[
    {
        "image_filename": "designv11_12_0000886_iros.2007.4399168-Figure4-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000886_iros.2007.4399168-Figure4-1.png",
        "caption": "Fig. 4. Two-body physical model of rover",
        "texts": [
            " \u2022 Rover moves by 4 wheel drive system (4WD) \u2022 4 wheel motors can generate driving force indepen- dently \u2022 All the wheel motors and wheels are the same \u2022 Movable property of the rover is restricted in x-y plane and not considering steer motion \u2022 Rover moves very slowly \u2022 Movement of payload is done by a linear actuator \u2022 Center of mass position of payload and body part is located in midle for z-direction \u2022 Terrain is perfectly symmetrical \u2022 Angle of terrain gradient is estimated by camera \u2022 Driving force is generated by static frictional force [ [ \u2022 Static friction coefficient between wheel and terrain is constant value \u2022 Wheels contact with terrain by single point (not consid- ering ditch like terrain) To solve the dynamic effect of a rover, dynamic equation should be introduced. Considering the assumption as mentioned above, the physical model of a rover is shown in Fig.4. Since all the situations are symmetrical, the model is shown as a side view. From this model, the dynamic equation of the robot can be easily solved as dynamic equation on the twobody model (payload and body). The dynamic equation of the two-body model can be solved as two dynamic equations. One is dynamic equation of the relative position between payload and body, another is one of body position. In the coordinate system shown in Fig.4, x-y coordinate system is the inertial coordinate system and X-Y coordinate system is the coordinate system based on body attitude of a rover. The physical parameters of the rover are as following. ~lbody[m]:Length of body, hbody[m]:Height of body lw[m]:Length of payload, hw[m]:Height of payload ~hr[m]:Height from center of wheel to bottom of body ~r[m]:Wheel radius, \u03c6[rad]:Pitch angle of body ~Mwheel[kg]:Weight of wheel ~k1:Proportion of payload weight against wheel k2:Proportion of body weight against wheel ~Xp, Yp[m]:X-coordinate and Y-coordinate of payload center of mass position ~xb, yb[m]:x-coordinate and y-coordinate of body center of mass position ~Xb, Yb[m]:X-coordinate and Y-coordinate of body center of mass position ~Xw[m]:Relative position between payload and body center of mass ~\u03b31[rad]:Angle of terrain gradient for front wheel ~\u03b32[rad]:Angle of terrain gradient for rear wheel ~u, v[N]:Force against payload from body in X-Y coordinate system ~F1, N1[N]:Driving force and normal force of front wheel ~F2, N2[N]:Driving force and normal force of rear wheel Dynamic equation of the relative position between payload and body in X-Y coordinate system is solved as following"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001937_iecon.2010.5675184-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001937_iecon.2010.5675184-Figure1-1.png",
        "caption": "Fig. 1. Schematic diagram of the four-rotor mini helicpter",
        "texts": [
            " Although the I&I approach has many superior features compared to other conventional adaptive control that relies on the so-called certainty equivalence, it is not straightforward to apply it to complex nonlinear systems such as a four-rotor mini helicopter. To overcome this problem, a relatively simple model parameterized with respect to a fewer unknown coefficients is developed for the adaptation law. Then an adaptive controller is designed. Numerical simulation illustrates the effectiveness of the adaptive controller for trajectory control in the presence of aerodynamic disturbance and parametric uncertainty. Fig.1, shows a schematic diagram of the four-rotor helicopter under consideration. The thrust ( )1, ,4jf j = is produced by the rotating rotor j . The main thrust f is the sum of the thrusts of each rotor. The rotors can be grouped into the front-back pair, and the left-right pair. Since the pairs rotate in opposite directions, one pair rotates clockwise and the other rotates anti-clockwise. Thus the yawing moment generated by the rotors cancel out to zero. A rolling moment can be generated by speeding up one of the rotors in the left- 978-1-4244-5226-2/10/$26"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001196_0094-114x(72)90004-3-Figure4-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001196_0094-114x(72)90004-3-Figure4-1.png",
        "caption": "Figure 4. A PCCC chain in isometric projection.",
        "texts": [
            " I f ~x, ---- 0, a~ must be positive and s, or s~+~ (the adjacent fixed translation must have a value of zero). Note that the last restriction states that any link of the R C C C chain which contains parallel axes must have both a positive link length and a revolute joint on one of its two axes for the linkage to be a single-degree-of-freedom mechanism. The knowledge of these unique mathematical models greatly simplifies the search for simpler linkages of the R C C C chain. I f the R C C C chain of Fig. l is replaced by the P CCC chain of Fig. 4, the only difference is in the constraint at joint 1. For the P C C C chain, the rotation 01 is fixed and the translation sl becomes the free motion at joint I. This removes restrictions 3 and 5 for the PCCC chain. With these relaxations, a complete and unique model set is produced for the PCCC chain. Note also that the equations which are generated to describe the RCCC chain are also valid for the PCCC chain with the small change in the treatment of 01 and sl. Thus the two chains will have similarities"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000766_j.automatica.2007.11.007-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000766_j.automatica.2007.11.007-Figure1-1.png",
        "caption": "Fig. 1. On local non-observability along a trajectory of a discrete-time dynamic system.",
        "texts": [
            " Henceforth, tensor notation is used, meaning that superscript indices refer to components of a contravariant tensor and subscript indices to components of a covariant tensor. The Einstein sum convention is arranged whenever the range of the used index is clear from the context. Furthermore, the abbreviation \u2202\u03b1x = \u2202 \u2202x\u03b1x k is applied. From the geometric interpretation of observability we deduce that system (2) is obviously not observable if the output functions c\u030c \u03b1y k = c\u030c\u03b1y (k, xk) are invariants of a non- trivial transformation group \u03a6k,\u03b5, c\u030c \u03b1y k = c\u030c \u03b1y k \u25e6 \u03a6k,\u03b5, see e.g., Olver (1993). In Fig. 1 the property of non-observability is demonstrated geometrically by means of an exemplary second-order discrete-time system. Here, despite two different trajectories associated with distinct initial conditions, the output yk retains unchanged. The two trajectories are related by a nontrivial transformation group (4), which has the output function as an invariant. As illustrated in Fig. 1 the points of the range of a symmetry group are included in the preimage of the output function for fixed k. In order to establish the existence of a transformation group such that the output functions are invariants, let us consider the corresponding infinitesimal criterion which results for the observability problem in vk ( c\u030c \u03b1y k ) = v \u03b2x k \u2202\u03b2x c\u030c \u03b1y k = \u2329 dc\u030c \u03b1y k , vk \u232a = 0 (5) with the one-forms dc\u030c \u03b1y k = \u2202\u03b1x c\u030c \u03b1y k dx\u03b1x \u2208 \u0393 (T \u2217(EX )), i.e. sections of the cotangent bundle (T \u2217(EX ), \u03c4 \u2217 EX , EX ), and the canonical product \u3008\u00b7, \u00b7\u3009 : T \u2217 (EX ) \u00d7 T (EX ) \u2192 C\u221e (EX )"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000698_j.ijmecsci.2008.02.003-Figure3-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000698_j.ijmecsci.2008.02.003-Figure3-1.png",
        "caption": "Fig. 3. The rub and impact forces.",
        "texts": [
            " Chen / International Journal of Mechanical Sciences 50 (2008) 1090\u201311131094 The resulting damping forces about the journal center in the radial and tangential directions are determined by integrating Eq. (4) over the area of the journal sleeve: f r \u00bc Z p 0 Z L 2 L 2 pR cos ydzdy (5) t \u00bc Z p 0 Z L 2 L 2 pR sin ydzdy (6) f e \u00bc f r; f j \u00bc f t Therefore, f e \u00bc mRL3 24c2Gz Z p 0 \u00bd \u00f0 o\u00fe 2 _j\u00de sin y\u00fe 2_ cos y cos y \u00f01\u00fe cos y\u00de3 dy (7) f j \u00bc mRL3 24c2Gz Z p 0 \u00bd \u00f0 o\u00fe 2 _j\u00de sin y\u00fe 2_ cos y sin y \u00f01\u00fe cos y\u00de3 dy (8) 2.2. Rub-impact force From Fig. 3, the radial impact force fn and the tangential rub force ft could be expressed as f n \u00bc \u00f0e d\u00dekc (9) f t \u00bc \u00f0f \u00fe bv\u00def n; if eXd (10) Then we could get the rub-impact forces in the horizontal and vertical directions: Rx \u00bc \u00f0e d\u00dekc e \u00bdX \u00f0f \u00fe bv\u00deY (11) C.-W. Chang-Jian, C.-K. Chen / International Journal of Mechanical Sciences 50 (2008) 1090\u20131113 1095 C.-W. Chang-Jian, C.-K. Chen / International Journal of Mechanical Sciences 50 (2008) 1090\u201311131096 Rx \u00bc \u00f0e d\u00dekc e \u00bd\u00f0f \u00fe bv\u00deX \u00fe Y (12) 2.3. Dynamic equations Fig",
            " Note that the time series data of the first 800 revolutions of the rotor are deliberately excluded from the dynamic behavior investigation to ensure that the data used correspond to the steady state. The following values for the non-dimensional parameters are used: x1 \u00bc 0:0525; x2 \u00bc 0:026; f \u00bc 0:2; Cp1 \u00bc 2:0; Com \u00bc 0:125 (29) The five methods are applied to analyze the dynamical system: the dynamic trajectories, power spectra, Poincare\u0301 maps, bifurcation diagrams and the maximum Lyapunov exponent. These feature properties are all manipulated together to determine the onset conditions for chaotic motion (Fig. 3). C.-W. Chang-Jian, C.-K. Chen / International Journal of Mechanical Sciences 50 (2008) 1090\u20131113 1107 C.-W. Chang-Jian, C.-K. Chen / International Journal of Mechanical Sciences 50 (2008) 1090\u201311131108 Figs. 4(a\u2013d) are the bifurcation diagrams of bearing center and rotor center in the horizontal and vertical directions with rotating speeds as control parameters. The simulation results show that the vibration amplitude of bearing and rotor center are great at low speed ratios and dynamic responses are synchronous with period-one"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000726_s11044-007-9072-4-Figure4-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000726_s11044-007-9072-4-Figure4-1.png",
        "caption": "Fig. 4 A hub-beams system in the initial state",
        "texts": [
            " The mass density of the hub is \u03c1c = 3 \u00d7 103 (kg/m3), and the length, width and height of the hub are 2a = 0.18 m, respectively. The mass and the inertia matrix of the hub are mc = 8\u03c1ca 3,J c = diag(J, J, J ), respectively, where J = 2mca 2/3. Runge\u2013Kutta integration method is employed for simulation in which the time step size is 10\u22125 (s) and the error tolerance is set to 10\u22128 (m). In this simulation, the gravitational force is not taken into account. 6.1 Effect of torsion Initially, the system is in static state without deformation. As shown in Fig. 4, the bodyfixed frame of the hub coincides with the inertial frame, and the body fixed frames of the two beams are parallel to the inertial frame, thus, rc = 0,\u0398c = 0,\u03982 = 0. The time history of the driving torque is given by Md(t) = [Md1(t) Md2(t) 0]T (Nm), where Md1(t) and Md2(t) takes the form Md1(t) = { 10 sin(pt) Nm t < 0.1 s, 0 t \u2265 0.1 s, (80) Md2(t) = { 10 sin(100\u03c0t) Nm t < 0.1 s, 0 t \u2265 0.1 s, (81) where p is a multiple of 10\u03c0 . Thus, F c = 0,Mc = Md(t). Each beam is divided into 6 elements"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003202_0022-2569(66)90030-9-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003202_0022-2569(66)90030-9-Figure2-1.png",
        "caption": "FIG. 2.",
        "texts": [
            " The dual angle between two lines in space is expressed by the quantity ~0 = q~o + og~P i (2) where ~Po is the angle between the lines and qh is the perpendicular distance between the lines. For the solution of this problem we use formulae defining the dual angle ~ between links 1-4 and 4--3, dual angle X between links 2-3 and 3-4, and dual angle 0 between links 1-2 and 2-3. These dual angles are dependent on the angle of rotation \u00a2p of the input link 1-2 relative to link 1-4, regarded as fixed. The designations are indicated in Fig. 2. Introducing the variables: t Translator's note: See also Brand, Vector and Tensor Analysis, Chapter II. Wiley. O = tan-~ = t a n - ~ + c o ~ ~ COS2(tpo/2 ) = ~ o + co~(1 + @o2), t n ~k . ~ o . ~kl l o g ~ ( l + ~ o Z ) ' = a ~ = t a n - ~ +co ~- c o s 2 ( ~ o / 2 ) = q % + X ~o X~ 1 + o Zl . I+Xo2) ' X = tan~ = t a n ~ + o ~ - X\u00ae \"cos2(Zo/2)- 0 0o 01 1 01 2 \u00ae = t a n ~ = tan-~ + a ,~ \u2022 cos2(-0o/2 ) = \u00aeo + o~ S ' ( I + \u00aeo), we write the necessary relat ionships developed earlier in reference [1]",
            " It is natural that for movable mechanisms on which one or another restrictive conditions are superimposed it is necessary first of all to satisfy the condition of movability for the corresponding spherical mechanism, or, in other words, to ensure the existence of a solution of the primary part of the equation at different values of the angles of rotation of the driving link. If, in addition, the dual part of the equation does not appear to be contradictory, then that condition will be sufficient and the given spatial mechanism will possess mobility. We examine a spherical mechanism with axes 1, 2, 3, 4, (Fig. 3) corresponding to the spatial mechanism under investigation (Fig. 2). We first clarify the necessary and sufficient conditions for the absence of rotation around axis 4 while preserving a single degree-of-freedom mechanism. In the general case the absence of rotation about axis 4 reduces to the 2 / / / rigid tetrahedronal pyramid 1, 2, 3, 4. However, if axis 3 is superimposed on axis 1 it is possible to have two opposite rotations around axes 1 and 3; moreover there will not be rotations around axes 2 and 4. Hence, we arrive at the conclusion that a necessary condition in the corresponding spherical mechanism (related to the investigated mechanism) is to constrain axes 1 and 3 to be parallel"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002634_s12541-012-0036-0-Figure4-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002634_s12541-012-0036-0-Figure4-1.png",
        "caption": "Fig. 4 Stair-stepping effect in 3-D microstructure",
        "texts": [
            " Sometime was required to settle the resin flow. The cross-sectional image was transferred to the DMD and the fabrication process was repeated until the last layer was complete. After completing the fabrication process, the redundant liquid resin on the microstructure was rinsed using a solvent such as alcohol and post-cured to enhance the solidity of the microstructure. In layered manufacturing, the stair-stepping effect necessarily occurs, and it deteriorates the surface quality of layered manufactured microstructure as shown in Fig. 4. This problem can be reduced by minimizing the layer thickness, but the thickness depends on the viscosity of the resin. This is because high viscosity resin needs considerable settling time for a flattened resin surface, and it is hard to control the thickness of the liquid layer. If the resin surface is not flattened and the layer thickness is not accurate, then shape error occurs in the completed microstructure. The Stairstepping effect is hard to reduce in a conventional P\u00b5SL process because one cross-sectional image per layer is usually used in the process"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001901_1.3591479-Figure5-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001901_1.3591479-Figure5-1.png",
        "caption": "Fig. 5 Skew four-bar mechanism with turning pairs on f ixed link",
        "texts": [
            " Spherical cardanic motion may also be envisaged as the motion of a link, two points of which are constrained by spherical guides, which are great circular arcs (equator and meridian). This corresponds to the case r = R \u2014 k = 90 deg. The coupler curve is a good deal simpler than in the general case, but remains of order 8 and genus 1. Skew Four-Bar Mechanisms (2 Turning Joints, 2 Spherical Jo ints ) Case (i) : Two Turning Joints F ixed. T h e re lat ion b e t w e e n the crank rotations is well known [4, 15, 16], The mechanism is shown in Fig. 5. In the figure, the offset between the crank axes is a (linear) and cx (angular). The offset of the plane of the crank circle, radius b, from the common perpendicular, is p and that of the plane of the C-crank circle, radius e, is q. The coupler, whose ends are balljointed, is of length c -f- d and P is a point on the coupler at a distance c from B and d from C. The crank rotations, 8, <j>, are measured from a direction at right angles to the common perpendicular AD. The origin, O, is halfway between A and D; the z-axis is along DA and the x-axis is inclined 0 A ) a to the crank axes, xyz being right handed"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003026_j.phpro.2012.10.037-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003026_j.phpro.2012.10.037-Figure1-1.png",
        "caption": "Fig. 1. A schematic drawing of the sample preparation by laser treatment",
        "texts": [
            "5%Fe), because of its low ductility at room temperature and high springback. This precipitation hardened aluminium alloy is extensively used in the aeronautical industry for engine parts, fuselage and wing skins [2]. The used sheet material is solution heat treated in the range of 488- cold rolled to 0.4 mm and naturally aged. To compare the effect of locally softened zones on the springback, different rectangular sheets with a length of 225mm and a width of 70mm were locally heat treated along a line in the centre of the sheet (Fig. 1) and bent to an angle of 81\u00b0 using a servo controlled hydraulic LVD-PPEB50/20 press brake. Localized softening of the material was performed by means of an Nd-YAG (1.06 \u03bcm) laser with a maximum power of 500W and a M2 value of 88.8.The intensity distribution of the laser beam in the laser spot is even (top hat distribution) and the laser energy is delivered using a flexible fiber-optic delivery system. The sheets were clamped tightly at the two opposing edges in the width direction (Fig. 1) and coated with Grapghite 33 for good absorption of the Laser energy. After which, a single or multi-path laser scanning strategy was used to heat treat the material along a line in the centre of the sheet. For preventing localized melting of the material close to the edges due to localized heat accumulation, the graphite coating was removed over a width of 5mm near the edges and laser heat input reduced accordingly (Fig. 1). In order to ensure good control of the temperature distribution during heat treatment, the surface temperature was recorded by means of an IR camera (Thermovision A20) with spectral range of 7.5 to 13 -radiated side of the test sample, at a distance of 40 mm from the sheet surface. In-process angle measurement during bending was performed by means of an LVD Easy-form laser assisted angle monitoring system in which the angle of the two bending flanges is measured in real time at multiple points"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003713_978-1-4614-2419-2-Figure28.3-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003713_978-1-4614-2419-2-Figure28.3-1.png",
        "caption": "Fig. 28.3 Scheme of a tracking-interferometer coordinate system",
        "texts": [
            " There are two right triangles with the corner angles a0 and a00 sharing their hypotenuse which leads directly to a \u00bc a0. As the same relationships can be shown for b and b0, the beam length is (A + a) + (B + b) \u00bc 2\u2219(A + a) and is hence proven not to be biased neither by reflector rotation nor by lateral shift. Tracking-interferometers (TI) were developed in the late 1980s by the American National Bureau of Standards (NBS) as a measuring system that can track an optical reflector and detect its 3D position in a spherical coordinate system [6] as is shown in Fig. 28.3. The distance to the reflector is measured with an incremental laser interferometer or an absolute distance measurement system (ADM). The angles (y, j) can be measured by using the rotary encoders of the two-axis tracking unit. In tracking mode, a lateral movement of the reflector results in a lateral beam shift that is detected by an integrated position sensitive device (PSD) and serves as input for the tracking unit. Although originally developed for robots, TIs currently serve also as portable coordinate measuring machines and for the precise calibration of coordinate machines and machine tools"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003713_978-1-4614-2419-2-Figure47.1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003713_978-1-4614-2419-2-Figure47.1-1.png",
        "caption": "Fig. 47.1 Dynamic wheel test rig",
        "texts": [
            " The specific objectives are (1) to develop a test rig to measure the hub transmitted force and static radial stiffness of road bike wheels; (2) to experimentally measure and analyze these quantities using several wheels in order to disclose intra and inter variability; and (3) to investigate if static radial stiffness could be correlated to the transmitted force at the hub. J. Le\u0301pine (*) \u2022 Y. Champoux \u2022 J.-M. Drouet Ve\u0301lUS, Department of Mechanical Engineering, Universite\u0301 de Sherbrooke, 2500 boul. de l\u2019Universite\u0301, Sherbrooke, QC J1K 2R1, Canada e-mail: julien.lepine@usherbrooke.ca R. Allemang et al. (eds.), Topics in Modal Analysis II, Volume 6, Conference Proceedings of the Society for Experimental Mechanics Series 31, DOI 10.1007/978-1-4614-2419-2_47, # The Society for Experimental Mechanics, Inc. 2012 465 Figure 47.1 provides a schematic diagram of the test rig developed for studying wheel dynamic characteristics. The dynamic wheel test rig measures the blocked force at both ends of wheel axle. These ends are respectively clamped to a force transducer (Sensortronics 60001 S-Beam, capacity: 1,000 lb) to measure transmitted vertical force. Both transducers are clamped to a very stiff rigid structure connected to the ground in order to measure the hub axle vertical blocked forces. An eight channels LMS SCADAS Mobile data acquisition system was used and the data was analysed using LMS Virtual",
            " The tire pressure is adjusted to 8 bars. For the purposes of this paper, only the front wheels were tested. To simulate the typical preload of a 75 kg cyclist, the actuator is positioned to generate a nominal preload force of 250 N on the wheel In all the tests we conducted, the preload was adjusted to compensate any wheel out of round variation. A white noise signal was provided to the actuator with an average displacement range of 0.25 mm. A typical displacement power spectrum density (PSD) measured by the LVDT is shown in Fig. 47.1: Dynamic wheel test rig Figure 47.2. The DC component of the curve is due to the preload force. The amplitude response of the actuator drops drastically above 100 Hz. This is caused by the actuator speed limitation. The bandwidth of interest is 0\u2013100 Hz, with the knowledge that human sensitivity to vibration diminishes drastically above 100 Hz [3]. 466 J. Le\u0301pine et al. Figure 47.3 illustrates the typical PSD of the total blocked force transmitted by the hub. All the PSDs in this paper are calculated using 60 averages of 1 s segments overlapped at 67%"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002262_ijtc2011-61234-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002262_ijtc2011-61234-Figure1-1.png",
        "caption": "Figure 1 Sketch of the test apparatus: (a) overall view, (b) measurements apparatus with laser devices, (c) exciting apparatus with shaker.",
        "texts": [
            " The authors think that the floating body approach is the most correct way to bring into contact flat surfaces and to study the plane contact parameters and related wear even if the majority of test rigs adopts the rigid body approach. A test rig to perform flat-on-flat contact has been designed and set up at the Department of Mechanical Engineering of the Politecnico di Torino within the laboratory LAQ AERMEC. The test rig can be split into three functional blocks: the contact mechanism, the excitation system and the measurement and control system. Fig. 1 shows the overall view of the rig. Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 01/13/2016 Terms of Use: http://www.asme.org/about-asme/terms-of-use 2 Copyright \u00a9 2011 by ASME Contact mechanism The two specimens, machined with plane contact surfaces, rub against each other with a small amplitude oscillatory motion, see Fig. 2. The specimens, namely S1 and S2, are tightly connected to the Mobile Support (MS) and to the Fixed Support (FS), respectively, as shown in Fig. 3. The Mobile Support is integral with the frame that is excited with an electromagnetic shaker, Fig. 1c, while the Fixed Support is simply supported by two rods, the rod R and the V-shaped rod shown in Fig. 3. The other end of both rods is simply supported by force transducers FT1 and FT2. A sphere with a small radius is machined at the ends of the rods, detail in Fig. 3, so that the constraints behave as spherical hinges, hinges from H1 to H5. These hinges are unilateral kinematical constraints then a preload, forces P1 and P2, is needed to avoid the rods losing contact with the force transducers and the fixed support",
            " The friction force F and reaction forces acting on the rod R and Vshaped rod lie in the same plane, namely xy-plane, so that they do not add any moments on the contact surface. The contact mechanism can work at high temperature, up to 1000 \u00b0C, by means of induction heating. Two thermocouples TH are introduced into the specimens, as close as possible to the contact point, to measure the contact temperature. Excitation system The mobile support and relative specimen S1 are integral with a square frame supported by four springs, Fig. 1c. An electromagnetic shaker excites the square frame in the plane of symmetry at frequencies equal to or below the first resonance. Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 01/13/2016 Terms of Use: http://www.asme.org/about-asme/terms-of-use 3 Copyright \u00a9 2011 by ASME Measurement and control system Two laser doppler vibrometers, one for each contact surface, measure the velocity of each specimen. The velocity is integrated and the relative displacement calculated by difference"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001265_tmag.2009.2012576-Figure9-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001265_tmag.2009.2012576-Figure9-1.png",
        "caption": "Fig. 9. Analysis model of eddy current loss.",
        "texts": [
            " The length is unknown. Then, let us define the following contact resistance coefficient : (6) The relationship between and the stress is shown in Fig. 8. The figure denotes that the contact resistance coefficient is decreased with the stress. When the stress is small, is dispersed considerably, and the average value of becomes large. The eddy current analysis considering the resistance of the contact part is carried out. The effects of the stress, the exciting frequency and the dimension of magnet are examined. Fig. 9 shows the examined model. A large magnet 50 mm 50 mm 10 mm is segmented into six parts. The figure shows the 1/8 part. The magnetic field is impressed in the -direction. The frequency of the impressed field is 1 kHz. The magnetic flux density when there is no magnet is 13.5 mT, and the current was kept constant. is assumed as 0.01, 0.1, and 1.0 mm. Fig. 10 shows the eddy current distributions in the magnets at 6 MPa and 90 MPa. It can be understood that each segmented magnet can be treated as almost insulated magnet when the compressive force is not large (6 MPa)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002531_s10800-010-0153-3-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002531_s10800-010-0153-3-Figure2-1.png",
        "caption": "Fig. 2 Cyclic voltammograms for PPy electrodeposition on a graphite pellet in a 0.2 M pyrrole ? 0.1 M NaClO4 acetonitrile solution\u2014 15 cycles\u2014scan rate: v = 100 mV s-1",
        "texts": [
            " The scanning electron microscopy (SEM) photos of the surface of composite material pellets were realized on a SEMFEG Zeiss Ultra 55 instrument. Impedance measurements were performed by using either a Radiometer analytical Voltalab PGZ 301 or a Solartron 1287/1250 electrochemical work station, in the typical frequency range 10 mHz\u2013100 kHz. The amplitude of the AC signal was 10 mV. 3.1 In situ electrodeposition of PPy For the PPy electrodeposition, cyclic voltammograms (CVs) were recorded between -1.5 and 1.5 V/SCE in a 0.2 M pyrrole ? 0.1 M NaClO4 acetonitrile solution (Fig. 2). Wide anodic and cathodic waves occurred respectively at 0.75 and -1.25 V/SCE. The peak intensity values increased with the number of cycles, which indicated the progressive PPy electrodeposition on the graphite pellets. The anodic current density values were found to increase from 1.5 to 2.5 mA cm-2, and the cathodic current density ones from -1.7 to -2.8 mA cm-2. Moreover, it appeared a dark coloration on the graphite pellets, confirming the formation of a polymer deposit at the electrode surface"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000597_978-3-540-73129-0-Figure6.150-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000597_978-3-540-73129-0-Figure6.150-1.png",
        "caption": "Fig. 6.150. Determination of amplifier parameters via the signal form. a Monotone signal, b complex periodic signal",
        "texts": [
            " Secondly, one has to acquire the values of the voltage, current and power required for the operation. Nominal Voltage of the Amplifier. The required output voltage of the power amplifier is determined by the parameters of the actuator. When the operating range is not entirely used, an amplifier with a smaller output voltage will suffice. Average Current, Continuous Output Power. The average of a sinusoidal current signal can be computed through the equation I = fCVpp, and the continuous output power through P = fCVppVD, VD being the nominal voltage of the amplifier, see Fig. 6.150a to the right. Complex signals, however, have periods that can contain several charges with varying voltage amplitudes. When calculating the average current which is required to drive a capacitive load, it is not the rate of rise or decay of charging and discharging that matters, but the sum of the individual charging processes per signal period, the actuator capacitance and the repetition frequency. A piezo actuator that has a large signal capacitance of 10\u00b5F is first to be charged to 120V, then discharged to 80V and finally recharged to 180V. Discharging takes place in the reverse order. The overall cycle is to be repeated at a frequency of 150Hz. The sum of the charging voltages is then 120V plus 100V during the charging process, and 40V during the discharging process, adding up to a total of 260V. The average current is computed via the product of 150Hz \u00b7 10\u00b5F \u00b7 260V, resulting in 390mA. Figure 6.150b to the right displays an example of such a composed signal. The average power is calculated using the 200V nominal voltage of the amplifier: 0.39A \u00b7 200V, which equals 78W. When dealing with ER actuators, one has to consider the additional energy required because of the electrical conductance. Maximal Current, Pulse Power. Computing the maximal current demands knowledge of the greatest slope in the voltage-time signal. This is gained by means of a curve tangent or by differentiating mathematically. The maximum current results from the charge equation dq = CdV , which, after rearranging, reads Imax = C(dV/dt)max for the numerical and Imax = C(\u0394V/\u0394t)max for the graphical solution (C is the so-called large signal capacitance). Figure 6.150 to the left shows two examples. The pulse power is determined via the product of the maximum current and the nominal voltage of the amplifier. When the ratio of the maximum current to the continuous current is high, one can usually neglect the current component related to the conductance of the ER fluid. A particularity of ohmic-inductive loads like magnetostrictive or magnetorheological actuators is that they \u2013 especially during dynamic operation \u2013 mainly require reactive power and only a little active power",
            " The parameters of the actuator to be driven form the basis for the choice of the appropriate power amplifier. The amplifier must be capable of providing the required current. The necessary operating voltage is determined by means of the greatest incline in the current-time signal that the amplifier has to produce and the load inductance. To this end, one applies a tangent to the geometric curve or differentiates it mathematically. The procedure is similar to the one described for power amplifiers used to drive capacitive loads: in the examples in Fig. 6.150 to the left simply replace V by I and vice versa and C by L. If power amplifiers are used that do not make use of variable or switchable operating voltages and which are not tuned to specific loads and signal forms, the continuous power of the analogue amplifier is determined by the greatest incline of the current signal and its corresponding voltage: P = VDImax. The need for a power amplifier for driving an unconventional actuator in general laboratory applications will in most cases result in the choice of an analogue amplifier without regard to the fact if there are solid state actuators or actuators with electrically controllable fluids: analogue amplifiers have the highest signal quality, the largest range of frequency and they allow for a wide value range of load impedance"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002043_pime_proc_1967_182_025_02-Figure9-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002043_pime_proc_1967_182_025_02-Figure9-1.png",
        "caption": "Fig. 9",
        "texts": [
            "comDownloaded from INFLUENCE 0 F COUPLED ASYMMETRIC BEARINGS ON THE MOTION OF A MASSIVE FLEXIBLE ROTOR 269 .................................. I I .................................. I Bay, 0 0 0 0 0 B,,, 0 0 0 0 0 0 Bb,, 0 0 0 0 0 B,,, 0 0 0 0 .................................. + - E a x x E,XY El 1 Ebvv E22 ........................ Eaw E U Y Y E b Y X These equations appear formidable but apart from being the eigenvector equations of the real rotor they are also the precise equations of motion of the comparatively simple system shown in Fig. 9 in which a number of masses are spring mounted with certain eccentricities at points on a rigid and massless shaft carried in the same bearings as the initial system. There is a one to one correspondence between these masses and the various modes of vibration of the real system which they fully represent. Subjecting each of the original rotor modes to the normalizing condition 2 mq5: = 1, the relationships between the coefficients, the substitute system of Fig. 9, and the real system are: A,, = A,, = M,(L-T~) /L M,. = (1 m+Y)2 c m+rz 2, = -Arb = Abr = M,.ZT,/L A,, = M , z: m9r 1 Ebxy i ............ I I E22 1 Ebyu El 1 It can be shown that and consequently Aaa+Aab+Aba+Abb = M = total mass It will be noted that the forcing vector postulated in the new general equations is rather different in form from that suggested by the author. For a real and imperfect rotor, in which the masses lie at variable distances w from the Proc Instn Mech Engrs 1967-68 Vol182 Pt 1 No 13 at NANYANG TECH UNIV LIBRARY on June 9, 2016pme",
            "comDownloaded from 270 COMMUNICATIONS straight line joining the centres of the journals, ex, and e,, are components of er = (2 m ~ h ) / ( 2 mdr) Having established this interpretation of the system the arguments of an earlier paper (16) become applicable wherein it was shown that the amplitudes of motion of the single mode system could be established by a simple graphical process. Apart from the non-dimensional bearing characteristics, and the rigid bearing critical speed of the rotor, it was shown in (16)~ and confirmed in (17) that the whole quality of the rotor-bearing interaction was much influenced by the ratio radial clearance x rotor stiffness/bearing force It would be useful in the present paper to have an indication of the values of these parameters considered. A final point of great interest may be deduced from the equivalent physical system of Fig. 9. Consider a rotor running in bearings of any dissipative linear characteristics and at one of its critical speeds, i.e. SZ = A,. The orbits of the mass centre G,, and the geometric centre C, of the mass M , and also the corresponding point of attachment P, to the shaft must all be finite ellipses. Taking an origin 0 at the position of static equilibrium of P, and sketching a diagram indicating inertia forces, spring deflection, etc. it is easily shown that z, = local \u2018shaft\u2019 displacement = G,C, = E7",
            " The reason for these restrictions (perfectly reasonable ones for the majority of rotors) is that a graphical solution for the fully coupled case is more difficult to apply. Dr Morrison\u2019s equations-substantially the same as those appearing in (5)-can be solved quite easily numerically; the present paper attempts however to give a qualitative interpretation of the solutions of such equations. Incidentally it appears that Dr Morrison\u2019s forcing vector is deficient in four terms associated with the bearing coordinates. These terms, >instead of zero, should read 2 Aa,ex,,, 2 AbrgXr, Aa,eyr, 2 AbrZY,. Physical visualiza- tions of the type shown in Fig. 9 condense the normal coordinates and give the illusion of simplicity. Unfortunately the governing equations still remain the same, and there is no way out of performing the appropriate arithmetical operations. One cannot for instance use the principle of superposition to treat one mode a t a time, although superficially this might appear possible. I fear that I must disagree with Dr Morrison\u2019s penultimate paragraph. To fix ideas, consider Fig. 14, which is a simplification of the case under discussion"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001882_j.commatsci.2009.02.005-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001882_j.commatsci.2009.02.005-Figure1-1.png",
        "caption": "Fig. 1. The schematic of laser light strip.",
        "texts": [
            " In this study, a mathematical model is developed to conduct numerical simulation of the temperature field and weld pool geometry for HPDDL welding with the strip-shaped beam output, and HPDDL welding experiments on thin sheets of mild steel are carried out to verify the analysis results. It lays good foundation to get insight of the new welding process mechanism and to get basic data for process development, defect explanation and process control. As aforementioned, unlike conventional laser welding or arc welding processes which generate round heating spots at the workpiece surface, HPDDL generates the inherent rectangular output beam profile. The strip-shaped beam, as shown in Fig. 1, is 12.5 mm in length and 0.5 mm in width. However, the heat intensity is not uniformly distributed within the rectangular strip. The users\u2019 manual of the product HPDDL ISL1000 gives the measured data of heat intensity distribution within the light strip [11]. To make the data be useful for general cases, they are normalized to the range of 0.0 and 1.0. The normalized data of heat intensity distribution along the length direction (ys) and the width direction (xs) are demonstrated in Figs. 2 and 3, respectively"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000707_s11044-006-9032-4-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000707_s11044-006-9032-4-Figure1-1.png",
        "caption": "Fig. 1 The 3-SPS/PU parallel manipulator and its coordinate system",
        "texts": [
            " In this paper, a novel analytic approach and CAD variation geometry approach are pro- posed for analyzing kinematics and dynamics of a 3-dof parallel manipulator with three SPS-type active legs and one PU-type passive leg by using virtual serial mechanism and virtual work theory. 2.1 The 3-SPS/PU parallel manipulator A 3-SPS/PU parallel manipulator includes a platform m, a base B, and three SPS-type active legs ri (i = 1, 2, 3) with the linear actuators, and one PU-type constrained passive leg ro, see Figure 1a. Here, m is an equilateral ternary link a1a2a3 with the sides li (i = 1, 2, 3) and a center point o. B is an equilateral ternary link A1A2 A3 with the sides Li and a center point O . Each of the SPS-type active legs ri connects m to B by a spherical joint S at ai , a active leg ri with a prismatic joint P, and S at Ai (i = 1, 2, 3), respectively. The constraint passive leg ro is perpendicular to m, and connects m with B by P on m at o and U on B at O . Let {m} be a coordinate system o-xyz fixed on m at o; let {B} be a coordinate system O-XY Z fixed on B at O ",
            " In the 3-SPS/PU parallel manipulator, the number of links is k = 9 for 1 platform, 3 cylinders, 4 piston-rods, and 1 base; the number of joints is g = 11 for 4 prismatic joints P , 1 universal joint U , and 6 spherical joints S; f1 = 1 for the prismatic joint, f2 = 2 for the universal joint, f2 = 3 for the spherical joint; the redundancy dof (degree of freedom) is M0 = 3 for 3 SPS-type active legs rotating about their own axes, and M0 has no influence on the kinematic characteristics. Therefore, based on the revised Kutzbach\u2013Grubler equation [1, 2], the dof of the 3-SPS/PS parallel manipulator is calculated as M = 6(k \u2212 g \u2212 1) + g\u2211 i=1 fi \u2212 M0 = 6 \u00d7 (9 \u2212 11 \u2212 1) +(4 \u00d7 1 + 1 \u00d7 2 + 6 \u00d7 3) \u2212 3 = 3. (1) 2.2 The simulation mechanism A simulation mechanism of the 3-SPS/PU parallel manipulator is created by using the CAD variation geometric approach [12] (Figure 1b). Its construction processes are described as follows. 1. Construct the base B in 2D sketch. The subprocedures are: a. Construct an equilateral triangle A1 A2A3 by using the polygon command. b. Coincide its center point O with origin of default coordinate, set its one side horizontally, and give its one side a fixed dimension Li = 120 cm. c. Transform A1 A2A3 into an equilateral triangle plane by using the planar command. 2. Construct the platform m in 3D sketch. The subprocedures are: a. Construct three lines li (i = 1, 2, 3), and connect them to form a closed triangle a1a2a3"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002120_rspa.2009.0118-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002120_rspa.2009.0118-Figure1-1.png",
        "caption": "Figure 1. Local contact model: (a) Derjaguin approximation and (b) current model.",
        "texts": [
            " However, non-local models are computationally demanding, so that, on the side of economy, a local model is preferred, provided that it can describe the relevant physical behaviour. In defining the contact law, we assume that the traction acting at a point on the surface S1 is a given function of the distance to the surface S2. To fully define the contact traction at a point, one must define: (i) the distance on which the traction depends, (ii) the direction of the traction, and (iii) the area to which the traction is referred. (i) Physical reasoning requires that the distance a in figure 1b, governing the traction p that surface S2 exerts on a point of S1, be the shortest distance from the point in question to the smooth surface S2, i.e. along the normal to S2, and so not, in general, normal to S1. (ii) To determine the rational direction of the traction p(a), consider the atomic forces shown in figure 1b. A mildly curved surface can be locally treated as a plane. The forces acting on a point on S1, from equidistant neighbours of the central atom on S2, are symmetrical and have the resultant directed along a. Thus, unless the curvature of S2 is very high, the direction of the traction on the point on S1 is reasonably assumed to be along the normal. Proc. R. Soc. A (2009) (iii) The reference area for the traction p(a) presents a more challenging problem. Two simplest (but not the only) options are the reference (undeformed) area and the current (deformed) area",
            " A (2009) The successive approximations to a non-local quasi-continuum model (Sauer & Li 2007a), developed for the purpose of numerical implementation, indicate that the first two assumptions are exact for flat, parallel surfaces, and are first-order approximations otherwise. By contrast, the Derjaguin (1934) approximation, commonly used in the normal contact between spheres (e.g. Greenwood 1997), implies that (i) the relevant distance is along the line parallel to the line that connects the centres of the spheres, as shown in figure 1a, (ii) consistently, the direction of the traction p(a 0) is along a 0 in figure 1a, and (iii) the resulting traction is given per unit area of the projection on the mid-plane. The local contact model used in the FE analysis is a surface-to-point contact law, i.e. the traction at a point on S1 is the result of action of all (in principle) atoms of solid 2, while the traction at a point on S2 is the result of action of all atoms of solid 1. Newton\u2019s third law must be satisfied by the resultants of the contact forces between the solids. In the present case, this is guaranteed by the equilibrium conditions implicit in the FE formulation"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000880_tmag.2007.891399-Figure4-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000880_tmag.2007.891399-Figure4-1.png",
        "caption": "Fig. 4. Three-dimensional mesh of EMI drive mechanism.",
        "texts": [
            " The ring magnet has 8 poles, and is inserted in the gap between the inner and outer conductors. The air gaps of both sides between the conductors and the ring magnet are 0.2 mm. The interlinkage flux of conductors is changed when the ring magnet rotates at the constant speed. The transient torque is produced by the eddy current flowing in the conductor when the ring magnet rotates. In order to obtain the big EMI torque, this mechanism has the divided conductors to use the resonance phenomena. Fig. 4 shows the 3-D mesh of EMI drive mechanism. The mesh of ring magnet is equally divided in steps of 1.5 for the automatic mesh rotation. In this calculation, it is supposed that the ring magnet is not inserted into the gap between the conductors at the initial condition in this impact drive mechanism. Therfore, the initial potentials of all edge elements are set to be zero in order to simulate under the conditions of that the ring magnet is rotated at constant speed and suddenly inserted into the gap between the conductors, and the ring magnet is rotated from the rotation angle of 45 "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002949_etep.1642-Figure6-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002949_etep.1642-Figure6-1.png",
        "caption": "Figure 6. The flux density distribution.",
        "texts": [],
        "surrounding_texts": [
            "4.1. Magnetic field distributions and typical air-gap flux density waveform at standstill Under the locked-rotor conditions, the frequency of rotor e.m.f. equals the supply frequency so that rotor-related skin effects and the saturation of the upper rotor tooth can be more severe than the rated load. Furthermore, the rotor magnetic field distribution is no longer symmetrical upon a broken bar fault, as illustrated in Figures 5 and 6. In addition to changes in the broken bar regions, the field and flux density distributions at other positions in the stator and rotor core are also distorted and increased to some extent. It can be seen that the presence of a broken bar in Figure 5(b) or two broken bars in Figure 5(c) results in highly saturated regions in the neighboring core iron that may cause a progression of the fault. If observed around the air gap, the radial and tangential components of flux density waveforms at one time instant can be obtained, as presented in Figure 7. In order to study the variation of the air-gap flux density at standstill, the frequency spectrum of the air-gap flux density was analyzed by the harmonic frequency analysis. Figure 8 shows that the air-gap flux density contains high spatial harmonics components and varies significantly with broken bar faults. It is also noted that broken bar faults affect the tangential and radial components differently. The fundamental component and each harmonic of the tangential component change little with the faults, but the impact on the radial components of each harmonic is more evident. For instance, the harmonic component increases in magnitude obviously in addition to the seventh and ninth. It is different from rated load wherein the tangential component of air-gap flux density is large at standstill. Taking the case of prototype in the paper, under healthy conditions, the tangential component and radial component ratio is 63.64% at standstill, but the ratio is 4.71% under rated load condition, so the tangential component should be taken into account in the air-gap flux density calculation at standstill. Measured and computed results of the fundamental component of the air-gap flux density are given in Table II. Measurements were obtained from the search coils installed on the stator, whereas calculated results were based on the FE method. From this table, two sets of results agree with each other reasonably well with a disparity of less than 5%. Besides, it can be found that the flux density in the case of broken bar faults is higher than that in the normal cage. Copyright \u00a9 2012 John Wiley & Sons, Ltd. Euro. Trans. Electr. Power (2012) DOI: 10.1002/etep Copyright \u00a9 2012 John Wiley & Sons, Ltd. Euro. Trans. Electr. Power (2012) DOI: 10.1002/etep 4.2. Stator current profiles The stator current waveforms at locked-rotor conditions were recorded in the test and in the simulation, as shown in Figures 9 and 10 for comparison. In theory, the rotor magnetomotive force (m.m.f.) asymmetry caused by broken rotor bar faults can produce a backward-rotating field and then introduce Copyright \u00a9 2012 John Wiley & Sons, Ltd. Euro. Trans. Electr. Power (2012) DOI: 10.1002/etep lower sideband components in the stator current spectrum at double slip frequency. When the machine is loaded with rated torque, the amplitude of stator currents periodically fluctuates with time. Nonetheless, when the tests are performed at standstill, one can notice that the periodical pulsation of stator current cannot be found in the motor with broken bars. It is because of the tests being performed at standstill (s = 1). The fault-specific sideband components of the machine\u2019s current do not appear near the fundamental. In the experiment, the current amplitude ranges from 22.2 (healthy rotor) to 21.2 (one broken bar) to 20.5 A (two broken bars), although in the simulation, these differences between the three cases are almost negligible. 4.3. Rotor current profiles It is clear that the harmonic content of the m.m.f. depends on the distribution of the stator conductors in the stator slots and the currents flowing in them. If fed with a balanced sinusoidal three-phase supply, Table II. The fundamental component comparison of air-gap flux density between experimental value and calculation value at standstill. Rotor type Calculated (T) Measured (T) Error (%) Healthy rotor 0.391 0.380 2.895 One broken bar 0.408 0.422 3.318 Two broken bars 0.448 0.471 4.883 (a) With healthy rotor (b) With one broken bar fault (c) With two broken bar faults Figure 9. The experimental stator current waveforms. Copyright \u00a9 2012 John Wiley & Sons, Ltd. Euro. Trans. Electr. Power (2012) DOI: 10.1002/etep the stator winding would in effect create a rotating constant flux that sweeps the rotor bars to generate the rotor e.m.f.. As a result, the e.m.f. and the induced current are identical in all the rotor bars around the rotor periphery. At standstill, however, the rotor bars are fixed in position with respect to the stator and are thus subjected to local variations in the stator m.m.f. [35]. The current amplitude of the healthy rotor varies with position around the rotor periphery and is not equal at standstill, which is different from the rated load. Figure 11 indicates the broken bar positions within the rotor geometry that is used throughout the paper, and Figure 12 shows the rotor current distributions in the three cases. It is easily understood that the currents would increase in the neighboring bars next to the broken ones. When bar 3 is broken, the currents of bars 2 and 4 are seen to rise, and in the case (b), the amplitude of the current in bar 4 presents a staggering increase. Similarly, bars 2 and 5 increase dramatically when bars 3 and 4 are broken in the case (c). Moreover, the rotor current distribution becomes more asymmetrical arising from magnetic distortions. These can be explained by the fact that the amplification of some rotor currents is directly related to the number of broken bars when the motor is operating either full load or upon a fault [19]. In fact, the increase in rotor current amplitudes at standstill is more significant than the rated load so that the fault is quicker to propagate and poses more severe a threat. Copyright \u00a9 2012 John Wiley & Sons, Ltd. Euro. Trans. Electr. Power (2012) DOI: 10.1002/etep Copyright \u00a9 2012 John Wiley & Sons, Ltd. Euro. Trans. Electr. Power (2012) DOI: 10.1002/etep 4.4. Torque profiles and magnetic force of rotor bar Any changes in flux distributions or rotor currents will bring about a change in torque and magnetic force of rotor bar developed by the machine. Figure 13 demonstrates the torque of the motor with the three rotor bars at locked-rotor conditions, whose torque distribution waveform shapes are almost identical. Through further investigation, it can be observed that the average torque is reduced at locked-rotor conditions from healthy, one broken bar to two broken bars (from 12.28 to 11.23 to 10.22Nm, respectively). It becomes obvious that the average torque continues to decrease with the number of the broken bars. This is the reason why the loading capability of the motor with broken bars is getting worse, and sometimes it cannot start. The magnetic force distributions of rotor bar at standstill conditions are shown in Figure 14. It can be seen that the bars with the highest magnetic force are those immediately adjacent to the broken bars, including the tangential and normal components of the force. Without a doubt, this excessive force gives rise to thermal and mechanical stresses, vibration and wear in the neighboring bars, which individually or in combination may cause fault propagation or even a breakdown of the motor if left untreated. Copyright \u00a9 2012 John Wiley & Sons, Ltd. Euro. Trans. Electr. Power (2012) DOI: 10.1002/etep"
        ]
    },
    {
        "image_filename": "designv11_12_0002440_10426914.2010.544806-Figure3-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002440_10426914.2010.544806-Figure3-1.png",
        "caption": "Figure 3.\u2014The concept of the titanium built-up welded beam.",
        "texts": [
            " A sizeable amount of material that is removed from a thick plate or billet is often wasted, and this directly contributes to an increase in the cost of the end product. Further, the thicker section bars are far more expensive and difficult to make, and are often not readily available with material suppliers. D ow nl oa de d by [ C hu la lo ng ko rn U ni ve rs ity ] at 0 9: 02 0 2 Ja nu ar y 20 15 The concept of built-up beams from thin plate elements is new to products made from titanium. This concept was tried for the first time in this research project. A schematic view of the built-up concept used in this study is shown in Fig. 3. Two flange plates (horizontal plates) are welded on to the two edges of a web plate (vertical plate) to form an I-shaped cross-section. The loading points and locations at the two end supports are provided with stiffeners to prevent and/or minimize any local buckling of the web under the influence of a concentrated load. A typical cross-section of the welded built-up beam is shown in Fig. 4. This figure also shows the stiffeners that were welded on to the web to prevent local web buckling under the direct influence of a concentrated load"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001003_j.engfailanal.2006.11.052-Figure3-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001003_j.engfailanal.2006.11.052-Figure3-1.png",
        "caption": "Fig. 3. Identification of the strain gauge on the gearbox outside side.",
        "texts": [],
        "surrounding_texts": [
            "In this work nine rosettes were bonded (45 rosettes \u2013 3 linear strain gauges) in the following positions: rosette 1 in the central zone of the cover of the housing to measure the stresses in the cover; rosette 2 in a zone away from weld areas. In the box body, to measure the nominal stresses; rosette 3 in the weld toe of weld zone on the lower reinforcement box, place where fatigue cracks were started; rosettes 4\u20139 were bonded in the six supports of the housing to obtained the reaction forces (Figs. 2 and 3). The same procedure of surface preparation has been conducted by the authors in the study of cast steel railway coupling used for coal transportation [7]. The data was collected with a portable PC and the HPVEE processing system was used for treatment and analysis of the signals. The data was obtained in service, in the Lisbon\u2013Porto Intercity passengers train, with a maximum speed of 160 km/h and in the Entroncamento\u2013Guarda freight train, with a maximum speed of 120 km/h (Table 3). The program of the rosettes readings were made in order to get the best possible comparative information. Those stages are the critical section of the journey, in what concerns speed, power and track oscillations. For the Lisbon\u2013Porto journey in the Intercity, the following 6 stages were selected: Entroncamento\u2013Fa\u0301tima: Acquisition time \u2013 894 s; Number of km \u2013 23.3 km; Fa\u0301tima\u2013Pombal: Acquisition time \u2013 1180 s; Number of km \u2013 40 km; Pombal\u2013Alfarelos: Acquisition time \u2013 1037 s; Number of km \u2013 29 km; Alfarelos\u2013CoimbraB: Acquisition time \u2013 900 s; Number of km \u2013 18.7 km;"
        ]
    },
    {
        "image_filename": "designv11_12_0000597_978-3-540-73129-0-Figure6.60-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000597_978-3-540-73129-0-Figure6.60-1.png",
        "caption": "Fig. 6.60. Parallel gripper with interchangeable jaws [88]",
        "texts": [
            " Small shape memory actuators can apply relatively high forces and strokes, can be well integrated into the grippers mechanical structure, and do not emit particles into the clean-room environment. Only small SM elements are required to actuate miniature grippers. Hence, fast opening and closing times of the gripping jaws can be expected. Two examples of micro grippers built according to these design principles are described in the next paragraphs. Parallel Gripper. The parallel gripper mechanism in Fig. 6.60 consists of a single piece of plastic with interchangeable jaws. For this prototype, flexure hinges were cut out of the material by micro milling, but for higher production quantities injection moulding is possible. The jaws are closed by heating the SM wires (length 12mm, diameter 0.15mm). The gripping mechanism translates the actuators stroke into a movement of the jaws of 1.5mm with a gripping force of 0.15N. The gripping-/loosening time averages by 0.3 s. SMA Actuated Miniature Silicon Gripper"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002184_cdc.2009.5399544-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002184_cdc.2009.5399544-Figure1-1.png",
        "caption": "Fig. 1. Visual sensor footprint",
        "texts": [
            " In this set-up, we assume the values Wq to be constant in time. 2) Modeling of the visual sensor nodes: Each mobile agent i is modeled as a point mass in Q, with location ai := (xi, yi) \u2208 Q. Each agent has mounted a pan-tilt-zoom camera, and can adjust its orientation and focal length. The visual sensing range of a camera is directional, limited-range, and has a finite angle of view. Following a geometric simplification, we model the visual sensing region of agent i as an annulus sector in the 2-D plane; see Figure 1. The visual sensor footprint is completely characterized by the following parameters: the position of agent i, ai \u2208 Q, the camera orientation, \u03b8i \u2208 [0, 2\u03c0), the camera angle of view, \u03b1i \u2208 [\u03b1min, \u03b1max], and the shortest range (resp. longest range) between agent i and the nearest (resp. farthest) object that can be recognized from the image, rshrt i \u2208 [rmin, rmax] (resp. rlng i \u2208 [rmin, rmax]). The parameters rshrt i , rlng i , \u03b1i can be tuned by changing the focal length FLi of agent i\u2019s camera. In this way, ci := (FLi, \u03b8i) \u2208 [0, FLmax] \u00d7 [0, 2\u03c0) is the camera control vector of agent i"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000569_tie.2008.2009991-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000569_tie.2008.2009991-Figure1-1.png",
        "caption": "Fig. 1. Two-link manipulator model.",
        "texts": [
            " In Section III, the relation between the 0278-0046/$25.00 \u00a9 2009 IEEE PD controller and time delay is described. Sections IV and V present strategies of decreasing and avoiding the destabilization. The experimental results are shown in Section VI. Finally, conclusions and future work are described in Section VII. This paper realizes a stabilization of an inverted pendulum which is placed on the end of the two-link manipulator. The manipulator is modeled on the relationship between the joint angles and the length of link as shown in Fig. 1. The position of the first drive joint is the origin of the world coordinate, and two links have masses m1 and m2 on the edge and lengths l1 and l2, respectively. \u03b81 and \u03b82 are the relative joint angles. On the basis of the model, the relation from the joint angle \u03b8 = (\u03b81, \u03b82) to the end position of the manipulator x = (x, y) is shown as follows: x = l1 sin \u03b81 + l2 sin(\u03b81 + \u03b82) (1) y = \u2212 l1 cos \u03b81 \u2212 l2 cos(\u03b81 + \u03b82). (2) The relations between workspace and joint coordinate space in velocity and acceleration are derived as follows by differentiating the earlier equations with respect to time x\u0307 =Jaco\u03b8\u0307 (3) x\u0308 =Jaco\u03b8\u0308 + J\u0307aco\u03b8\u0307 (4) where Jaco is the Jacobian matrix"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003467_gt2013-95424-Figure15-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003467_gt2013-95424-Figure15-1.png",
        "caption": "Figure 15: V1 CFD domain for heat transfer analysis",
        "texts": [],
        "surrounding_texts": [
            "In this section the validated methodology is used to predict the wear for all three seal configurations shown in Figure 4. The first step was to determine the heat transfer coefficients for the seal fins. Subsequently these results were used in the FEA to simulate the wear an the seal strips."
        ]
    },
    {
        "image_filename": "designv11_12_0000715_s11249-007-9287-9-Figure17-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000715_s11249-007-9287-9-Figure17-1.png",
        "caption": "Fig. 17 Schematic illustration of CVT variator with longitudinal grooves",
        "texts": [
            " (2) In an actual variator, the magnitude of the required traction coefficient varies depending on the transmission ratio. Increasing the traction coefficient is desired most in the vicinity of a transmission ratio of 1.2:1. Accordingly, the increasing effect can be obtained even if the formation of the longitudinal grooves is limited only to the area of the disks used as the traction surfaces near a transmission ratio of 1.2:1. In that case, it would also avoid usage of that area under the high contact pressure condition of a low ratio. As shown in Fig. 17, longitudinal grooves were continuously formed in the surfaces of the input/output disks at a groove pitch of 200 lm and with a 4-mm radius of curvature of the convex portion. The maximum depth of the grooves was 2 lm near a transmission ratio of 1.2:1 and the depth was gradually reduced in an arc-like pattern toward the ratio positions of 1.8:1 and 0.82:1. 5.2 Experimental Procedure In an actual toroidal CVT, a loading cam mechanism produces loading force in proportion to the input torque. In order to measure the traction coefficient of the toroidal variator used in this study, a test box [14] was built with an increased hydraulic loading system that allowed the loading force to be set at an arbitrary level"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001695_10402000903420803-Figure3-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001695_10402000903420803-Figure3-1.png",
        "caption": "Fig. 3\u2014Upstream (left), downstream (middle), and side (right) view of hybrid brush seal (HBS) with brushes on the upstream and the Electric Discharged Machining (EDM) spring elements on the downstream side.",
        "texts": [
            " Continuous rubbing not only will wear the bristle tips but, most importantly, it induces excessive frictional heating that causes rotor thermal growth and rotor bow. This leads to more friction and more rotor growth until the seal fails, a typical thermal instability. If a seal is installed with too little interference, however, it can result in unacceptable leakage, thus negating the seal usefulness. The novel hybrid brush seal (HBS; Justak and Crudgington (10)) integrates spring-supported cantilever pads into the brush seal bristle bed, as shown in Fig. 3. Upon rotor spinning, the soft-spring-supported pads or shoes lift off due to the generation of a hydrodynamic gas film underneath. Hence, the HBS can withstand shaft forward and backward rotations and, most importantly, has noncontacting operation without wear and local thermal distortion. Justak and Crudgington (10) showed that D ow nl oa de d by [ U ni ve rs ity o f A uc kl an d L ib ra ry ] at 2 0: 15 0 6 D ec em be r 20 14 shoed brush seal, a predecessor to the HBS, maintains an effective clearance of 0"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001009_1.3085886-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001009_1.3085886-Figure2-1.png",
        "caption": "Fig. 2 Electromagnetic actuator",
        "texts": [
            " Moreover, the validity of the proposed method for vibration control and unbalance estimation is confirmed experimentally for both linear and nonlinear rotor systems. 2 Vibration Control and Unbalance Estimation by Disturbance Observer 2.1 Nonlinear Rotating Shaft Model and Equation of Motion. In this paper, vertical Jeffcott rotor shown in Fig. 1 is considered. Coordinate system O-xyz is used. In this coordinate system, z axis is on the bearing centerline and x and y axes are on the plane of disk rotation. Electromagnetic actuators are set as shown in Fig. 2. Considering control forces fmx and fmy of electromagnetic actuators, the equation of motion 9 of the nonlinear Jeffcott rotor is shown as mx\u0308 + cx\u0307 + kx + Nx = me 2 cos t + fmx my\u0308 + cy\u0307 + ky + Ny = me 2 sin t + fmy 1 Here, m is the mass of the disk, c is the damping coefficient, k is the spring constant of the elastic shaft, is the rotational speed, and t is time. e is the static unbalance of the disk, and an unbalance force with magnitude me 2 acts on the disk. Nx and Ny show the system nonlinearities caused by various factors such as clearance in a ball bearing",
            " These experimental results coincide qualitatively with the results of numerical simulation shown in Figs. 4 and 5. JUNE 2009, Vol. 131 / 031010-5 hx?url=/data/journals/jvacek/28900/ on 04/12/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use T t s 1 , a F e 0 Downloaded Fr 4.3 Unbalance Estimation Using Disturbance Observer. he magnitude and phase of unbalance were estimated from the ime history of disturbance observer force fobs and reference pulse ignal obtained from vibration suppression experiment. Figure 2 a shows the complex FFT spectra of disturbance control force fobs for the case of nonlinear rotor system. These disturbance con- 0 500 1000 1500 2000 2500 0.0 0.2 0.4 Balancing (1\u00b5m) Balancing (10\u00b5m) Without D.O. Rotational speed \u03c9 (rpm) D ef le ct io n r m ax (m m ) ig. 9 Effect of sensor resolution on accuracy of unbalance stimation 31010-6 / Vol. 131, JUNE 2009 om: http://vibrationacoustics.asmedigitalcollection.asme.org/pdfaccess.as trol force fobs data were obtained at point A of rotational speed =833 rpm"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003761_9781118818428-Figure5.6-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003761_9781118818428-Figure5.6-1.png",
        "caption": "Figure 5.6 Karabulut crank-slider mechanismmodified by inclusion of variable offset z\u2032 between slider track and crank-pin",
        "texts": [
            " On the other hand, one of the conditions for minimum friction is self-evident: minimum loads on moving surfaces in contact. A step in this direction \u2013 regardless of the detail of the mechanical embodiment \u2013 is achievement of minimum cyclic pressure swing per unit of indicated cycle work. The Larque engine demonstrates smooth running down to low rpm. This is consistent with the kinetic energy in a slowly rotating flywheel readily overcoming the compression stroke \u2013 in other words, a low pressure ratio. The drive mechanism of the Karabulut engine merits further study. Figure 5.6 shows the mechanism modified somewhat by incorporation of variable offset z\u2032 of slider track from crank-pin. Applying Pythagoras\u2019 theorem gives piston gudgeon pin location yp above the crank-shaft centre line: yp = rcos\u03c6 +\u221a {lrp 2 \u2212 (sin\u03c6 \u2212 d)2} (5.7) Angle \u03b4 made by the slider with the horizontal is: \u03b4 = atan{(cos\u03c6 + Y)\u2215(X \u2212 d + sin\u03c6)} The inclination of rod rd to the horizontal is \u03b4 \u2212 \u03c2. A design aim would be to keep rod lrd as nearly vertical as possible, so that instantaneous height yd of the gudgeon pin driving the displacer is given approximately by: yd = \u2212Y + rdsin(\u03b4 \u2212 \u03c2) + lrd (5"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003581_aim.2011.6027001-Figure8-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003581_aim.2011.6027001-Figure8-1.png",
        "caption": "Fig. 8. Structures of the robot\u2019s parts",
        "texts": [
            " Each unit has a space called a chamber between the bellows and the artificial muscle. When air is supplied to the chamber, each unit is contracted. The air tube, electrical wires, and a treatment tool are arranged inside the bellows to avoid any hindrance in movement. B. Fabrication of the Prototype In order to confirm that there is adequate space for the electrical wires and treatment tool in the bellows, we fabricated a prototype. Each unit has a single air tube except the last unit, which has two tubes. Figure 8 shows the structures of the robot\u2019s parts. Table I shows the specifications of one unit, and Fig. 9 shows the exterior of the prototype, and Fig. 10 shows its head. The space for the treatment tool of a conventional endoscope is an aperture about 3 mm in diameter and the diameter of conventional endoscope is 12.5 mm. From Table I, it is noted that outside and inside diameter of the prototype are almost the same as outside and aperture\u2019s diameter of a conventional endoscope. Therefore, the robot can attach conventional endoscope and supply the space needed for a treatment tool"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001968_09507110902836879-Figure10-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001968_09507110902836879-Figure10-1.png",
        "caption": "Figure 10. Schematic of wire extension and interaction with laser beam.",
        "texts": [
            " This being so, and taking into account the desirability of experimentally investigating laser absorption characteristics in order to understand the vaporization of droplets and the melting of base metal, the discussion that follows concerns the absorption characteristics for laser beams, using the physical characteristics of the arc and focusing primarily on wire melting in a steady state, and there is also an attempt to quantify these. Wire extension during MIG arc welding and the present hybrid welding is shown in the schematic drawing in Figure 10. In the case of MIG arc welding, the wire melting rate is determined by both the heat input from the arc and also the resistance exothermy in the wire. It is known that the former of these is proportional to the arc current and the latter is proportional to the square of the arc current. If the arc current is known, the wire melting rate can be expressed by Equation (1)32 Vw \u00bc a\u00b7IM \u00fe b\u00b7EX\u00b7I2 M: \u00f01\u00de Here, Vw, wire melting rate; a and b, proportional constants; IM, arc current; and EX, wire extension length"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002293_s00707-009-0262-4-Figure3-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002293_s00707-009-0262-4-Figure3-1.png",
        "caption": "Fig. 3 a Cervix under uniform radial tension, b Upper half of cervix showing tensile stress",
        "texts": [
            "18) Let J (t, \u03b1) be the creep compliance corresponding to the stress relaxation modulus G (t, \u03b1), at a constant value of \u03b1. It is noted here, for later use, that Eq. (3.14) can be inverted to give \u03c3\u0302 (\u03b5(t)) = \u03c3(0)J (t, \u03b1) + t\u222b 0 J (t \u2212 s, \u03b1) d ds \u03c3(s)ds. (3.19) Figure 2 shows a uterus, fetal head, and birth canal [20, p. 139]. During birth, the uterus contracts and applies pressure to the amniotic fluid and thereby to the fetus. The fetal head then applies pressure to the opening at the birth canal. Figure 3a shows a ring of tissue at the opening of the uterus at the birth canal (the cervix). The cervix is subjected to a force distribution that acts outward as a result of tension from the uterus during contractions and pressure from the fetal head. This dilates the cervix and allows the fetus to pass through the birth canal. Leppert [21] described the cervix as a predominantly extracellular connective tissue matrix whose biomechanical response is viscoelastic. In [21], Leppert further discussed the complex changes in the cervix during pregnancy and birth, a process referred to as \u201cripening\u201d",
            " As the cervix is viscoelastic, it is interesting to use the notion of a viscoelastic clock model to represent aspects of this ripening process. It has already been shown in Fig. 1 that the viscoelastic clock model can represent softening. The example presented here shows some additional consequences of the model in the context of birth. For convenience in illustrating the underlying idea, the force distribution is assumed to be radially symmetric. The radial force per unit length is P(t). This induces tensile stress \u03c3(t) in the ring as shown in Fig. 3b. The initial radius of the ring is ro, the cross sectional area of the ring is A, and the radial displacement is u(t). The circumferential strain in the ring is given by \u03b5(t) = u(t) ro , (4.1) and the force balance equation gives \u03c3(t) = P(t) ro A . (4.2) Assume that during a contraction the radial force distribution on the ring, P(t), can be described by the constant rate radial loading and unloading history described by, P(t) = \u23a7\u23a8 \u23a9 Pmax t\u2217 t, t \u2208 [ 0, t\u2217 ] , Pmax t\u2217 (2t\u2217 \u2212 t) , t \u2208 [ t\u2217, 2t\u2217 ] , 0, t > 2t\u2217, (4"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000150_robot.2005.1570677-Figure12-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000150_robot.2005.1570677-Figure12-1.png",
        "caption": "Fig. 12 MRI-coordinate system",
        "texts": [
            " It was verified that there was no discernible magnetic attraction force on the prototype when the motor was operating inside the MRI gantry, which suggests that the proposed motor is MR safe [3]. In this study, MR-compatibility was evaluated with respect to two points: the influence of MRI\u2019s strong magnetic field and the influence of motor\u2019s operation on MRI imaging. The former was evaluated through thrust force measurements, while the latter was evaluated through the signal-to-noise ratio (SNR) of the MRI images. The MRI coordinate system defined in Fig. 12 is used to describe the measurement conditions. The origin of the system is set at the edge of the RF-coil that was positioned at the center of the gantry. The X-axis is defined as the direction parallel the direction of the MRI\u2019s static magnetic field. Since the materials used in the prototype are paramagnetic, the prototype is not affected by the surrounding magnetic field when the motor is not operating. However, when the motor is operating, the surrounding magnetic field induces a Lorenz force on electric current in the slider, as was discussed in the theoretical evaluation section"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001866_bf00251592-Figure14-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001866_bf00251592-Figure14-1.png",
        "caption": "Fig, 14. Special case, common to Watt=l, Stephenson-1 and Stephenson-2 mechanism",
        "texts": [
            " It follows that there are 4 for which u is fixed and t takes two values. The total number of 28 double points agrees with the value obtained by considering the genus. The number of double points on a curve of order 16 and genus 7 is 89 x 15 x 1 4 - 7 =98, and I and J account for 35 each (28 for an 8-fold point, 6 for a 4-fold point in the neighborhood and 1 for a 2-fold point in the neighborhood), so the expected number is 9 8 - ( 2 x 35)=28. The three mechanisms have a special case in common, as shown in Fig. 14. It is obtained from Stephenson-1 by making D and F coincide, from Stephenson-2 by making C and D coincide, and from Watt-1 by making D and F coincide. It is interesting to see what happens to the singular foci and double points when we specialize. Singular Foei. In the special case, we have 3 at A, 1 each at E and K, and 2 others. (i) Stephenson-1. H coincides with A, and one of the three singular loci, obtained in section 4 (iv) of Chapter II of this paper (page 45) coincides with A. 5 Arch. Rational Mech"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001370_icsens.2009.5398292-Figure8-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001370_icsens.2009.5398292-Figure8-1.png",
        "caption": "Fig. 8. Overview of the experimental equipment",
        "texts": [
            " Figure 7 shows the relationship between the length and the circumference of each MPA. Table II shows the estimated length b and turns n of the fiber of each MPAs. The red lines in Fig. 7 show the estimated length L using the estimated b and n in (8). The estimated lengths are closed enough with the measured results and the values of b and n will also be used in next experiments. The flexible sensor is installed to MPA as shown in Fig. 1. The validity of the MPA length estimation method is examined by comparing the estimated length and the true length. Figure 8 and 9 show the experiment devices and the configuration of control system. Control system consists of an air compressor, an air regulator, a control computer, an A/D convertor, a D/A convertor, a DC power source, a MPA and a flexible sensor. The estimation length is caluclated from the circumference displacement measured by the flexible sensor using (8). The true value is measured by the potentiometer. L=150[mm], \u03c6=15[mm] MPA is used in this experiment. The air pressure in MPA is controlled with sine and square signals"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002123_iros.2009.5354225-Figure5-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002123_iros.2009.5354225-Figure5-1.png",
        "caption": "Fig. 5. Robots following a sweep line with \u03b4 overlap and splitting into two sweep lines at a critical point. Robots with solid disks are moving towards future positions marked as robots with dashed disks. Note that the four robots require additional robots to reach the dashed positions.",
        "texts": [
            " In practice, block lines are often at the position right after the split on pk,i,j , in particular when one or more of Ci, Cj , Ck are points. Fig. 4 shows how multiple robots maintain a line oriented towards the left site, i.e. robots are placed uniformly on a sweep line starting from the left site and placed at distance r \u2212 \u03b4 from each other with the first robot having distance r\u2212\u03b4 2 from the left site. As the distance between the obstacles grows another robot will have to be added. Due to the left bias this is trivially achieved. Following the split lines is also easy and an illustration is shown in figure 5. A further discussion of the significance of \u03b4 and how to arrange robots on the line is found in VI. IV. IMPLEMENTATION The actual implementation to use the line clearing approach to construct a surveillance graph from an occupancy grid map proceeds in several stages. First, to smoothen the map we convert it to a polygon by computing its \u03b1 shape using the CGAL library [14]. These shapes are frequently used to reconstruct the shape of a dense set of points. Once we get the polygon boundary from the occupied grip points we apply the Ramer-Douglas-Peucker line-simplification algorithm [15] to get a polygon boundary with less line segments",
            " Given the sequence of vertices we can construct a sequence of moving lines. The first vertex is cleared by blocking one edge and clearing it coming with a new sweep line starting from that edge, leading to blocks on all its edges. From there on every next vertex is a movement of one blocking line towards the critical point for the direction the line is coming from. These continuously moving lines can be followed by a team of robots. To account for localization and navigation errors robots can be spaced with sensors overlapping as seen in Figure 5. The \u03b4 parameter also helps to offset possible approximation errors during the conversion of the grid map into a polygon. To validate the algorithm from section IV we ran it in a variety of configurations on the grid maps from [2]. The first grid map was obtained with a Pioneer P3AT mobile platform equipped with a SICK PLS200 laser range finder. The robot was driven through the 2nd floor of the UC Merced Engineering and Science building to collect laser data. The data was then used to build the map with the Gmapping software [16]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001287_cdc.2007.4434251-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001287_cdc.2007.4434251-Figure1-1.png",
        "caption": "Fig. 1. Variables in the homicidal chauffeur game. The shaded region around the pursuer indicates its capture disc.",
        "texts": [
            " We assume that all the players have unlimited sensing capabilities. The pursuers have identical motion abilities and possess greater speed than that of the evader. However, the evader can make arbitrarily sharp 1-4244-1498-9/07/$25.00 \u00a92007 IEEE. 4857 turns while the pursuers cannot turn more than a minimum turning radius. We assume that the instantaneous position and velocity of the evader is available to all pursuers. Let e(t) and pk(t), for k \u2208 {1, . . . , N}, denote the positions of the evader and the kth pursuer in R 2 at time t, as shown in Figure 1. Let ve and vp denote the speeds of the evader and all the pursuers, respectively. Let ve and vp,k denote the velocity vectors of the evader and the kth pursuer, respectively. Given a minimum turning radius \u03c1 > 0, the mathematical model for this problem can be described as follows [2]. For evader: e\u0307x(t) = ve cos \u03b8e(t), e\u0307y(t) = ve sin \u03b8e(t). For pursuers: p\u0307k,x(t) = vp cos \u03b8p,k(t), p\u0307k,y(t) = vp sin \u03b8p,k(t), \u03b8\u0307p,k = vp \u03c1 up,k, (1) where \u03b8e(t) and \u03b8p,k(t) are respectively the angles made by the velocity vectors of the evader and of the kth pursuer with reference to a global X axis",
            " In that time, the evader\u2019s reachability set is the dotted circle of radius \u03b3(lp + 2\u03c0\u03c1), centered at e(0), as shown in Figure 4. Thus, to compute lp, we impose the condition that the minimum distance between the evader\u2019s reachability set and the circular portion of the path of p1 must be (1+\u03b3)lst. Using elementary geometry, the equation (4) for lp follows. We now prove the following property of the Align phase. Lemma IV.3 (Align phase) The Align phase of CONFINE strategy terminates after finite time with the evader aligned with p1 and \u2016p1 \u2212 e\u2016 = (1 + \u03b3)lst. Proof sketch of Lemma IV.3: Consider the system as shown in Figure 1 with k = 1. At the end of the Pre-align phase, angle \u03c6 = 0. Using the control law for Align phase, we show that \u03c6 = 0, for all subsequent time instants as long as distance L(t) \u2265 \u03b3\u03c1. We now obtain an upper-bound for the distance between the first pursuer and evader at the end of the Chase phase. Lemma IV.4 (Upper bound on d) Let d denote the distance \u2016p1 \u2212 e\u2016 at the end of the Chase phase. Then, d \u2264 \u03b3 \u221a 1 + \u03b3 1 \u2212 \u03b3 lst, or, equivalently, for \u00b5 , d\u22121 \u03c1 and lst = 2\u03c0\u03b3\u03c1 1\u2212\u03b3 , \u00b5 \u2264 2\u03c0\u03b32 (1 \u2212 \u03b3) \u221a 1 + \u03b3 1 \u2212 \u03b3 \u2212 1 \u03c1 "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000388_14644193jmbd97-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000388_14644193jmbd97-Figure1-1.png",
        "caption": "Fig. 1 Grinding spindle and angular contact ball bearing",
        "texts": [
            " The overall stiffness of deflection can be found by the summation of two relative movements: outer raceway/ball and ball/inner raceway as given in equation (6) below. K \u00bc 1 \u00f01/Ki\u00de 2=3 \u00fe \u00f01/Ko\u00de 2=3 3=2 \u00f06\u00de where Ki and Ko are inner and outer raceways to ball contact stiffnesses, respectively, [5]. The objective here is to calculate the net force on a bearing due to the displacement of the spindle centre since this force can then be used in the equations of motion to observe the motion of the spindle centre. In order to calculate the total force, the deflection at the ith ball in Fig. 1 will be calculated first and this will be used in the calculation of the total force. As seen in Fig. 1, the ball is rotating between the inner and outer rings. During this rotation, the ball is continuously in contact with different points in the circular grooves in each race. In the initial position, without any preload, the loci of raceway groove centres of curvature will produce circles as shown in Fig. 2. The figure is a three-dimensional representation of the inner and outer ring raceway groove curvature loci, with two-dimensional crosssection of an inner and outer race overlaid on it. Proc",
            " The deflections and contact angles for the RHS bearing are, respectively \u00f0di\u00deR \u00bc Bdb cos\u00f0a0\u00de \u00fe x cos\u00f0ui\u00de \u00fe y sin\u00f0ui\u00de b1\u00f0fi\u00de cos\u00f0ui\u00de \u00feb1\u00f0ci\u00de cos\u00f0fi\u00de sin\u00f0ui\u00de \u00feR\u00f01 cos\u00f0fi\u00de\u00de cos\u00f0ui\u00de \u00feR\u00f01 cos\u00f0ci\u00de\u00de sin\u00f0ui\u00de 0 BBBBBB@ 1 CCCCCCA 2 \u00fe Bdb sin\u00f0a0\u00de \u00fe z0 \u00fe z \u00fe b1\u00f01 cos\u00f0fi\u00de\u00de cos\u00f0ui\u00de \u00fe b1\u00f01 cos\u00f0ci\u00de\u00de cos\u00f0fi\u00de sin\u00f0ui\u00de \u00f0fi\u00deR cos\u00f0ui\u00de \u00fe \u00f0ci\u00deR sin\u00f0ui\u00de 0 BBB@ 1 CCCA 2 2 66666666666666664 3 77777777777777775 1=2 Bdb \u00f021\u00de \u00f0ai\u00deR \u00bc tan 1 Bdb sin\u00f0a0\u00de \u00fe z0 \u00fe z \u00fe b1\u00f01 cos\u00f0fi\u00de\u00de cos\u00f0ui\u00de \u00fe b1\u00f01 cos\u00f0ci\u00de\u00de cos\u00f0fi\u00de sin\u00f0ui\u00de \u00f0fi\u00deR cos\u00f0ui\u00de \u00fe \u00f0ci\u00deR sin\u00f0ui\u00de 0 BB@ 1 CCA Bdb cos\u00f0a0\u00de \u00fe x cos\u00f0ui\u00de \u00fe y sin\u00f0ui\u00de b1\u00f0fi\u00de cos\u00f0ui\u00de b1\u00f0ci\u00de cos\u00f0fi\u00de sin\u00f0ui\u00de \u00feR\u00f01 cos\u00f0fi\u00de\u00de cos\u00f0ui\u00de \u00feR\u00f01 cos\u00f0ci\u00de\u00de sin\u00f0ui\u00de 0 BBBB@ 1 CCCCA 2 666666666666664 3 777777777777775 \u00f022\u00de The deflection for the ith ball was calculated in the previous section. The force due to this deflection for a single ball can be found and then the total force in the bearing can be calculated as this ball rotates around the inner ring. In the model, the contacts of balls to the inner and outer races are represented by non-linear contact springs, i.e. balls act as massless springs. The elastic model of the bearing is represented in Fig. 10. To calculate the total contact force on the ith ball in Fig. 1, the reference axes should be determined and the total deflection, and hence the forces with respect to these axes, should be calculated. The deflection for each ball can be calculated as described earlier. Having calculated the deflection for the ith ball in its contact direction, the force in the same direction Proc. IMechE Vol. 222 Part K: J. Multi-body Dynamics JMBD97 # IMechE 2008 at UNIV PRINCE EDWARD ISLAND on August 5, 2015pik.sagepub.comDownloaded from (Fi) can easily be found. This force can be split into two components in the radial (Fri) and axial (Fai) directions for angular contact ball bearings"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003246_1.4004116-Figure8-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003246_1.4004116-Figure8-1.png",
        "caption": "Fig. 8 Geometric sketch of the right tooth profile of noncircular gear generated by shaper cutter",
        "texts": [
            "asme.org/about-asme/terms-of-use The length of profile normal BN can be derived by jBNj \u00bc jB2Bj cos a \u00bc B1B _ cos a, where a is the profile angle of the rack. Similarly, the length of profile normal AN in Fig. 6 can be represented as jANj \u00bc A2A _ cos a \u00bc A1A _ cos a. If the intersection point of tooth profile and profile normal is outside the pitch curve of noncircular gear as shown in Fig. 6, the vector angle of AN is u\u00fe l a. On the contrary, if the point is inside the pitch curve as shown in Fig. 8, the vector angle of AN is represented as u\u00fe l a\u00fe p. Using Eq. (24), the formulas of left tooth profile of noncircular gears can be derived as follows xNL \u00bc r cos u6 A1A _ cos a cos\u00f0u\u00fe l a\u00de yNL \u00bc r sin u6 A1A _ cos a sin\u00f0u\u00fe l a\u00de ( (25) where tan l \u00bc r=r0\u00f0u\u00de. Similarly, the right profile formulas are represented as follows xNR \u00bc r cos u A1A _ cos a cos\u00f0u\u00fe l\u00fe a\u00de yNR \u00bc r sin u A1A _ cos a sin\u00f0u\u00fe l\u00fe a\u00de ( (26) The upper and lower signs in Eqs. (25) and (26) correspond to the points outside and inside the pitch curve on the tooth profile"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002947_17452759.2013.790599-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002947_17452759.2013.790599-Figure1-1.png",
        "caption": "Figure 1. Representation of non-supported feature and the need for support material.",
        "texts": [
            " This stage is called the post-processing stage (Volpato 2007, Gibson et al. 2010). In addition to the waste of material, it is possible that the removal procedure deforms or breaks some object features, compromising the quality and the precision of the final product. It is important to note that even though in some cases the support material is a water-soluble composition, the cost of the material is extremely high, making the process more expensive. This subject has already been emphasised in others studies, such as (Crockett et al. 1999). Figure 1 illustrates two very common non-supported features that are found in plastic parts design (a cantilever, *Email: marloncunico@yahoo.com.br Virtual and Physical Prototyping, 2013 Vol. 8, No. 2, 127 134, http://dx.doi.org/10.1080/17452759.2013.790599 # 2013 Taylor & Francis or snap-fit, and a bridge), showing the negative surface that depends on the current deposition of a support material. Additionally, it can be observed that these two features are extreme cases of negative surfaces because they are aligned with the construction process orientation (horizontal)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002198_pime_conf_1967_182_045_02-Figure21.4-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002198_pime_conf_1967_182_045_02-Figure21.4-1.png",
        "caption": "Fig. 21.4. Mechanism of grease redism\u2018bution in ball-bearing assembly",
        "texts": [
            " SCARLETT The grease in the outer race track is forced ahead of the balls and as the balls rotate they carry a thin film of grease (dependent on the clearance in the cage ball pockets) around with them. Some of this grease is deposited on the track edges of the inner race where it travels up the track by centrifugal force and is then pumped under the cage bore as already described. During this process a small amount of grease is splashed on to the cage bore between each ball (the region where no grease was packed originally) where it adheres and forms pads of grease as shown in Fig. 21.4. The accommodation of excess grease into static pads takes some hours to complete, as indicated by the running period before a normal settled temperature is reached. Unsatisfactory greases do not form complete pads but circulate continuously and the bearing temperature then remains high. When the running temperature eventually falls to a reasonable level (usually quite suddenly), circulation under the cage bore ceases. The pads of grease are now firmly attached to the cage bore and accommodated in a static position"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002975_c3sm27906e-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002975_c3sm27906e-Figure2-1.png",
        "caption": "Fig. 2 Geometry used to determine the liquid crystal force contribution to the colloidal nodes, based on the local torque, Gz, applied to the liquid crystal by the colloidal surface.",
        "texts": [
            " Pu bl is he d on 2 4 A pr il 20 13 o n ht tp :// pu bs .r sc .o rg | do i:1 0. 10 39 /C 3S M 27 90 6E colloid. To accomplish this, we calculate Gz at the midpoint locations between particle nodes, and apply forces to pairs of adjacent nodes i, and i + 1, according to F i \u00bc Gzn\u0302 l F i\u00fe1 \u00bc Gzn\u0302 l ; (14) in order to produce an equal and opposite torque acting on the local surface element between the two nodes. Here l is the distance between the particle nodes, and n\u0302 is the surface normal at the midpoint location (see Fig. 2). This results in zero net force acting on the particle, but may lead to deformations of the particle shape. In addition to the forces that result directly from the particle surface anchoring, distortions in the liquid crystal generate forces that would be felt even by particles with neutral surface anchoring. These forces are calculated using the symmetric liquid crystal stress tensor according to Fa,i \u00bc nbsabDSi, (15) where DSi is the portion of the colloidal surface represented by node i, and n is the surface normal"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001954_j.jsv.2009.01.057-Figure3-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001954_j.jsv.2009.01.057-Figure3-1.png",
        "caption": "Fig. 3. Nodal displacement initiating fold mode: (a) the top members folding mode as g1, (b) the middle members folding mode as g2, (c) the bottom members folding mode as g3.",
        "texts": [
            " For the static problem the folding patterns can be found by considering the eigenproblem of the Jacobian. For the folding patterns that are identified no contact between nodes is assumed. Hence for the localised folding as shown below although the gap between nodes does reduce to zero no contact occurs. For all singular paths at the BP the Jacobian equates to zero, i.e. when the increasing deformation v\u03041 reaches a critical value det J \u00bc 0. For the particular case when det J\u00f0v\u0304BP1 \u00de \u00bc 0 of Eq. (18) three eigenvectors are obtained gi and shown in Fig. 3. The different collapse modes or folding patterns corresponding to these eigenvectors are shown in Fig. 3(a)\u2013(c) as follows: (a) Eigenvector g1 shows localised folding of the top member. (b) Eigenvector g2 shows localised folding of the middle member. (c) Eigenvector g3 shows localised folding of the bottom member. These folding patterns are those identified by the bifurcation paths. The vertical gap between nodes prior to loading is 2h. During the folding process when nodes come into contact the vertical gap reduces to zero. In the following four subsections each of the eigenvectors shown in Fig. 3(a)\u2013(c) are discussed. Because of symmetry of the folding pattern the discussion is limited to the left-hand half of the truss only. The static behaviour compares with the dynamic behaviour described below and behaviour observed in experiments as seen in Figs. 6, 8,12. The mode shape for this eigenvector shows proportional folding of all the members and corresponds to the primary unstable equilibrium path shown in Fig. 2(a). When the top node (node 1) has displaced by a vertical distance 3h then, due to proportional displacement of all nodes, all the members are horizontal. This corresponds to the point in Fig. 2(b) when the primary path graph crosses the horizontal axis at 3h. 2.4.2. Folding mode g3: localised folding initiated by bottom member (bifurcation path) The mode shape for this eigenvector corresponds to the bottom member folding shown in Fig. 3(c). When this members folds the gap between nodes 2 and 4 reduces to zero. For this fold pattern the members joining nodes 2 to 3 and 4 to node 3 will be collinear and adjacent. 2.4.3. Folding mode g2: localised folding initiated by middle member (bifurcation path) The mode shape for this eigenvector corresponds to the middle member folding shown in Fig. 3(b). When this member folds the gaps between nodes 1 and 3 and nodes 2 and 4 reduces to zero. Hence for this fold pattern the members joining nodes 1 to 2, nodes 3 to 2 and nodes 3 and 4 will be collinear and adjacent (Table 1). ARTICLE IN PRESS I. Ario, A. Watson / Journal of Sound and Vibration 324 (2009) 263\u2013282270 2.4.4. Folding mode g1: localised folding\u2014top member (bifurcation path) The mode shapes for these eigenvectors show similar behaviour and correspond to the top member folding shown in Fig. 3(a). When the top member folds the gap between nodes 1 and 3 reduces to zero. For these fold patterns the members joining nodes 1 to 2 and 3 to node 2 will be collinear and adjacent. The dynamic analysis equation for the folding truss combines mass, damping and nonlinear stiffness F\u00f0v\u00de in the following equation: M \u20ac\u0304v\u00f0t\u00de \u00fe C _\u0304v\u00f0t\u00de \u00fe F\u00f0v\u0304\u00f0t\u00de\u00de \u00bc 0, (25) where M 2 RN N is the mass matrix; C 2 RN N is the damping matrix; F\u00f0 \u00de is the nonlinear stiffness vector; f \u20ac\u0304vi\u00f0t\u00deg T \u00bc \u20ac\u0304v\u00f0t\u00de 2 RN is normalised acceleration; f _\u0304vi\u00f0t\u00deg T \u00bc _\u0304v\u00f0t\u00de 2 RN is the velocity; fv\u0304i\u00f0t\u00deg T \u00bc v\u0304\u00f0t\u00de 2 RN is the normalised displacement and N is the total number of degrees of freedom in the system",
            " The individual stiffness of each (real) member is unchanged; however, due to an increase in system stiffness, caused by contacting nodes, a value of EA \u00bc 1000 for the dummy elements was found to give good results for the folding pattern behaviour of the pantographic truss. The analysis allows for separation of nodes during the dynamic behaviour. As such the length of the dummy element, \u2018\u0302dummy, can increase from zero. As soon as there is a gap between nodes the stiffness of the dummy element becomes zero again. Hence there is no \u2018extra stiffness\u2019 during any noncontact behaviour occurring after any contact behaviour. If local snap-through behaviour occurs there will be contact for the restabilised state. For each of the folding processes shown in Fig. 3(a)\u2013(c) contact occurs between nodes. When local folding is initiated by the top member this initially leads to contact between nodes 1 and 3. Further loading will eventually result in the bottom member folding and contact between nodes 2 and 4. When local folding is initiated by the bottom member there is contact between nodes 2 and 4. Further loading will eventually result in the top member folding and contact between nodes 1 and 3. For middle member initiated folding results in contact between nodes 1 and 3 plus nodes 2 and 4"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001065_13506501jet487-Figure10-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001065_13506501jet487-Figure10-1.png",
        "caption": "Fig. 10 Cavitated bearing, \u03c0 film",
        "texts": [
            " In the calculation of the journal centre velocities, the average angular velocity of the journal and bush differs between the big-end bearing analysis and the main bearing analysis in their respective computational frame [39]. For the big-end bearing w = \u03c9 2 ( 1 + Ra L cos(\u03b8) ) (13) and for the main bearing w = \u03c9 2 (14) The initial value of the eccentricity ratio used is zero. The product of the journal centre velocity and the time interval will give the corresponding new eccentricity ratio for each step. The program outputs bearing friction power loss at every 10\u25e6 of crankshaft rotation. The program calculates bearing losses for both \u03c0 (Fig. 10) and 2\u03c0 film extents. Figure 11 shows the bearing friction model flow chart. JET487 \u00a9 IMechE 2009 Proc. IMechE Vol. 223 Part J: J. Engineering Tribology at University of Ulster Library on March 31, 2015pij.sagepub.comDownloaded from The predicted engine bearing friction power loss was calculated using two different approaches, the short bearing analysis and the finite width method. Comparing the predicted results for both the methods (short bearing and finite width) (Figs 12 to 14), it can be seen that there is very little difference in the results for both the cavitation conditions, \u03c0 film as well as 2\u03c0 film (fully flooded)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001347_bfb0109668-Figure2.3-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001347_bfb0109668-Figure2.3-1.png",
        "caption": "Fig. 2.3. The orthographic projection of a ground curve on the image plane. Here sol = 7z and ~2 = 00-~-.",
        "texts": [
            " It can be shown that in the above setting, as long as the tilt angle r > 0, there is a simple diffeomorphism between these two types of projection images (for a detailed proof and an explicit expression for the diffeomorphic transformation see [6]). Consequently, the dynamics of the orthographic projection image curve and tha t of the perspective one are algebraically equivalent. Further on we will use the orthographic projection to study our problem. The orthographic projection image curve of F on the image plane z = 1 is given by (Tx(y , t ) ,y) T E lR 2, denoted by/~, as shown in Figure 2.3. We define: ~i+l ~- OiT~(Y't) e ]It, ~ ---- (~1,~2, . 9 ,~i) T E ]a i, ~ ~-- ( ~ 1 , ~ 2 , . . ) T e ]R ~176 Oyi 9 . Since ~/,(y,t) is an analytic function of y, 7x(y , t ) is completely determined by the vector ~ evaluated at any y. 2 .1 D y n a m i c s o f G e n e r a l C u r v e s While the mobile robot moves, a point a t tached to the spatial frame F I moves in the opposite direction relative to the camera frame Ft. Thus, from (2.2), for points on the ground curve F = (Tx(Y, t), y, 7z(y)) T, we have: (2",
            " Using the homography between the image plane and the ground plane the controllability could be studied on the ground plane alone. However we have chosen to use vision as a image based servoing sensor in the control loop. Studying the ground plane curve dynamics alone does not give the sort of explicit control laws that we will obtain. Our task is to track the given ground curve F. Note that ~ is still a function of y besides t. It needs to be evaluated at a fixed y. According to Figure 2.2 and Figure 2.3, when the mobile robot is perfectly tracking the given curve F, i.e., the wheel keeps touching the curve, the orthographic image curve should satisfy: 7~(y,t)ly=-dcoso =_o (3.1) o~. (~,0 Jy=-dcos ~ -- 0 oy Thus, if ~ is evaluated at y = - d c o s r the task of tracking F becomes the problem of steering both ~1 and ~2 to 0. For this reason, from now on, we always evaluate ~ at y -- - d c o s r unless otherwise stated. T h e o r e m 3.1. ( L i n e a r C u r v a t u r e C u r v e C o n t r o l l a b i l i t y ) Consider the system (2"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000802_1.2908921-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000802_1.2908921-Figure2-1.png",
        "caption": "Fig. 2 Axial cross section of a ball bearing",
        "texts": [
            " It is important to note that parameters rbx and rbz are negative ue to concave surfaces in the ball-ring elliptical contact. In adition, the equations related to the inner and outer rings have to be eparately formulated by using the equations presented above. El- iptic integrals of the first \u0304 and second \u0304 kinds presented in Eq. 6 can be expressed as follows: \u0304 = 1.0003 + 0.5968 Rx Rz , \u0304 = 1.5277 + 0.6023 ln Rz Rx 10 all bearing forces can be calculated from the relative radial dislacements between the rings, which are denoted as ex and ey, orrespondingly. In Fig. 2, a ball bearing with eccentricities in the - and Y-directions is shown. The corresponding radial eccentricty in the direction of ball i can be stated as follows: ei r = ex cos i + ey sin i 11 here i is the attitude angle azimuth angle of ball i. In Eq. 4 , the total elastic deformation can be expressed as ollows: i tot = 2r + h0 in + h0 out \u2212 Rout + Rin + ei r 12 here r is the radius of the ball, h0 in and h0 out are the lubricant film hicknesses between the contact surfaces, and Rout is the outer and in is the inner raceway radius"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000854_s0022112006002709-Figure8-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000854_s0022112006002709-Figure8-1.png",
        "caption": "Figure 8. The region of parameter space in which multiple solutions exist (shaded light) and the stability boundary (solid line) are plotted in the (Dr, H)-plane for a viscosity law with M = 0.1 and K = 12. Special care should be taken when interpreting the stability boundary because the shaded region contains three solutions. The points labelled a\u2013g are included for easy comparison with figure 6(d).",
        "texts": [
            " Thus, one can clearly see that as H increases from 1.3 to 1.7, a small window of instability develops, grows in size, and eventually merges with the draw resonance boundary. The region in which the force decreases with increasing draw ratio is also shown (shaded light). Even though it is not important for the overall stability, we also plot the boundaries where the instability becomes non-oscillatory (dashed line). For easy comparison with figures 2 and 6(d), we also label the points a\u2013g for H = 1.6. In figure 8, we plot the stability boundary (solid line) in the (Dr, H)-plane along with the region in which multiple solutions exist (shaded light). For clarity, in this case, we omit the boundaries where the instability becomes non-oscillatory. One should take particular care when viewing this figure because in the shaded region, there are three possible solutions and, hence, three different manifolds. The two branches of the stability boundary (solid line) are embedded in different parts of the manifold (see figure 2)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002335_detc2010-28926-Figure11-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002335_detc2010-28926-Figure11-1.png",
        "caption": "Figure 11. The resultant of AGs after applying several times the two types of extensions.",
        "texts": [
            "org/about-asme/terms-of-use 6 Copyright \u00a9 2010 by ASME GRAPHS IN 2D In this section we show that it is possible to derive all the AGs in 2D by only two operations. All the AGs, although there is an infinite amount, are arranged in a very unique order as shown in the map appearing in this section. This map is proved to be complete and sound, i.e., all the AGs appear in this map and all the graphs that appear in the map are AGs. The Assur Graphs are arranged in a table with infinite rows and infinite columns; they are all derived from one basic Assur Graph, called dyad, as shown in Figure 11a, by applying only two types of extensions. The first extension, termed fundamental extension, produces all the Assur Graphs in the first row, called also the fundamental Assur Graphs. This operation is done by replacing a ground edge by a triangle and two new ground edges, as shown in 10(a,a1). The fundamental AGs can also be related as representatives, since from each one of them it is possible to derive an infinite number of different AGs. This is done by applying a second extension, termed regular extension, that divides, splits, one of the edges (x,y) by a new vertex, z, and adding a new edge (z,t) for some vertex t/=x,y, as shown in Figure 10(b,b1). Figure 11 depicts example of AGs that are the result of applying a sequence of extensions, starting from the basic AG \u2013 the dyad (Figure 11a). The first row presents the fundamental AGs, called also the representatives, all derived from the basic dyad through applying the fundamental extensions. For example, the fundamental AG in Figure 11b known also as Triad, is obtained by replacing the ground edge (A,O2) with the triangle <A,B,C> with the two new ground edges (C,O2) and (B,O3). All the other infinite fundamental AGs are obtained in the same way; each time a ground edge is replaced with a triangle and two new ground edges. Now, that we can generate all the fundamental AGs, each one of them defines an infinite number of new AGs, all derived by applying successive regular extensions. For example, the AG in (b1) is derived from the fundamental AG, the triad, by applying the regular extension on edge (A,B) by adding vertex D and adding the additional edge \u2013 the ground edge \u2013 (D,O4)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000020_jjap.45.4241-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000020_jjap.45.4241-Figure1-1.png",
        "caption": "Fig. 1. (a) Screen-printed Ag/AgCl electrode, (b) painless needle,",
        "texts": [
            " Secondly, in electronical blood collection, the blood vessel is visualized using an array of near-infrared (NIR) light emitting diodes with a wavelength of 850 nm, the potential change against elapsed time in seconds shown is used to detect the depth of vessel and all these operations can be performed by only focusing on a computer display without actually observing the blood vessel in the arms. Thirdly, new screen-printed carbon and KCl-saturated Ag/AgCl reference electrodes coated on carbon-silver wires have also been developed (Fig. 1). Finally, the screen-printed electrode is then attached to an injection molded polycarbonate plate with a channel pattern (Fig. 1) assembled with a painless needle to complete the healthcare chip. In accordance with the channel design, the operation of the healthcare chip is solely carried out using centrifugal force. As a result, 6 ml of blood is collected using the painless needle and the blood is centrifuged in one direction to obtain plasma and carry out potentiometric measurement later. After measuring, the chip is rotated 90 deg and centrifuged to discard the waste. The cost of the chip design is kept low as it is, only used once and is disposable",
            " As for the two types of anionic additive used, differences between two compounds are TFPB with four fluorine atoms attached to four benzene rings forming borate complexes with sodium ions and K-TCPB with chlorine atoms bonded to four benzene rings in borate complexes with potassium ions. As a result, two anionic additive compounds significantly affect detection. The most important part of this study is the use of TD19C6 to create rigidity of the cyclic compound and introduce blocking efficiency forming a complex with ammonia ions. The electrode used is a screen-printed electrode, which is constructed with a screen-printed Ag layer followed by a screen-printed AgCl layer, Fig. 1(a). The layers of 0.1M ammonium chloride in the presence of 0.5% PVP (polyvinylpyrrolidone) and the ammonia-sensing membrane are deposited using an autodispenser owning to the uniform surface characteristics achieved, which is helpful during the measurement (Fig. 3). As shown in Fig. 1(a), the screenprinted AgCl layer is designed to have an open diameter of 250 nm, which is ideal for use in future miniaturized device purpose. Figure 1(c) shows an injection molded microfluidic system of collecting whole blood using a painless needle [Fig. 1(b)] and of obtaining plasma by applying a centrifugal force. Measurement followed by waste removal can be carried out in one simple chip and most importantly the cost is not high. ISE sensors undergo a series of potentiometric tests to meet the requirements for in vitro clinical use. As reported in the i-STAT system,11,12) 100 mL of blood is required for checking various markers. In our study, a new type of chip has been developed in which only 4 ml of blood is necessary to run the analysis. In the case of ammonia sensors, measurements are carried out firstly in an open-field environment followed by closed-field measurement on the chip"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001265_tmag.2009.2012576-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001265_tmag.2009.2012576-Figure1-1.png",
        "caption": "Fig. 1. Schematic view of solenoidal coil.",
        "texts": [
            " In this paper, the resistance between magnets under various compressive stresses is measured. Then, the effect of stress on the contact resistance between magnets is clarified by solving simultaneous equations. Moreover, the eddy current analysis is performed using the 3-D finite-element method by considering a contact resistance between magnets. The effects of the stress and the exciting frequency on the eddy current loss of such a segmented sintered magnet are also examined. A solenoidal coil which is prepared for measuring the eddy current loss is shown in Fig. 1. The solenoidal coil is wound 320 turns (five layers, 64 turns per layer) using a wire with rectangular cross section (1 mm 3 mm). The B coil with 20 turns wound on the magnet is used to measure the flux density. The Manuscript received October 07, 2008. Current version published February 19, 2009. Corresponding author: N. Takahashi (e-mail: norio@elec. okayama-u.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"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002181_s00707-010-0369-7-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002181_s00707-010-0369-7-Figure1-1.png",
        "caption": "Fig. 1 Plots of scalar spin functions of classical corotational rates of z: a plots of functions r Z (z), r G(z), rd (z), and h(z); b plots of functions gZ (z), gG(z), gd (z), and h\u0304(z)",
        "texts": [
            "3 All the spin tensors considered in Table 2, except for \u03c9n and \u03a9N , belong to the families of continuous spin tensors associated with the families of objective continuous corotational rates considered in Sect. 8. The behavior of the spin tensors listed in Table 2 depends on the parameter z = \u03bbi/\u03bb j . Therefore, it is useful to compare plots of the scalar spin functions defining the behavior of these spin tensors of z. We are especially interested in the behavior of these spin functions for coincident eigenvalues of stretch tensors (i.e., as z tends to 1). Plots of these scalar spin functions of z are given in Fig. 1. Hill [21] (see also [8]) suggested the following equality: E\u0307(0) \u2248 D ( = E(0) D ) for 0.75 < z < 1.25. (129) For the Eulerian counterparts of Lagrangian tensors in (129), the similar equality has the form e(0)G \u2248 d ( = e(0) d ) for 0.75 < z < 1.25. Hill\u2019s approximate equality [21] leads to approximate equalities for the corotational rates of Eulerian tensors h and Lagrangian tensors H: hd \u2248 hG, HD \u2248 H\u0307 for 0.75 < z < 1.25. Hill\u2019s approximate equality implies approximate equalities for the scalar spin functions generating the spin tensors associated with these corotational rates rd(z) \u2248 0 ( = r G(z) ) , gd(z) \u2248 gG(z) for 0.75 < z < 1.25. Hill\u2019s approximate equalities are supported by plots of these functions given in Fig. 1 in the interval 0.8 < z < 1.25 of the argument z.20 Using scalar spin functions of classical corotational rates, one can generate new spin tensors of corotational rates as follows: the functions g(z) of spin tensors associated with classical corotational rates are used as functions r(z) and vice versa, the functions r(z) of spin tensors associated with classical corotational rates are used as functions g(z). The mechanical meaning of corotational rates generated by such spin functions is not clear, but some spin tensors generated in such a way are presented in the literature. As an example, in [57\u201359] a spin tensor of the form \u03c9L \u2261 w + \u03a8 h(V,d)+ \u03a8\u0303 (V,d) = \u03c9n \u2212 \u03c9R + w is considered, which, as the spin tensor \u03c9n , does not belong to the family of continuous spin tensors introduced in Sect. 8. In [53], a skew-symmetric tensor \u03c9A \u2261 R \u00b7 W \u00b7 RT = w \u2212 \u03c9R = \u03a8 r Z (V,d) 20 For the lower bound of the variable z, we use the value of 0.8 instead of 0.75 for the purpose of symmetric arrangement of the bounds of this interval (marked by dashed vertical lines in Fig. 1) on the axis ln z with respect to the value ln 1 = 0 (ln 1.25 = \u2212 ln 0.8 = 0.2). is introduced. This tensor is an Eulerian tensor and is used in the cited paper to define the constitutive relations for a viscous liquid. This tensor does not belong to the family of continuous spin tensors associated with Eulerian corotational tensor rates because the necessary conditions for the objectivity of the Eulerian corotational rates of symmetric tensors (107) for the tensor \u03c9A are not satisfied. A key issue in defining families of continuous spin tensors associated with families of continuous objective corotational tensor rates is the use of the kinematic tensors U, D, and V, d as arguments of the tensor functions \u03a8 r (U,D) and \u03a8 r (V,d)/\u03a8 g(V, d), respectively"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000586_978-1-4020-8889-6_6-Figure4-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000586_978-1-4020-8889-6_6-Figure4-1.png",
        "caption": "Fig. 4 UAV system during flight \u2013 hardware overview",
        "texts": [
            " Forces can only be applied to the UAV mass point and need to be generated by changing the orientation and the adjustment of FMR 3 . The controlled orientation and therefore the generation of forces can be approximated by a third order system. The force generation model and the mass point pendulum model were combined and the resulting model was used for the design of a PI-state-feedback controller. For the multi UAV load transportation it was possible to use the translation controller described in [3, 4] combined with the extended orientation controller described above, which is using the torque compensator. In Fig. 4 one of the UAVs, used for the slung load transportation experiments, is shown during flight. The UAVs are based on commercially available small size helicopters. The helicopters have a rotor diameter of 1.8 m, a main rotor speed of approximate 1,300 RPM and are powered by a 1.8 kW two-stroke engine. The UAVs can carry about 3 kg of additional payload, whereas the weight of the UAV itself is 13 kg. The different components necessary to achieve autonomous flight capabilities are mounted to the helicopters, using a frame composed of strut profiles. Through the use of these profiles, the location of hardware components can be altered and new hardware can be installed easily. This allows quick reconfiguration of the UAVs for different applications, easy replacement of defective hardware and alteration of the position of different components to adjust the UAVs center of gravity. The components necessary for autonomous operation are shown in Fig. 4: A GPS, an IMU, a control computer and a communication link using WLAN. Due to the strong magnetic field of the engine an additional magnetic field sensor is mounted on the tail. All UAVs are equipped with Load Transportation Devices (LTD), which are specially designed for the transportation of slung loads using one or more UAVs. The LTD is composed of a two axis cardan joint with two magnetic encoders attached to each axis. After the joint a force sensor is mounted. After the force sensor a release mechanism for the rope is attached, which is composed of a bolt, inserted into a tube"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000375_02286203.2008.11442485-Figure3-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000375_02286203.2008.11442485-Figure3-1.png",
        "caption": "Figure 3. Proposed mechanical analog of the hysteresis model.",
        "texts": [
            " If we replace the variable x by the torsional torque \u03c4 and the variable y by the angular displacement \u03b8, (1) becomes: d\u03b8(t) dt = h(\u03b8)g[\u03c4(t)\u2212 f(\u03b8(t))] (3) This equation can be rearranged into the form \u03c4(t) = g\u22121 [ \u03b8\u0307(t) h(\u03b8(t)) ] + f(\u03b8(t)) (4) Equation (4) can be interpreted as the mechanical dynamic equation across the flexspline describing the parallel combination of a nonlinear torsional spring and a nonlinear viscous damping. The function f(\u00b7) determines the stiffness curve while the function g\u22121(\u00b7) represents the nonlinear dynamic friction as shown in Fig. 3. The validity of the nonlinear model (4) will be then established. This consists of first showing that the postulated model exhibits the same significant properties as the actual system and then verifying that the model gives realistic responses to one or more test signals. In order to study the dynamic behavior of the complete harmonic drive system, the model of hysteresis will be combined with the wave generator and load dynamic models. The following set of equations represent the complete model of the harmonic drive: J1\u03b8\u03081 +B1\u03b8\u03071 + \u03c4f + \u03c4(\u03b8\u0307, \u03b8) N = \u03c4m (5) J2\u03b8\u03082 +B2\u03b8\u03072 + \u03c4L \u2212 \u03c4(\u03b8\u0307, \u03b8) = 0 (6) \u03c4(\u03b8\u0307, \u03b8) = g\u22121 [ \u03b8\u0307 h(\u03b8) ] + f(\u03b8) (7) \u03b8 = \u03b81 N \u2212 \u03b82 (8) where J1 is the total motor and wave generator inertia, J2 is the total load inertia, \u03b81 is the motor position, \u03b82 is the load position at the end side of the flexspline, N is the reduction gear ratio, B1 and B2 are the viscous damping coefficients at the motor side and the load side, \u03c4 is the transmitted torque across the flexspline, \u03c4m is the driving torque applied by the electric motor, and \u03c4L is the load torque"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000597_978-3-540-73129-0-Figure6.116-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000597_978-3-540-73129-0-Figure6.116-1.png",
        "caption": "Fig. 6.116. Electrostatic valve according to [318]",
        "texts": [
            " A particle filter with a pore size of 10\u00b5m must be used. The operating temperatures are limited to the range 0 . . . 55\u25e6C, and the switching times are typically of the order of one second. In an approach pursued by Bosch, a valve reed in a normally closed design is opened by electrostatic forces (Fig. 6.115). Pressures of the order of 10 kPa and a throughput of 0.2 l/min are controlled with an operating voltage of 200V [317]. In another approach realized by Hitachi, a thin conductive film is electrostatically set into motion over an opening (Fig. 6.116). The valve with 3/2-way functionality can be activated to control a maximum pressure of 30 kPa, and it requires nearly zero power at an operating voltage of 200V [318]. A thermally driven bimetallic valve was developed at the Institute for Microtechnology and Information Engineering (IMIT) of the Hahn-Schickard Society, Villingen-Schwenningen, Germany [307]. The valve is suitable for use with both liquids and gases (Fig. 6.117). The valve consists of two components, a flexible valve reed and a valve seat"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003246_1.4004116-Figure15-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003246_1.4004116-Figure15-1.png",
        "caption": "Fig. 15 Optimized pitch curves of deformed limacon gear pair applied to generate nonsinusoidal wave",
        "texts": [
            " When c \u00bc 2 m=min, the negative slide time can be calculated as tN \u00bc 0:916 s. The mold oscillation cycle is obtained by T \u00bc 2p=x \u00bc 0:4189 s and the corresponding vibration amplitude is 3 mm. All the above parameters satisfy the process requirement of continuous casting, which states that new noncircular gears derived by genetic algorithm optimization can be applied to generate the nonsinusoidal oscillation in continuous casting machine. The optimized pitch curves of the deformed limacon gear pair are shown in Fig. 15. The module of the gears is 6 mm and the tooth number is 30 as shown in Figs. 16 and 17, respectively. When the driving gear rotates counterclockwise and clockwise, the maximal pressure angles are 35:52 and 32:56 , respectively. Fig. 14 Curve of optimized mold velocity Journal of Mechanical Design JUNE 2011, Vol. 133 / 061004-7 Downloaded From: http://mechanicaldesign.asmedigitalcollection.asme.org/pdfaccess.ashx?url=/data/journals/jmdedb/27948/ on 03/26/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use The pitch curves of limacon gears are created by the simple cosine rule"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000597_978-3-540-73129-0-Figure6.125-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000597_978-3-540-73129-0-Figure6.125-1.png",
        "caption": "Fig. 6.125. Electrostatic micromotor fabricated in polysilicon surface micromachining technology [328]",
        "texts": [
            " The piezoelectric drive has a bimorph cantilever movable normal to the wafer surface [327]. Its second direction, the in-plane movement, employs a compliant mechanism in order to enlarge the very small strains of a piezoelectric monomorph. It contains a set of elastic hinges arranged as two-stage gear. Figure 6.124 shows the structure and the kinematic principle of the compliant gear. Electrostatic micromotors built in silicon based surface micromachining technology were first presented by Berkeley University in 1989 [328]. A typical example is shown in Fig. 6.125. The rotor built out of polycrystalline silicone has a diameter of about 100\u00b5m and includes a number of radial teeth. It is surrounded on its periphery by electrodes that can be addressed individually. Their number is larger than the number of teeth (e. g. in a ration of 4:3) so that an attractive force arises between the activated electrodes and nearby rotor teeth, due to an induced electric charging on the electrically insulated rotor. The motors have been brought to spin at rates higher than 10 000min\u22121"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003961_s1068366613040119-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003961_s1068366613040119-Figure1-1.png",
        "caption": "Fig. 1. Calculation model.",
        "texts": [
            " The joint applica tion of the methods aims to extend the possibilities of the method of triboelements via using the findings of finite element analysis of the stress strain state of tribo joint elements as data for determining parameters of the triboelements model of wear. The joint application of these methods made it possible to eliminate some of the limitations in the calculation models. A contact of a rigid shaft of radius with a cylin drical elastic antifriction layer of z thickness in engage ment with a rigid bush (Fig. 1) was considered. The shaft is positioned at angle \u03b1 to the bush. The axis is directed along the bearing axis. The bearing wear and contact pressures depend on the position of contact points. It is taken that only the antifriction layer is worn. General theses of the algorithm for the solution of wear contact problems using the method of triboele ments with ANSYS package are discussed in [6]. In order to solve wear contact problems in a spatial set ting, one should take into consideration that, in this case, a cubic spline of the worn surface is constructed"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002032_s00202-010-0182-2-Figure4-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002032_s00202-010-0182-2-Figure4-1.png",
        "caption": "Fig. 4 Generation of space vector-based ADPWM techniques. a Continual clamping and b split clamping",
        "texts": [
            " Another accepted method clamps every phase during the middle 30\u25e6 for every 90\u25e6 of its fundamental voltage, which is well known by the name DPWM3. DPWM1 uses \u03c3 = 0 for the first half and \u03c3 = 1 in the second half of the sector-I. DPWM3 employs the reverse. So change in the zero state is made at the middle of every sector in the case of the DPWM1 and DPWM3 and is pictorially represented in Fig. 3b and c. If the change in the zero state is made at any spatial angle \u03b1 = \u03b3 , where \u03b3 is between 0\u25e6 and 60\u25e6 as shown in Fig. 4a, each phase is clamped continually for a period of 60\u25e6, whereas if it is as shown in Fig. 4b, each phase clamps for a period of 60\u25e6 but the clamping period splits into two parts one with a width of \u03b3 in the first quarter and other with a width of (60\u25e6 \u2212 \u03b3 ) in the next quarter in every half cycle. These techniques are termed as \u201ccontin- ual clamping\u201d and \u201csplit clamping\u201d techniques, respectively, [17]. Here onwards, DPWM1 and DPWM3 will be referred as continual and split clamping methods with the specified value of \u03b3 . The modulating waveforms of both continual and split clamping techniques for \u03b3 = 45\u25e6 are illustrated in the Figs"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001090_s11370-008-0025-4-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001090_s11370-008-0025-4-Figure2-1.png",
        "caption": "Fig. 2 Robot and sensor modeling",
        "texts": [
            " The unicycle model is assumed as the kinematic model for the robot: xk = g(xk\u22121, uk\u22121, ns,k\u22121) = xk\u22121 + \u23a1 \u23a3 cos \u03c6\u0303k\u22121 0 sin \u03c6\u0303k\u22121 0 0 1 \u23a4 \u23a6 uk\u22121 + ns,k\u22121 (3) where uk\u22121 = (\u03b4sk, \u03b4\u03c6k) is the system input (the vehicle displacement and the vehicle rotation respectively), \u03c6\u0303k\u22121 = \u03c6k\u22121 + \u03b4\u03c6k\u22121/2 is the average robot orientation, and nk\u22121 is a white zero mean noise. The robot is assumed to be equipped with 8 laser rangefinders arrayed on 360\u25e6. Given an environment entirely described by a list M of pairs of points, the related observation model is: z j,k = h(xk,M, ns,k\u22121) = |ar lx j + br l y j + cr | |ar cos \u03b8 j + br sin \u03b8 j | + nb,k\u22121 (4) where (ar , br , cr ) are the coefficients of the r th segment and (lx j , l y j , \u03b8 j ) is the configuration of the laser beam detecting the segment considered. Figure 2 depicts the proposed robot and sensor modeling. In Sect. 3 the independent evolution technique proposed in [10] for the autonomous localization has been recalled. Successively, the collaborative procedure proposed for the multi-robot scenario has been described. In this section, a qualitative comparison is provided in order to highlight the advantages introduced by the collaborative procedure against the independent evolution. 5.1 First environment The first simulation has been carried out in a simple environment with only one wall and two landmarks"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001546_j.jmatprotec.2010.12.002-Figure4-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001546_j.jmatprotec.2010.12.002-Figure4-1.png",
        "caption": "Fig. 4. (a) 3D model of 75 mm spiral welded tube for one complete",
        "texts": [
            " The spiral weld path is defined by n discrete points equally posiioned along the spiral path. The x, y, z points are created to cover nly the 30 mm tube section which represent the ring. To generate he weld path, the weld joint model developed before is revolved dragged) along a spiral weld path. A user subroutine was develped to create these points for spiral weld path and then creating he line. After creating the weld joint area and the spiral line, draging has been performed in order to create the weld volume. The nal model for the 75 mm spiral welded ring is shown in Fig. 4 for ne complete spiral pitch. The geometric parameters of pipe used n this work are given in Table 2. .1. Material model An accurate modeling of temperature dependant material roperties is a key parameter to the accuracy of computational echanics and has been a challenging job due to scarcity of mateial data at elevated temperature. A hypoelastic rate-independent able 3 emperature-dependent thermal conductivity of steel (Mahapatra et al., 2006). Temperature (K) 273 373 573 723 Thermal conductivities (W/m K) 51"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001512_s11668-010-9398-8-Figure5-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001512_s11668-010-9398-8-Figure5-1.png",
        "caption": "Fig. 5 Second kind of gear tooth configuration: (a) geometrical dimension, (b) solid model and (c) cooling model",
        "texts": [
            " 4 were drilled at an angle oriented along the gear axis. Fresh, cooling air flowed from one surface to the other as the gear rotated at a constant speed. The size of the holes that were drilled in the vicinity of the pitch diameter affected the quantity of air mass and also reduced the mechanical properties of the gear. However, the holes were carefully positioned to minimize the effect of the holes on mechanical behavior. Second Type of Configuration for the Reduction of the Plastic Gear Tooth Surface Temperature Figure 5 shows the geometry of the gears with two different cooling holes. The second type of gear has one hole ([ 5 mm) located at a 76-mm radial position, and an additional two cooling holes were drilled along the radial direction of the tooth body ([ 2.5 mm), as shown in Fig. 5. Thus, it was assumed that the heat that was generated due to the cooling hole formed on the gear tooth body was removed by fresh air circulation. According to Bernoulli\u2019s law, a moving fluid generates a lower pressure than a stationary fluid. Therefore, as the gear rotates, it will rotate the surrounding air molecules in accordance with the \u2018\u2018no-slip\u2019\u2019 condition. If the gear rotates at a constant speed (W), then the points on the gear move at different speeds according to their radial positions. The holes (Fig. 5) have different pressures, and the speeds have the relationship of (Va \\ Vb) and (Pb \\ Pa) that depend on inertial forces on the air and the pressure difference. Mass transfer occurred between the holes b and a because of the effects of pressure difference and centrifugal forces. Therefore, the flowing air removes heat via convection. The cooling holes also reduce the strength and stiffness of the gear, but evaluation of the reductions in mechanical behavior is beyond the scope of this study. Unfortunately, the holes which are drilled to reduce thermal loading on the gear and thereby improve gear performance also reduce the mechanical response of the gear, and this may result in reduced performance",
            "1 N/mm advance because of the effects of heat in reducing the properties of the plastic materials if the gear tooth stress had scaled up with the size of the hole and the gear tooth stress. Second Type of Gear Tooth Configuration The experiments with the first configuration demonstrated that the cooling holes did not significantly improve the thermal damage characteristics. Therefore, additional holes were created on the body of the tooth to achieve a second type of gear configuration (Fig. 11). If the gear rotates at a constant speed (W), then the points on the gear move at different speeds according to their radial position. Therefore, the holes (Fig. 5) have different pressures, and the speeds have the relationships (Va \\ Vb) and (Pa \\ Pb). Owing to inertial forces on the air and the pressure difference, mass transfer will occur between holes b and a. Therefore, the transferred air will expel the accumulated heat via convection. Owing to the additional holes, the heat transfer from the body of the tooth increased, which resulted in a decrease in the tooth body temperature. Therefore, the lowest tooth temperatures were obtained using the second type of gear configuration"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001247_0954406jmes321-Figure17-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001247_0954406jmes321-Figure17-1.png",
        "caption": "Fig. 17 Ease-off topography of the face gear for example 2",
        "texts": [
            "68E-3) Parameter of about contact line angle \u03b6 (\u25e6) (left, right) (9.0, \u221232.9) Parameter of face gear about bias angle \u00b5 (\u25e6) (left, right) (\u221260.0, 60.0) JMES321 \u00a9 IMechE 2007 Proc. IMechE Vol. 221 Part C: J. Mechanical Engineering Science at UNIV OF CONNECTICUT on April 13, 2015pic.sagepub.comDownloaded from the face-gear set to achieve adequate contact patterns through an increase in the shaping cutter tooth number. However, the contact path and size of the contact pattern could be determined directly with the proposed profile modification methodology. As Fig. 17 shows the amount of ease-off along the contact path is the same for the left and right flanks, but the bias angles are different for the purpose of improving the contact ratio. According to the TCA results shown in Figs 18 and 19, the face-gear set permits large axes misalignment without any edge contact. The current paper first proposed a double-crowning methodology for a moulded face gear and then investigated the consequences of varying the design parameters. Based on the numerical examples above, the following conclusions can be drawn"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002110_rnc.1483-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002110_rnc.1483-Figure1-1.png",
        "caption": "Figure 1. The principal forces and moments of the robotic helicopter with respect to its body frame.",
        "texts": [
            " The robustness of the proposed controller is investigated in Section 4 for the mass variation during the airdrop mission. In Section 5, several numerical simulations are conducted to demonstrate the merits of the proposed backstepping control method. Section 6 concludes the paper. This section is aimed to briefly review the complete nonlinear dynamic model of a robotic helicopter with single main rotor and tail rotor. Like the model in [16], a robotic helicopter model is considered as a rigid body incorporating with a simple component force model for the rotor dynamics and illustrated in Figure 1. For detailed mathematical modeling for the full-scale and small-scale helicopters, the reader is referred to [1\u20133]. Therefore, the complete dynamic model of the robotic helicopter incorporating with the force and moment equations is summarized by [10, 16] p\u0307i = i (1) \u0307i =ge\u0302i3+ 1 m ( )ubT + 1 m ( )KubM (2) \u0307( )= ( )sk( b) (3) Im\u0307b=\u2212 b\u00d7 Im b\u2212Qtre\u0302 b 2\u2212Qmre\u0302 b 3+PubM (4) where m\u2208 R and Im \u2208 R3\u00d73, respectively, denote the total mass and the moment inertial matrix of the robotic helicopter. As illustrated in Figure 1, the position and velocity of the robotic helicopter center of gravity are given by pi =[x y z]T and i =[vx vy vz]T with respect to the inertial frame in North-East-Down (NED) orientation (with an upper sub-index i). The helicopter angular rate vector b=[ x y z]T and the Euler angle vector b=[ ]T defined in the roll-pitch-yaw sequence are with respect to its body frame (with an upper sub-index b). Furthermore, ( ) is the helicopter\u2019s rotation matrix from the body axes to the inertial axes and sk( b) means the skew-symmetric matrix of the body angular rate"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001931_iros.2010.5649323-Figure13-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001931_iros.2010.5649323-Figure13-1.png",
        "caption": "Fig. 13. Arrangement of areas in bj(z)",
        "texts": [
            " If adjusting the relation between a sink m and a velocity uj , the chasing motion of the moving obstacle can be avoided as shown in Figs. 5(a)\u2013(d). However, the velocity of a moving obstacle is not continuous and is irregular in real environment. If m changes frequently up and down according to the changing uj , the velocity of a mobile robot changes too frequently though the motion of chasing the moving obstacle does not occur. Therefore, it is necessary for the method of using no changing the value of m to avoid the chasing motion. That is, bj(z) is a necessity function. Fig. 13(a) shows the arrangement of areas of 1 and 0 by using the previous bj(z). These two values of 1 and 0 are changed on the point where the angle between the direction of the mobile robot and the direction of the obstacle is \u03c0 2 . So, Dj(z) is neglected immediately when the moving obstacle has passed. To solve this problem, the following new bj(z) is redefined instead of the previous bj(z) as shown in Eq. (4). bj(z) =  1, for \u03b1j \u2212 \u03c0 2 \u2264 6 (z \u2212 zj) \u2264 \u03b1j + \u03c0 2 1 + cos 2\u03b2j 2 , for \u03b1j + \u03c0 2 < 6 (z \u2212 zj) < \u03b1j + \u03b3 or \u03b1j \u2212 \u03b3 < 6 (z \u2212 zj) < \u03b1j \u2212 \u03c0 2 0, for other (21) where \u03b1j denotes the direction of the moving obstacle, \u03b3 denotes the terminal angle for using cosine function, \u03b2j is represented by the following. \u03b2j = 6 (z \u2212 zj) \u2212 ( \u03b1j + \u03c0 2 ) (22) A cosine function is installed at the discontinuous point, so bj(z) is changed smoothly from 1 to 0 in Eq. (21). Fig. 13(a) shows the arrangement of areas of 1 and 0 by using the previous bj(z). Fig. 13(b) shows the arrangement of areas of 1, (1 + cos 2\u03b2j)/2 and 0 by using the new bj(z). Figs. 14(a)\u2013(d) show the flow fields by using the new correction function bj(z), where Figs. 14(a),(b) use \u03b3 = \u03c0 and Figs. 14(c),(d) use \u03b3 = 3\u03c0 4 . The overshooting paths, that is chasing curve to the moving obstacle, by using \u03b3 = 3\u03c0 4 are smaller than the overshooting paths by using \u03b3 = \u03c0. Therefore, we decide to use \u03b3 = 3\u03c0 4 in the simulation of the next chapter. Compare Figs. 14(a),(c) and Fig. 4 by using the circle"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000702_bf02441577-Figure7-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000702_bf02441577-Figure7-1.png",
        "caption": "Fig. 7 Finite-element discretisation of the left ventricle",
        "texts": [
            " Stress analysis The left ventricular chamber shape is obtained from a single-plane cine angiogram taken at midejection. The chamber dimensions and wall thickness are taken from G m s ~ and SANDL~R (1970). The axis of revolution of the shell is assumed to be the line joining the left ventricular apex and the centre of the aortic valve opening. As the left and right side of the chamber is not symmetrical, both are revolved around the axis in turn and the stress distribution is determined. The finite-element discretisation of the symmetrical shell for both the cases is shown in Fig. 7. In this analysis, the deformation of the aortic valve and mitral rings, due to internal pressure at the instant of taking the angiogram, are neglected. The same isotropic material properties of the myocardium, used in their analysis by HAmD and GntSWA (1974), have been used here. Layered left ventrieular wall It is known (RusHM~R, 1961) that the left ventri- cular wall consists of three layers, that the muscle fibres are wound round the chamber at different angles and that the major portion of the muscle mass is concentrated in the middle layer"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002930_978-3-642-31988-4_26-Figure7-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002930_978-3-642-31988-4_26-Figure7-1.png",
        "caption": "Fig. 7 Required workspace of the six-cable driven parallel manipulator",
        "texts": [
            " \u03b5ll = f (L , H, h) = ( a1 \u00d7 L2 \u2212 a2 \u00d7 L \u2212 b1 \u00d7 h2 \u2212 b2 \u00d7 h ) \u00d7 10\u22126 + ( \u2212c1 \u00d7 10\u22124 \u00d7 H2 + c2 \u00d7 10\u22122 \u00d7 H \u2212 d ) \u00d7 10\u22126 (34) The compensation polynomial coefficients for the line equation can be optimized by genetic algorithm in Fig. 6. From Fig. 6, we obtain the expected error result of this modeling is about RMS 0.78841mm. Therefore, the compensation polynomial for the line equation is: Fig. 8 Experimental trajectories-line 3B 2B 5B 4B 6B 1B Fig. 9 Experimental trajectories-arc 3B 2B 5B 4B 6B 1B \u03b5ll = ( 8.3721 \u00d7 L2 \u2212 3.5328 \u00d7 L \u2212 9.8747 \u00d7 h2 \u2212 7.1498 \u00d7 h ) \u00d7 10\u22126 + ( \u22125.0527 \u00d7 10\u22124 \u00d7 H2 + 0.0894 \u00d7 10\u22122 \u00d7 H \u2212 3.2483 ) \u00d7 10\u22126 (35) In Fig. 7, required workspace of the six-cable driven parallel manipulator is presented as a sphere crown surface for studying the modeling error and error compensation methods. For proving the feasibility of the error compensation method of the six-cable driven parallel manipulator, three experimental trajectories are introduced in Figs. 8, 9, 10. The three experimental trajectories are line, arc and circle respectively. In Fig. 8, the line trajectory is from G1 = (0, 0, 8.4 m) to G2 = (0, 0, 9.4 m), and the kinematic velocity is v = 2"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002425_vppc.2012.6422784-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002425_vppc.2012.6422784-Figure2-1.png",
        "caption": "Fig. 2 (a) shows a cross-sectional figure and a simplified model of 2-pole PMSM with the concentric winding. In the figure, the a-phase winding is divided into two parts, i.e., healthy winding and fault winding. Fig. 2 (b) shows the magnetic circuit for inductance calculation of 2, 4, ... pole PMSM, where (A,2,PJ2, N, P, and R denote the magnetic flux of the slots, the turn number of one slot, the pole number, and the magnetic reluctance, respectively. x and \ufffd denote the rate of healthy turn rate in one slot and the rate of the fault turn of one phase. x and \ufffd satisfy",
        "texts": [
            " In this paper, dynamic d-q model of the PMSM under the inter turn short fault is derived in the positive and the negative sequence rotating frame. Based on the model, a new fault detection scheme is proposed also. The proposed 978-1-4673-0954-7/12/$31.00 \u00a920121EEE The turn number of main circuit a-phase is reduced by the fault. Hence, the motor model changes depending on the faulted turn number. To establish the main circuit model, the equivalent magnetic circuit is used. where Nshort denotes the fault turn number. By the magnetic circuit of Fig. 2(b), the flux density at the faulted phase is derived as (2) Nl(%P-Ij-Nl( f -2 j - XNl IA,3,Pl2 = R(%P-lj +R (3) By eq. (2) and (3), the self magnetizing inductance of the turn short phase Lmt is obtained such that (4) If the PMSM has no fault (\ufffd= 0), the self magnetizing inductance is Lm = PN2 I 3R. But, if the PMSM has an inter turn short (1 Z \ufffd > 0), the decrease of the self magnetizing inductance has a quadratic function. Fig. 3 shows the decrease of the magnetizing inductance along the fault ratio"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003847_ecc.2013.6669739-Figure5-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003847_ecc.2013.6669739-Figure5-1.png",
        "caption": "Fig. 5. Horizontal flight trajectory of the quadrocopter over two rounds of the circular trajectory. The dashed blue line shows the flight path before learning, the solid red line shows flight after ten iterations of learning. The nominal trajectory is shown in dotted black. Errors in height are not shown; they were reduced from a maximum error of 0.42 m to a maximum of 0.01 m.",
        "texts": [
            " Figure 4 shows the Euclidian norm of the error Fourier series coefficients over ten iterations. It can be seen that initial tracking performance is relatively poor with peak tracking errors of 64 cm. The errors are then quickly reduced over the first three iterations, after which non-repeatable disturbances cause them to vary from iteration to iteration while small improvements are made. After ten iterations of the learning algorithm, the peak tracking error is reduced to 11 cm, and subsequent iterations do not improve tracking performance significantly. Figure 5 shows the flight path of the vehicle in the horizontal plane before learning and after the tenth learning iteration. Note that the initial flight path shows large errors, with the flown circle being much too large, shifted from the desired centre point, and warped. After ten learning iterations, the tracking performance has improved considerably, although remaining, largely unrepeatable, disturbances still prevent the vehicle from following the trajectory perfectly. This paper evaluated an iterative adaptation scheme that improves tracking performance when periodic disturbances cause poor tracking under feedback control"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001884_s11431-010-3100-y-Figure3-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001884_s11431-010-3100-y-Figure3-1.png",
        "caption": "Figure 3 The non-coincident links between every two adjacent loops.",
        "texts": [
            " We then propose a general formula expressed by the virtual-loop rank to look for a general mobility formula. To simplify analysis, a link group is regarded as the combination of non-coincident links between any independent loop and its adjacent loop. The mobility of a link group can be equal to 0, or less than 0, or more than 0. The links of the link group can be driven or motive. The link group defined here is a generalized group, not the Assur group. For the planar 9-bar linkage shown in Figure 2, it contains four independent loops, and has four link groups, ABCD, EFG, HK and PRM, as shown in Figure 3. To describe the motion transmission manner of two adjacent loops by a terminology, we define the concept of virtual kinematic pair. Assume that links A and B are responsible for the motion transmission between two adjacent loops. No matter whether they are adjacent, it is supposed that they are connected by a kinematic pair, which is regarded as a virtual kinematic pair, short for virtual pair. The virtual pair is a generalized pair. When A and B are adjacent, the virtual pair is a real pair; when they are not adjacent, the virtual pair is a generalized pair, i"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001370_icsens.2009.5398292-Figure5-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001370_icsens.2009.5398292-Figure5-1.png",
        "caption": "Fig. 5. Flexible sensor installed in a cylinder",
        "texts": [
            " In initial state, The length of MPA is 150[mm] and the diameter is 15[mm]. The ratio of circumference direction is 2 times larger than the ratio of axial direction. Measuring with the flexible sensor we developed, the S/N ratio is 3.1[dB] in axial direction and 14.4[dB] in circumference direction. In circumference direction, the S/N ratio is also larger than in the axial direction and higher accuracy could be obtained by measuring the circumference displacement than directly measuring the axial displacement. Figure 5 shows the model of flexible sensor The flexible sensor was winded around a cylinder-approximated MPA (Diameter is D). A and B are the contact points for measuring the resistance. Between A and B, sensor is divided as two parallel resistors with same resistance, RAPB = RAQB . The total resistance RAB between A and B can be calculated as RAB = RAPB \u00b7 RAQB RAPB + RAQB = 1 2 \u03c1 \u03c0D/2 Sc . (5) Where D is the diameter of MPA, \u03c0D/2 is the length of each half sensor, and Sc is the cross-section area of the electroconductive rubber"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001301_s00466-009-0394-3-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001301_s00466-009-0394-3-Figure2-1.png",
        "caption": "Fig. 2 Linear segment on the lubricated boundary \u2202l , here sketched for m p = 3 so that there are m p + 1 = 4 pressure nodes per element",
        "texts": [
            " The important consequence of using low-order solid elements is that the boundary, including the lubricated boundary \u2202l , is discretized into linear segments with piecewise-linear approximation of displacements. This implies linear interpolation of the film thickness h within each segment (first-order interpolation, mh = 1) as the rod surface is rigid and, due to axial symmetry, represented by a line. However, arbitrary interpolation order m p can be adopted for the pressure p. Accordingly, the interpolation within a typical segment Sel is, cf. Fig. 2, x = 2\u2211 i=1 N (1) i xel i = 2\u2211 i=1 N (1) i (Xel i + uel i ), (26) and p = m p+1\u2211 i=1 N (m p) i pel i , (27) where xel i , Xel i , uel i , and pel i are the nodal quantities, and N (k) i are polynomial shape functions of order k. Clearly, in view of Eq. (26) we have x\u0304 = 2\u2211 i=1 N (1) i x\u0304el i , h = 2\u2211 i=1 N (1) i hel i , (28) where the nodal quantities x\u0304el i and hel i follow from straightforward considerations. Importantly, increasing the pressure interpolation order results only in a small overhead on the total number of unknowns and on the overall computational cost"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000400_1.2709514-Figure5-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000400_1.2709514-Figure5-1.png",
        "caption": "FIG. 5. Color online a Schematic diagram of the experimental setup and b optical images of droplet mixing using discontinuous ferrofluid pattern. Droplet size is 8 l.",
        "texts": [
            " Next, we investigated the mixing of water droplets after their transportation. We found that ferrofluid covers the surface of water droplets during the process and influences the mixing dynamics increasing the mixing time . The creation of a discontinuity in the ferrofluid pattern is expected to reduce or avoid the presence of the ferrofluid at the interface, thus improving the mixing efficiency. In order to create the discontinuity in the ferrofluid pattern at the mixing spot, we placed a smaller strip magnet at that location. Figure 5 a shows the setup and Fig. 5 b shows the mixing of two water droplets using this procedure. The mixing time of two 8 l water droplets with/without the discontinuity in the ferrofluid pattern was measured at different rotation speeds 200, 300, and 500 rpm , see Table I. As can be seen, the mixing time does not depend on the rotation speed of the magnetic stirrer, but it is drastically reduced by using the discontinuity. This clearly shows that the presence of discontinuity pushes away the ferrofluid from the interface of two water droplets giving them a greater chance to touch and mix rapidly"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001884_s11431-010-3100-y-Figure7-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001884_s11431-010-3100-y-Figure7-1.png",
        "caption": "Figure 7 JH//PT, JH=PT.",
        "texts": [
            " It is proved that dX is the rank of kinematic constraint equations. The actions of these two definitions are identical. Example 4. Figure 6 is a 2-PRU/PR(Pa)R mechanism formed by three loops. Loop I is the loop ABCDFE and lies on plane O-yz. It consists of a link group with 2-PRU (C, D are U-pair). Loop II is the GMJHKDEF and contains a link ~ ~ group with PRR//RR. Loop III is the loop HJPT and includes a link group with 2R. Since JH//PT and JH = PT, Loop III is the parallelogram mechanism with two cranks (Figure 7). In loop I, all links cannot rotate about x-axis nor translate about z-axis. Its rank is dI = dX I (, 0 y z)=4, FI = PI i=1fi dX I =84=4. Since link 3 cannot rotate about x-axis nor translate about z-axis, the rank of generalized pair G3,11 I is d3,11 I ( , 0 y z)=4. In loop II, links 3 and 6 form a revolute joint K, links 9 and 10 form a revolute joint M. The two axes are perpendicular to plane O-xz. The rank of the link group with PRR//RR is dgz II (, x y z)=6, so dX II=d3,11 I ( , 0 y z)+d gz II (, x y z)=d X II (x y z)=6, FII= PII i=1 fi d X II =56=1"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002469_s11804-012-1144-z-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002469_s11804-012-1144-z-Figure1-1.png",
        "caption": "Fig. 1 Sketch map of platform motions",
        "texts": [
            " Recently, much attention has been paid to the MPC method for the control systems of ships, and other ocean structures (Wang, 2006; Song, 2005). In this paper, the model predictive control approach is proposed to solve the control problem of DPS based on a linear motion model of the platform with many kinds of system constraints taken into account. A simulation is provided to illustrate the effectiveness and the stability of the controller. Generally speaking, the motion of ocean structures often has 6 degrees of freedom (DOF) as shown in Fig. 1; they are surge, sway, heave, roll, pitch, and yaw. The reference coordinates of the platform are shown in Fig. 1. The motion model of the platform can be divided into two parts: the high-frequency motion model and low-frequency motion model. The high frequency motion model is caused by a first-order wave which can only lead to oscillation of the platform without shift. Generally speaking, the propulsion system of DPS does not need high-frequency control because it would greatly increase energy consumption and accelerate Hongli Chen, et al. Model Predictive Controller Design for the Dynamic Positioning System of a Semi-submersible Platform 362 wear of the thruster"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001065_13506501jet487-Figure8-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001065_13506501jet487-Figure8-1.png",
        "caption": "Fig. 8 Journal bearing geometry and co-ordinate system",
        "texts": [
            " It should be noted that the short bearing mobility method is subjected to all the assumptions applied to its own derivations (neither elastic nor thermal distortion is considered for either the shaft or the bush and the shaft is well aligned with the bearing). Booker [33, 34] developed curve fits based on the short journal bearing approximation to calculate the journal centre velocity components. In a similar manner, Goenka [36] presented some curve fits generated by applying the finite width bearing approximation, which takes into consideration the lubricant pressure flow in both axial and circumferential directions. Details can be found in Dickenson [37]. Martin [38] presented the bearing friction power loss equations based on Fig. 8. These power loss equations predict instantaneous losses at a single crank angle condition. Considering only the shear effects, the power loss can be determined as H = \u03b7r3b c J 00 1 (wj \u2212 wb) 2 (10) where \u2018J 00 1 \u2019 is a journal bearing integral and can be found in Booker\u2019s table of journal bearing integrals [33]. For the shaft staying concentric to the bush, the power loss is given as H = 2\u03c0 \u03b7r3b c (wj \u2212 wb) 2 (11) The above equation is known as the Petroff equation. Proc. IMechE Vol. 223 Part J: J"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001345_robot.2008.4543604-Figure7-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001345_robot.2008.4543604-Figure7-1.png",
        "caption": "Figure 7. The grasping force",
        "texts": [
            " The closed-loop control system of the position (6), (7), (12) is stable if the fluid pressures control law in the chambers of the elements given by: tektekatu i j ii j ijiji 21 (29) tektekbtu qi j qiqi j qijiqji 21 , (30) where ; , with initial conditions: 2,1j Ni ,,2,1 000 211121 iiiii ekkpp (31) 000 211121 qiqiqiqiqi ekkpp (32) 00ie (33) 00qie , (34)Ni ,,2,1 and the coefficients , , , are positive and verify the conditions ik qik mn ik mn qik 2111 ii kk ; (35)2212 ii kk 2111 qiqi kk ; , (36)2212 qiqi kk Ni ,,2,1 Proof. See Appendix 1. B. Force Control The grasping by coiling of the continuum terminal elements offers a very good solution in the fore of uncertainty on the geometry of the contact surface. The contact between an element and the load is presented in Figure 7. It is assumed that the grasping is determined by the chambers in -plane. The relation between the fluid pressure and the grasping forces can be inferred for a steady state from [2], 8 ~~~~ 21 0 00 2 2 dSpp dssTsTsfds s sk l s T l (37) where sf is the orthogonal force on the curve ,bC sf is sF in -plane and sFq in q-plane, respectively. A spatial discretization is introduced and121 ,,, lsss ii ss 1 , ii s , 1,,2,1 li (38) For small variation i around the desired position id , in -plane, the dynamic model (6) can be approximated by the following discrete model [11], eiiidid didiidiiiii FfdqH qHcm , ,, , (39) where Smi , 1,,2,1 li , did qH , is a nonlinear function defined on the desired position did q, , diii qcc ,, , ,0ic q, (40) is the viscosity of the fluid in the chambers"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002335_detc2010-28926-Figure8-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002335_detc2010-28926-Figure8-1.png",
        "caption": "Figure 8. Determinate Truss and its corresponding decomposition graph.",
        "texts": [
            " Search for an Assur Graph that can be removed and at least one force, which is acting on one of its vertices. 2. Remove this AG, add ground vertices to its outer vertices with ground vertices and calculate the forces in its internal edges. 3. Replace all the ground edges of the removed AG with external forces with the same magnitude and direction forces in its ground edges. 4. Go to 1. An example of applying this analysis process appears in Figures 8 and 9. The determinate truss for which the analysis is applied consists of three triads and one dyad, as appears in Figure 8a. Figure 9 depicts the process of analyzing the determinate truss appearing in Figure 8a, each time an AG is being analyzed. First, the triad (A,B,C) can be removed and since on one of its vertices, vertex B, acts an external force thus this AG is the first to be removed and analyzed (Figure 9b). The inner forces in the three ground edges, (AK), (CD) and (B,K), of the latter AG become external forces that act on the remaining determinate truss: PBK, PCD, PAD, as shown in Figure 9c. This process continues and is applied on the dyad K (Figure 9c), then triad (G,H,I) as shown in Figure 9d, and ends with the analysis of the triad (E,D,F) upon which four external forces act (Figure 9e)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002390_jmems.2012.2194777-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002390_jmems.2012.2194777-Figure1-1.png",
        "caption": "Fig. 1. Device structure of an ID and its operation mechanism.",
        "texts": [
            " The contact of liquid and liquid can generate a smooth interface and minimize the effect of gravity on distorting the shape of the liquids. By changing the aperture of the ID, the shape of the liquid surface is changed. Therefore, the focal length of the lens is tunable. Adaptive lenses based on our cell structure own 1057-7157/$31.00 \u00a9 2012 IEEE the advantages of compact structure, easy operation, scalable aperture, multistability, and good performance. In our lens cell, the key part is the ID which consists of the base plate, blades, and rotatable disc, as shown in Fig. 1(a)\u2013(c), respectively. Some holes are drilled near the inner border of the base plate. These holes are used to fix the blades. Each blade in Fig. 1(b) has two pins. One pin (pin-1) points to the inside, and the other one (pin-2) points to the outside. The rotatable disc has some slides and a handle. Fig. 1(d) shows the arrangement of one blade positioned on the base plate. By pushing the pin-2 in clockwise direction, the blade will rotate in clockwise direction, as shown in Fig. 1(e). Fig. 1(f) shows the arrangement of five blades positioned on the base plate. For the five blades, their pins pointed to the outside are positioned in the slides of the rotatable disc. By rotating the handle of the rotatable disc, the aperture of the base plate can be changed. To demonstrate a lens, a commercial ID is chosen, as shown in Fig. 2. It exhibits the same operating mechanism of the device as shown in Fig. 1. Fig. 2(a) shows the ID with a larger aperture, and Fig. 2(b) shows that with a smaller aperture. Its aperture can be tuned from 15 to 0.8 mm. The outer diameter of the ID is 22 mm. The ID has ten blades, and each blade is very thin. The surface of overlapped blades is rather smooth. The total thickness of the ID is 5 mm. The cross-sectional structure of the lens cell and its operation mechanism are shown in Fig. 3. From the bottom to the top, it consists of the glass plate, liquid-1 (L-1), liquid-2 (L-2), and the top glass plate"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003001_00405001003696464-Figure3-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003001_00405001003696464-Figure3-1.png",
        "caption": "Figure 3. Scheme for calculation of the parameters of the internal structure of a plain weave fabric.",
        "texts": [
            " Let d be the wire diameter, p 0 be the distance between the central lines of the warp and weft wires, and h 0 be the crimp heights of the warp and weft wires (subscript 0 denotes fabric parameters before the deformation). Assuming that (a) the contact points of the wires are not moved during the tensile deformation, and (b) the shape of the central line of the wires z ( x ) between the intersections can be approximated with a polynomial of the third order, the following geometrical relations hold (Figure 3): where l ( p , h ) is the length of the central line of the wire between two intersections (Equation (2)). Figure 3. Scheme for calculation of the parameters of the internal structure of a plain weave fabric. Equation (2) allows calculating the curvature of the wires, the average deformation of the wires, and the angle of inclination of the centre lines of the wires Equations (2)\u2013(7) have two unknowns \u2013 crimp heights h 1 and h 2 \u2013 related with the constraint Equation (3). They are calculated from the condition of minimum energy of the whole structure (de Jong & Postle, 1978; Hearle & Shanahan, 1978): where \u201cbend\u201d and \u201ctens\u201d refer to the energy of bending and tensile deformations, respectively"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000041_j.mechmachtheory.2006.05.006-Figure16-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000041_j.mechmachtheory.2006.05.006-Figure16-1.png",
        "caption": "Fig. 16. Schematic diagram and geometrical parameters of the limbs: (a) Stewart platform, (b) Hexaglide.",
        "texts": [
            " The solution to the second problem is useful in a practical sense, as it would not only maintain a maximum constant cutting force for the entire machining process but would also eliminate possible discontinuities and sudden jumps of control forces to avoid structural damage. The numerical algorithms developed for solving these problems are based on the method of approximate programming, and taking into account the non-linear bounds imposed on the control forces and the rotation angle of the moving platform. The effectiveness of these algorithms is demonstrated by two numerical examples. The financial support of this work by the National Science Council of the Republic of China under Grant 94-2212-E-011-007 is gratefully acknowledged. As shown in Fig. 16(a) and (b), each limb of the Stewart platform and the Hexaglide consist of two links. For the Stewart platform, the virtual displacement of the mass center of two links can be, respectively, written as dP1i \u00bc dWi q1ili \u00f0A:1\u00de and dP2i \u00bc dPi dWi q2ili; \u00f0A:2\u00de where li is a unit vector along the direction of the limb and dWi is the virtual rotation of the limb. dPi is the virtual displacement of the center point of the spherical joint pi attached to the moving platform, and which is related to the virtual rotation and virtual displacement of the center of mass of the moving platform by dPi \u00bc dP\u00fe dW ri: \u00f0A:3\u00de On the other hand, since the absolute position of joint pi can also be expressed as Pi \u00bc Bi \u00fe qili; \u00f0A:4\u00de in which Bi is a constant vector representing the absolute position of the universal joint bi attached to the fixed link, therefore the virtual displacement of Pi can also be written as dPi \u00bc dqili \u00fe dWi qili: \u00f0A:5\u00de Taking the cross-product of both sides of Eq",
            "5) with li and noting that dWi should be perpendicular to li in order to be kinematically admissible, yields dWi \u00bc li dPi=qi: \u00f0A:6\u00de Substituting Eqs. (A.3) and (A.6) into Eqs. (A.1) and (A.2), the relationship between the virtual displacement and rotation of links and that of the moving platform can be combined together and written into matrix form as dbXi \u00bc \u00bdJi dbX; \u00f0A:7\u00de where dbX i \u00bc \u00bddPT 1idWT i T for the first link \u00bddPT 2idWT i T for the second link ( and the transformation Jacobian matrix is given by : Similarly, according to the geometric property of the vectors shown in Fig. 16(b), the relationship between the virtual displacement and rotation of the two links of each limb of the Hexaglide and that of the moving platform can also be derived into the general form of Eq. (A.7), and the corresponding transformation Jacobian matrices of the links in each limb are given by : The purpose of this appendix is to present a simple method for constructing the initial feasible trajectory of the rotation angle / for use in the computational algorithms. By denoting the initial trajectory as / = /(f), where f 2 [0, 1] is a normalized parameter of the execution time, the boundary conditions imposed on the two end points of the trajectory can be written as /\u00f00\u00de \u00bc _/\u00f00\u00de \u00bc 0; \u00f0B:1\u00de and _/\u00f01\u00de \u00bc \u20ac/\u00f01\u00de \u00bc 0: \u00f0B:2\u00de To ensure the feasibility of the trajectory, the maximum value of / can not exceed the minimum bound /s of the rotation angle as evaluated along the given machining contour, or /\u00f0n\u00de 6 /s 8n 2 \u00bd0; 1 : \u00f0B:3\u00de These five conditions can be satisfied with a fourth-order polynomial as /\u00f0f\u00de \u00bc X4 k\u00bc0 akf k; \u00f0B:4\u00de where ak (k = 0\u20134) are constant coefficients to be determined"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000193_icma.2005.1626722-Figure3-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000193_icma.2005.1626722-Figure3-1.png",
        "caption": "Fig. 3 Schematic view of Trinocular optical CCD-based detector",
        "texts": [
            " As a result, the location of the vision system with respect to the substrate holder and the component placed on it should be carefully selected to prevent any possible collision between the vision system and the substrate. 3) The computational time for the image processing is typically high. In order to use the vision system in a real-time control system, it is crucial to develop a simple but realistic algorithm to reduce the computational time as much as possible. In order to address these issues, a trinocular optical CCDbased detector is designed as shown in Figure 3. Three optical CCD-based detectors are considered to monitor the process zone in which appropriate neutral and interference filters are integrated [6]. The polar angles between the optical CCD- based detectors are 120 and the horizontal angle between an optical CCD-based detector and the substrate plane is set to 15 . Three digital cameras with resolution 1000 \u00d7 1000 are connected to digital frame grabbers within a PC equipped with a QNX real-time platform. In the frame grabber, the images are pre-processed and chopped to size of 350\u00d7350, in which each pixel is associated with dimension of 0"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001375_idaacs.2009.5342900-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001375_idaacs.2009.5342900-Figure1-1.png",
        "caption": "Fig. 1. The remotely controlled mobile robot.",
        "texts": [
            " But even though Java consists of many standardised classes for data transport over networks, the access to the underlaying operation system can differ. Bluetooth is the best example. A common API for accessing Bluetooth services, like device discovery, exists, but can behave differently on different devices, depending on the implementation of the Java Bluetooth stack interface. The authors have benchmarked the different wireless data communication technologies, and implemented a prototype of a remote control service of a mobile robot from mobile devices. For the benchmarks different mobile devices where target onto a mobile robot (see Fig. 1). This robot, connected to the facultie\u2019s Wireless LAN, is accessible over TCP/IP, and acts as the counterpart to the mobile device for the benchmark. For Bluetooth connections a USB Bluetooth dongle is connected to the robot. 978-1-4244-4881-4/09/$25.00 \u00a92009 IEEE 625 The GSM and UMTS benchmarks where made using T-Mobile\u2019s UMTS and GSM networks, which T-Mobile Germany kindly provided access to. Fig. 2 shows an simplified GSM/UMTS network setup, with the mobile device on the left. The transfered data then flow, after entering the providers network over the air, through some provider systems, before it reaches the robot, that is connected to the Internet",
            " Only some of the devices include Wireless LAN, while Bluetooth is supported by all of them. III. IMPLEMENTATION OF AN EXAMPLE SERVICE After measuring and benchmarking the different wireless technologies, a proof of concept service was implemented. Not only should this work show, what is theoretically possible in a laboratory setup. It also shows, on a real world implementation, the implications of the acquired data and gains informations about the hard to predict problems when running such a setup. The implemented service is a remote control for a mobile robot (shown in Fig. 1). The robot can be driven forwards and backwards, turned clockwise and counterclockwise as well as the PTZ camera (Pan Tilt Zoom), that delivers a live picture stream, can be controlled remotely. The server, which runs directly on the onboard computer of the robot, is now written in C++, as an earlier version written in Java had to many problems accessing system resources through the wrappers needed for Java (e.g. the Linux Bluetooth stack and the robotic programming toolkit ARIA). Direct access to the Bluetooth stack (Linux included Bluez) and standard TCP/IP sockets are combined with a multithreaded programming model using POSIX threads",
            " As a result the display sizes differ significantly. To not waste valuable bandwidth on the communication channel or processing power on the mobile device to transfer and scale oversized images, the client sends the desired size on the webcam image with the request. The server scales the webcam image to the requested size and then adds the attribution layer. This transparent overlay layer contains information about the current position and rotation of the robot relative to the starting position (see Fig. 1). UMTS/HSDPA and Wireless LAN provide good data transfer rates, but regardless of the used network technology there is never to much bandwidth. A high rate picture stream can cause a significant load on the network. If then the frequencies, used by the chosen wireless network, is jammed, the buffers kept by the involved network devices could fill up and destroy the interactivity needed to steer a robot securely. The worst case would be, if such a network jam drops in right before the user wants to issue a total stop command to prevent the robot from driving down a stairway"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000041_j.mechmachtheory.2006.05.006-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000041_j.mechmachtheory.2006.05.006-Figure1-1.png",
        "caption": "Fig. 1. Kinematic structure of fully parallel linearly actuated platform manipulators: (a) Stewart platform, (b) Hexaglide.",
        "texts": [
            " With this approach, the kinematically admissible virtual displacements and rotations of the individual links and the driving joints can all be easily transformed into the task space by means of transformation Jacobian matrices, and hence a concise parametric form of the dynamic equations in terms of the task space redundant DOF can be obtained in a more straightforward manner. The kinematic structure of fully parallel LAPs can be generally classified into two types based on the arrangements of the linear actuators. As illustrated in Fig. 1(a) and (b), the linear actuators of the Stewarttype LAP are attached to the limbs, and those of the Hexaglide-type [18] LAP are attached to the sliders along the rails fixed on the base platform. Although the procedure for kinematic analysis of the two types of robots is different [19], their dynamic equations can nevertheless be derived by using the same approach as described below. From the dynamics point of view, the LAP together with the spindle attached to the moving platform can be considered as a system of interconnected rigid bodies, and the generalized principle of D\u2019Alembert states that for such systems the total virtual work done by the effective forces along any kinematically admissible virtual displacement and rotation is zero [17], or dW a \u00fe dW e \u00bc 0; \u00f01\u00de where dWa is the total virtual work done by the gravity forces, inertia forces, and inertia moments, and dWe is that done by the externally applied forces and moments",
            " If s > 1 and jfs fs 1j 6 e, then the outer loop is converged, terminate the iteration, and then output the result; otherwise if s = 1 then set f = f + Df, else set s = s + 1 and f = f + (1/2)sDf and then go to step 2. Step 7. Set f = f (1/2)sDf and s = s + 1, and then go to step 2. Limb number i B0i \u00f0cm\u00de UVWri (cm) \u2018i (cm) qd i \u00f0cm\u00de X Y Z U V W Y X Z W V U k (0,0,15cm) 10 cm O P )(t\u03c6)(t\u03b8 P Fig. 4. Definition of the desired machining contour. The LAPs used for the examples are the Stewart platform and the Hexaglide as shown in Fig. 1. The geometric dimensions of these robots are given in Tables 1 and 2. The specified machining contour is a circle with 10 cm radius parallel to the x\u2013y plane of the fixed frame and centered at (0, 0, 15 cm), as shown in Fig. 4. The total execution time of the trajectory is specified to 10 s, and the orientation matrix and the angular velocity and acceleration of the moving platform are, respectively, specified by \u00bdR0p \u00bc \u00bdR\u00f0k; h\u00f0t\u00de\u00de 0:832 0 0:555 0:0 1 0:0 0:555 0 0:832 264 375; x0p\u00f0t\u00de \u00bc _h\u00f0t\u00dek and a0p\u00f0t\u00de \u00bc \u20ach\u00f0t\u00dek; where h\u00f0t\u00de \u00bc p\u00f01:2t3 1:6t4 \u00fe 0:6t5\u00de; 0 6 t < 1 s; 0:2pt; 1 s 6 t 6 9 s; 2p p\u00f01:2\u00f010 t\u00de3 1:6\u00f010 t\u00de4 \u00fe 0:6\u00f010 t\u00de5\u00de; 9 s < t 6 10 s 8><>: and k = [0 0 1]T"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001866_bf00251592-Figure8-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001866_bf00251592-Figure8-1.png",
        "caption": "Fig. 8. Cusp positions of Str mechanism generating a symmetrical six-bar curve: A B C G and G F D E a r e eollinear",
        "texts": [
            " Cusps The condition for a cusp in the general case is that A C G and E F G be collinear. The analysis of this case is difficult, except when limited to curves with 1 or 2 cusps. As in the Watt-1 mechanism, therefore, we confine our analysis Six-Bar Motion. II 47 to the symmetrical six-bar curve for which the joints of the triangular links are collinear. To any cusp position, then, there also corresponds a mirror-image cusp position (reflection about A E). The special case of a cusp position for a symmetrical Stephenson-1 curve is shown in Fig.8. The cusp configurations may occur with CG and GF parallel or antiparallel to A C and FE, respectively. Any cusp, therefore, lies at one of the intersections of the following four circles (see Fig. 9): Circle C1 : Center A, radius rl = s + / ; Circle C2 : Center A, radius I s - I I ; Circle C3 : Center E, radius s' + l'; Circle C4: Center E, radius Is'-l ' l . Since there are at most eight real intersections of these four circles, the maximum number of real cusps for the symmetrical six-bar curve cannot exceed eight. To find these, we apply the cosine law to triangles GBD and GAE of Fig.8: cos2= (s+l)2+(s'+l')Z-m2 (s+l-r)2+(s '+l ' -r ' )Z-n2 - ( 6 ) 2(s+t)(s'+r) 2(s + l-r)(s ' + l ' -r ' ) Equation (6) is the condition for the existence of a cusp in the position shown. This equation can also be written with 1 and l' replaced by ( - / ) and/or ( - l ' ) , resulting in a total of four equations for the four cusp positions 1, 2, 3, 4 of Fig. 9 (and their mirror images). The equations are necessary, but not sufficient to insure the existence of eight cusps. For each equation must not only be satisfied, but the value of cos 2 determined from the equation must lie within ( - 1, 1)",
            " This condition leads to the equations 12=S,(s,_r,) ( s ' r -r ' s)[s ' (s ' -r ' )-s(s-r)] ( r - s ) ( r ' - s ' ) - s s ' (7) l '2=I2-s ' (s ' -r ' )+s(s-r) , m2=2(s2 +s'2)-(r s+r' s'), n 2 = ( r - s ) 2 - s ~ + ( r ' - s ' ) 2 - s ' 2 + m 2. (8) (9) (lo) Thus when r, s, r', s' are given, equations (7 ) - (10) determine the remaining mechanism proportions. It is possible also to obtain analogous equations when r, n, r', m are given, rather than r, s, r' and s'. In order to facilitate the search for a Stephenson-1 curve with 8 real cusps, an alternative approach is used to derive the locus of G in terms of bipolar coordinates, R t and R2, as shown in Fig. 8 (this locus is, in fact, none other than the fixed centrode of the coupler motion of the four-bar mechanism A B D E). The equation for G in bipolar coordinates is: where URI(R~-R~-m2)+VR:(R~-R~-m2)+WRIRz=(R~+R~-m 2) (i) U=l/r; V=lJr'; W=(lJrr')(r2-ne+r'Z+m2). Six-Bar Motion. II 49 Equation (1) can be written for each of the cusp locations 1, 2, 3, 4 (or 1', 2', 3', 4') with the following values of R1 and R 2\" R I ( > 0 ) R 2 ( > 0 ) s+l s'+l' s + I ( s ' - l') or ( I ' - s') ( s - l) or (1- s) s' + 1' ( s - 1) or (1- s) ( s ' - 1') or (1 ' - s')",
            " However, this is possible in the special case, because the equation corresponding to (2), with ~0' instead of ~, becomes independent of t, so there are only three equations for X, Y, u, instead of four. 9. Cusps The general condition for a cusp is that G lies at the intersection of the fines through A F and ED. As in the earlier analysis of the Watt-1 and Stepheuson-1 six-bar curves, we confine our discussion to the case of a symmetrical Stephenson-2 Six-Bar Motion. II 67 curve (sin c~=sin ct'=0). For the symmetrical Stephenson-2 curve, the cusp configuration is shown in Fig. 15. This figure differs from the corresponding figure (Fig. 8) for the Stephenson-1 six-bar curve essentially only in that the triangular link on the right leg of triangle A GE, is floating. The cusp configuration in Fig. 15 is defined by the collinearity of A B F G and E D G C . Algebraically, it is defined by equating cos 2 in triangles G B C and GAE. The four cusp configurations which, when subsisting together, yield an eight-cusped, symmetrical Stephenson-2 curve are defined by the proportions R~ R 2 s' + m s + l s ' + m [ s - l l I s ' - m l s + l I s ' - m l [ s - I [ ",
            " S te ph en so n2 cu rv es 3 $2 - 05 4 \\ \\ 5 2 -I 0 1 / $2 -O 26 h W \\ \\ / / $2 - 1 02 'U $2 -0 45 $2 -2 17 All Stephenson-2 curves have been obtained via kinematic inversion of correspondingly curves have been shown. C u r v e m n l 1' 9 s S 2-002 13.715849 0.74980073 0.5 0.25 8.4926757 8.5 S 2-050 1.5 2.1428571 0.5 ]/I'J 2.1428571 3.5 s 2-101 2.2655 2.5681989 1.24 9.840612 1.9867692 6.43 S 2-102 1.2655 1.4688473 0.74 11.382064 0.93969688 7.93 S 2-217 0.2655 1.70299 1.74 8.712644 1.708423 2.93 these conditions yields the curve $2-050. It is possible also to set up a correspondence between Stephenson-1 and Stephenson-2 cusp configurations (Fig. 8, 15) by relabeling link lengths. Curves $2-101, $2-102, $2-217 were obtained in this fashion. 10. Multigeneration Since point A (Fig. 13) is the only 4-fold singular focus and point E the only 2-fold singular focus of the Stephenson-2 six-bar curve, these would have to remain the fixed pivots for any cognate mechanism (cognate meaning that point G of the cognate mechanism describes the same six-bar curve as the point G of a given Stephenson-2 mechanism). Suppose we use primes to denote the two mechanisms, i"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002716_s11340-011-9514-z-Figure3-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002716_s11340-011-9514-z-Figure3-1.png",
        "caption": "Fig. 3 (a) Experimental setup. Typical side view of a sample lens subject to (b) parallel plate compression and (c) central load compression via a ball bearing",
        "texts": [
            " When an external compressive indentation load is applied to the porous contact lens, water in the hydrogel matrix is squeezed out and the force measurements yield not only the elastic modulus of the matrix but also the bulk modulus and transport properties of water [20]. Moreover, indentation only samples the minute volume beneath the indenter which cannot represent the shell behavior under large elastic deformation as in our loading configurations. It is therefore safe to deduce that the conventional tensile test is more relevant in the present context. Experimental Setup for Parallel Plate and Central Load Compression Figure 3(a) shows the experimental setup of an Agilent T150 Universal Testing Machine with force and displacement resolutions of 30 nN and 10 nm respectively. A laser beam illuminates the sample, while an orthogonal long focal length microscope captures the shape of the deformed profile. All experiments are displacement-controlled with the shaft staying fixed and the base platform moving upwards. An aluminum cylindrical plate with radius rp=10 mm was fabricated for parallel plate compression, and a shaft with a stainless steel ball with radius rb=1 mm at the tip for central load",
            " The isotonic solution container sat on a micrometer stage with two horizontal degrees of freedom to align the loading shaft with the shell apex, while the entire setup was placed on a vibration isolation table. The buoyancy force and the liquid surface tension acting on the plate/shaft were carefully measured and subtracted from all force measurements. Great care was taken to accurately determine the moment when the probe first came into contact with the sample. It was taken to be the point when the load cell recorded a sudden jump of ~10\u03bcN, which was also confirmed by a long-focal digital camera that captured a complete side view of the sample (Fig. 3b-c). The camera was positioned horizontal and in the line of sight of the lens. To raise the optical contrast, a laser beam was used to illuminate the shell from the side so that the lens profile was highlighted. Images were then processed by using software ImageJ, a public domain Java image processing program inspired by National Institute of Health. The deformed profile was extracted by pixel intensity differentiation. Dimensions were measured by pixel scaling. Parallel Plate Compression Figure 4(a) shows the mechanical response, F(w0), of five samples, and the theoretical solution"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002390_jmems.2012.2194777-Figure7-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002390_jmems.2012.2194777-Figure7-1.png",
        "caption": "Fig. 7. Simple setup for measuring the focal length of the liquid lens.",
        "texts": [
            " For a collimated beam of light, the relationship between the transverse spherical aberration and the radius of lens aperture is expressed by [24] W (\u03bb, \u03c1) = W040\u03c1 4 (10) where \u03c1 is the incident beam diameter divided by the clear aperture of the lens and W040, which depends on wavelength (\u03bb), is the wavefront aberration coefficient. For a fixed wavelength (\u03bb), W040 is a constant. From (10), W (\u03bb, \u03c1) can be decreased largely by shrinking the aperture of the lens ID. Such a conclusion is still valid if the incident beam is white light. The tradeoff is that decreasing the lens aperture will cause the light intensity through the lens to decrease. The focal length of the lens with different apertures can be measured using the simple setup as shown in Fig. 7. The liquid lens was mounted on a linear metric stage. A He\u2013Ne laser beam (\u03bb = 0.633 \u00b5m) was collimated and expanded by lens-1 and lens-2. The beam was used to illuminate the liquid lens. After passing through the liquid lens, the beam was converged. The light intensity at the focal point was received and analyzed using a CCD camera (SBIG ST-2000XM). From the focal point to the liquid lens center, the distance (f) is the focal length of the lens. The focal length versus the variable aperture is plotted in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002449_ijcat.2010.032200-Figure5-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002449_ijcat.2010.032200-Figure5-1.png",
        "caption": "Figure 5 The illustration of the three-dimensional coaxial powder feeding system",
        "texts": [
            " Two other advantages of the coaxial powder supply are the controllable heating of the powder before it enters the molten pool and the high powder feed efficiency (Jouvard et al., 1997; Lemoine et al., 1994; Li et al., 1995; Vetter et al., 1994; Nan et al., 2006). Figure 4 shows the working coaxial powder nozzle of the LRS. To perform three-dimensional repair, a corresponding powder feeding nozzle has been developed. In this research, a three-dimensional coaxial powder feeding subsystem is integrated in the cladding system (Figure 5). The LRS is based on the integration of laser cladding and reverse engineering (Figure 6). Figure 7 shows the LRS machining developed by Shenyang Institute of Automation, Chinese Academy of Sciences. To start the repair process, a high-powered laser beam strikes a small spot (approximately 0.5 mm wide) on the surface of the damaged metal component, producing a molten pool. Then, the powder feed nozzle delivers metal powder into the molten pool to increase its volume. This process is repeated within a plane to create a single metal layer",
            " With laser remanufacturing technology, there is no interference between tools and formed parts, so the dimension shape can be very complicated (Figure 14). Moreover, the as-deposited parts are fully dense, hold rapid solidified microstructure, and meet the requirements for direct usage. During the machining process, the substrate moves in the X\u2013Y plane beneath the laser beam to deposit a thin cross-section, thereby creating the desired geometry for each layer. After deposition of each layer, the position of the powder feed nozzle and the focusing laser beam are incremented in the positive Z-direction (see Figure 5). As a result, successive layers are deposited to produce the entire three-dimensional component volume of fused metal representing the desired CAD model. In nature, this forming procedure is multi-layer laser cladding. This opens the door to fabricate replacement components for parts that are not repairable at the time they are needed. With the ability of one-step manufacture, this technology can greatly reduce the lead-time and investment cost of mould and die design, the fabrication of hard or rare metal components, the repair of refractory and costly components (Li et al"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003869_ls.1222-Figure3-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003869_ls.1222-Figure3-1.png",
        "caption": "Figure 3. Drawing and schematic view of the test gearbox.",
        "texts": [
            " Copyright \u00a9 2013 John Wiley & Sons, Ltd. Lubrication Science 2013; 25:297\u2013311 DOI: 10.1002/ls Copyright \u00a9 2013 John Wiley & Sons, Ltd. Lubrication Science 2013; 25:297\u2013311 DOI: 10.1002/ls Speed and torque sensors were mounted before and after the test gearbox, represented by Sen.1 and Sen.2, in Figure 2. The sensor Sen.2 was mounted in a movable platform that allows the adjustment of its height and depth. The speed range of the test rig goes from 100 to 4300 rpm, and the maximum input torque is 1330Nm. Figure 3 shows, schematically, the test gearbox used in this work. The gearbox had two kinematic relations, obtained using five gears mounted on three shafts. Pinions 4 and 5 were mounted on needle roller bearing, and the transmission relation was obtained through the engagement of one of these pinions. The test gearbox was used as a speed multiplier with a transmission ratio of 0.44, engaging the gear pairs 1\u20132 and 3\u20135, with shaft 1 connected to Sen.1, whereas the slave gearbox was used as a speed reducer"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003019_c0ay00632g-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003019_c0ay00632g-Figure2-1.png",
        "caption": "Fig. 2 Schematic diagram of the new ECL cell.",
        "texts": [
            " The three different solutions were mixed in three-limb tubes, and then were injected into the cell, which was placed in front of the detection window of the photomultiplier tube (PMT) ( 800 V) of the system. The cyclic voltammetry mode of CHI 760D electrochemical unit was used for the electrolysis. When the potential is higher than 0.9 V, an obvious ECL emission can be observed. The ECL emission was detected with PMT and recorded by the computer employing an IFEL-E analysis software system. The ECL cell31 (Fig. 2) was fabricated by a top cuboid PMMA, a bottom columniform PMMA, an inlet, an outlet, an Ag/AgCl reference electrode (RE), a Pt counter electrode (CE) and an ITO working electrode (WE). The inlet, outlet, RE and CE were located in the top cuboid PMMA. The 1.0 cm i.d ITO slide glass 1164 | Anal. Methods, 2011, 3, 1163\u20131167 was sonicated for 10 min to clean the ITO surface in each of the following solvents: acetone, 10% NaOH/ethanol and distilled water, and was then used as the working electrode. The cleaned pieces were kept in methanol and rinsed with doubly distilled water before use"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001884_s11431-010-3100-y-Figure6-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001884_s11431-010-3100-y-Figure6-1.png",
        "caption": "Figure 6 The 2-PRU/PR(Pa)R mechanism.",
        "texts": [
            " In loop II, all R-, P-axes of the link group RRRP are parallel to z0-axis, the link group cannot rotate about x, y-axes, its rank is dgz II (0 0, x y z)=4, dX II = d A,B I (0 0 0x y z) + dgz II (0 0 , x y z)=dX II (0 0 , x y z)=4, FII = PII i=1fi dX II= 44= 0, F=FI + FII = 3+0=3.From eq. (4), F=12(54)=3. With the method of virtual loop, we have L j=1d X j =5+4 = 9, which is consistent with the result by Gogu [5], who obtained it by solving the rank of the kinematic equations via a 12\u00d712 matrix. It is proved that dX is the rank of kinematic constraint equations. The actions of these two definitions are identical. Example 4. Figure 6 is a 2-PRU/PR(Pa)R mechanism formed by three loops. Loop I is the loop ABCDFE and lies on plane O-yz. It consists of a link group with 2-PRU (C, D are U-pair). Loop II is the GMJHKDEF and contains a link ~ ~ group with PRR//RR. Loop III is the loop HJPT and includes a link group with 2R. Since JH//PT and JH = PT, Loop III is the parallelogram mechanism with two cranks (Figure 7). In loop I, all links cannot rotate about x-axis nor translate about z-axis. Its rank is dI = dX I (, 0 y z)=4, FI = PI i=1fi dX I =84=4"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001623_elan.201000221-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001623_elan.201000221-Figure2-1.png",
        "caption": "Fig. 2. CV at : (a) GCE electrode, (b) PB-SiO2, (c) APTES/PBSiO2, (d) CS-nanoAu/APTES/PB-SiO2, (e) anti-NSE/CSnanoAu/APTES/PB-SiO2 and (f) BSA/anti-NSE/CS-nanoAu/ APTES/PB-SiO2modified electrodes at 50 mV/s in 10 mM phosphate buffer solution (pH 7.0) and all potentials are given vs. SCE.",
        "texts": [
            " The TEM images of Figure 1(a and b) illustrate that the SiO2 and PB-SiO2 nanocomposites. As shown in Figure 1a, particlelike SiO2 nanoparticles with particle diameter of 50\u2013 100 nm are observed. After coated with PB, a relatively compacted and smooth nanoparticles can be found, and the size of PB-SiO2 become slightly larger than SiO2 nanoparticles. From the SEMs of SiO2 and PB-SiO2 nanocomposites, we can observe the SiO2 still retains the highly dispersive structure (Figure 1c), while the PB-SiO2 (Figure 1d) is more uniform and smoother than SiO2. Figure 2 displays cyclic voltammetry electrochemical behavior of different modified electrodes. It is difficult to achieve the typical redox peaks at bare GCE in pH 7.0 phosphate buffer solution as the lack of redox probes (Figure 2a). Then a pair of well-defined redox peaks was appeared in Figure 2b with PB-SiO2 nanocomposites modified GCE, indicating the immobilized PB-SiO2 may act as a redox probes. With the adsorption of APTES, the CV current decreased a little (Figure 2c) owing to the fact that APTES could obstructs electron and mass transfer. Compared with curve c, the current of curve d increased after adsorption of CS-nanoAu is most likely due Scheme 1. Schematic drawing of the stepwise immunosensor fabrication process. (a) dropping of PB-SiO2 membrane; (b) functionalizing with APTES; (c) assembly of CS-nanoAu; (d) immobilization of anti-NSE; (e) blocking with BSA. Electroanalysis 2010, 22, No. 21, 2569 \u2013 2575 2010 Wiley-VCH Verlag GmbH & Co. KGaA, Weinheim www.electroanalysis.wiley-vch.de 2571 to the fact that the CS-nanoAu may facilitate improve the conductivity of APTES/PB-SiO2 composite film. After the immobilization of anti-NSE onto the electrode surface, the redox peaks decreased as the immune-protein film acts as an inert electron and mass transfer blocking layer (Figure 2e). Subsequently, BSA was employed to block the possible remaining activity sites, thus, a further decrease can be observed in Figure 2 curve f. The catalytical activity of anti-NSE/CS-nanoAu/APTES/ PB-SiO2 nanostructural multifilms modified electrode was investigated. The curve a in Fig. 3 shows the quasireversible cyclic voltammogram of the immunosensor. Upon addition of 0.75 mM H2O2 in N2-saturated phosphate buffer solution, the CV of the reduction current increased, while the oxidation current density decreased obviously (Figure 3b), indicating the good electrocatalytic activity toward reduction of H2O2, leading to a very low over potential [23]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001531_s11012-009-9232-0-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001531_s11012-009-9232-0-Figure1-1.png",
        "caption": "Fig. 1 Denavit\u2013Hartenberg parameters",
        "texts": [
            " Excepting the Schatz 6R mechanism, called \u201cTurbula\u201d [17] and some deployable structures of Gan [10] and Chen [7], we have no knowledge of existence of other industrial applications of 6R overconstrained mechanisms. We consider the Denavit-Hartenberg formalization for a mechanism with revolute joints [9]. The twist angle between two successive joints (i) and (i + 1) is noted \u03b1i , the bar lengths is noted ai and the offset distance between two elements (i \u2212 1) and (i) is noted di . The fourth parameter is the angle between two successive elements (i \u2212 1) and (i), noted \u03b8i . Figure 1 shows the schematic representation of the revolute joints with the Denavit and Hartenberg notations, in which ai, \u03b1i and di are geometrics parameters and \u03b8i is the kinematical variable. The homogeneous form for the transfer matrix is: i\u22121Qii = \u23a1 \u23a3 cos \u03b8i \u2212 cos\u03b1i \u00b7 sin \u03b8i sin\u03b1i \u00b7 sin \u03b8i ai \u00b7 cos \u03b8i sin \u03b8i cos\u03b1i \u00b7 cos \u03b8i \u2212 sin\u03b1i \u00b7 cos \u03b8i ai \u00b7 sin \u03b8i 0 sin\u03b1i cos\u03b1i di 0 0 0 1 \u23a4 \u23a6 (1) The closure condition for a 6R single loop mechanism expresses that the six transfer matrix product is equal with the unity matrix"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002057_20090819-3-pl-3002.00080-Figure3-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002057_20090819-3-pl-3002.00080-Figure3-1.png",
        "caption": "Fig. 3. Hydraulic motor.",
        "texts": [],
        "surrounding_texts": [
            "As for modelling, the considered mechatronic system is divided into a hydraulic subsystem and a mechanical subsystem, which are coupled by the torque generated by the hydraulic motor (Schulte [2007]). Here, the actuator dynamics of the displacements units is taken into account by a first order lag element, i.e., a PT1-system. Consequently, the dynamics of displacement unit of the pump, which is realised as a tiltable swashplate, is governed by TuP \u02d9\u0303\u03b1P + \u03b1\u0303P = kP uP, (1) whereas the dynamics of the motor according to a bend axis design is described by TuM \u02d9\u0303\u03b1M + \u03b1\u0303M = kM uM. (2) Here, a normalised swashplate angle \u03b1\u0303P = \u03b1P/\u03b1maxP and a normalised bend axis angle \u03b1\u0303M = \u03b1M/\u03b1maxM have been introduced. The volume flow rate of the pump can be stated as a nonlinear function of the swashplate angle qP = VP(\u03b1P)nP = V\u0303P(\u03b1P)\u03c9P, (3) with V\u0303P = VP/2\u03c0 . Nevertheless, the reasonable assumption of a small swashplate angle |\u03b1P| \u2264 18\u25e6 , c.f. Kugi et al. [2000], allows for the approximation qP = V\u0303P\u03b1\u0303P\u03c9P, (4) with \u03b1\u0303P \u2208 {\u22121,1}. Accordingly, the volume flow rate into the hydraulic motor can be formulated as a nonlinear function of the bend axis angle qM = VM(\u03b1M)nM = V\u0303M(\u03b1M)\u03c9M, (5) with V\u0303M = VM/2\u03c0 . As before, the assumption of small angles |\u03b1P| \u2264 20\u25e6 leads to a simplified relationship qM = V\u0303M\u03b1\u0303M\u03c9M, (6) with \u03b1\u0303M \u2208 {0,1}. The pressure dynamics of the hydrostatic transmission in closed-circuit configuration can be modelled by two capacitances (Jelali and Kroll [2004], Kugi et al. [2000]), which account for both the fluid compressibility and the elasticity of the connecting hoses qCA = CA p\u0307A, qCB = CB p\u0307B. (7) The internal leakage oil flow is modelled as a laminar flow resistance (Jelali and Kroll [2004]), which depends linearly on the difference pressure qleak = kleak(pA\u2212 pB) (8) characterized by the leakage coefficient kleak. The external leakage could be introduced as well but shall be neglected in the control-oriented design model. A volume flow balance leads directly to the resultant volume flows in the two capacitances qP\u2212qM\u2212qleak = qCA , (9) \u2212qP +qM +qleak = qCB . (10) A model order reduction becomes possible, if some reasonable symmetry assumptions are made. Considering identical capacitances CA = CB =: CH leads to an order reduction by one and results in a differential equation for the difference pressure \u2206p\u0307 = 2 CH ( V\u0303P \u03b1\u0303P\u03c9P\u2212V\u0303M \u03b1\u0303M\u03c9M\u2212 kleak\u2206p ) . (11) The longitudinal dynamics of the working machine is governed by the equation of motion. The vehicle with the drive chain system (vehicle mass mv, wheel radius rw, gear box transmission ratio ig, rear axle transmission ratio ia, damping coefficient dvc at the drive shaft, see also Fig. 4) can be described by the following second order differential equation mv r2 w i2a i2g\ufe38 \ufe37\ufe37 \ufe38 JV \u03c9\u0307M + dg i2g\ufe38\ufe37\ufe37\ufe38 dvc \u03c9M = V\u0303M\u2206p \u03b7M\u03b1\u0303M\ufe38 \ufe37\ufe37 \ufe38 \u03c4M \u2212 \u03b7g\u03c4L iaig\ufe38 \ufe37\ufe37 \ufe38 \u03c4U . (12) Here, the torque \u03c4M = V\u0303M\u2206p\u03b7M\u03b1\u0303M of the hydraulic motor depends on its mechanical-hydraulic efficiency \u03b7M . The overall system model involves four first order differential equations. Introducing the normalized angles of the displacement units \u03b1\u0303i, i \u2208 {P,M}, the difference pressure \u2206p, and the motor angular velocity \u03c9M as state variables, the corresponding state-space representation becomes  \u02d9\u0303\u03b1P \u02d9\u0303\u03b1M \u2206p\u0307 \u03c9\u0307M =  \u2212 1 TuP \u03b1\u0303P + kP TuP uP \u2212 1 TuM \u03b1\u0303M + kM TuM uM \u2212 2kleak CH \u2206p+ 2V\u0303P CH \u03c9P\u03b1\u0303P\u2212 2V\u0303M CH \u03c9M\u03b1\u0303M \u2212 dVC JV \u03c9M + V\u0303M\u03b7M JV \u2206p\u03b1\u0303M\u2212 \u03c4U JV  . (13) The input voltages ui, i \u2208 {P,M}, of the corresponding proportional valves for the displacement units serve as physical control inputs. Feedfoward compensation of the actuator dynamics using its approximated inverse in form of a proper transfer function, as seen in Fig. 5, leads to a simplified control design with u\u0304i \u2248 \u03b1\u0303i, where u\u0304i represents the inverse input. Therefore, only the differential equations for the difference pressure and the motor angular velocity are used for feedback control design. The remaining model uncertainties are taken into account by the disturbance torque \u03c4U . On the one hand, these uncertainties stem from external load forces and drive resistances acting on the vehicle. On the other hand, a varying vehicle mass mv contributes to parameter uncertainty."
        ]
    },
    {
        "image_filename": "designv11_12_0000546_pes.2007.386135-Figure11-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000546_pes.2007.386135-Figure11-1.png",
        "caption": "Fig. 11. Flux distribution when the rotor with iron core is used in the bearingless motor with short-pitch winding.",
        "texts": [
            " 9 shows the computed rotational torques in the shortpitch winding model. The torque ripple is small because this machine type is a brushless DC motor. The difference between the computed results in the rotors with and without iron core is also small. Fig. 10 shows the computed results of suspension force in the x- and y-directions. The difference between the suspension forces in the rotors with and without iron core is quite small though it was large in the full-pitch winding model as shown in Fig. 8. Fig. 11 shows the flux distribution when the rotor with iron core is used in the short-pitch winding model. The rotor angular position is 12 [deg] when the suspension windings sv1 and sv2 are excited, in which the suspension flux is superposed on the PM field flux. It is seen that both of the PM flux and suspension flux does not go through the iron core within the PM in the rotor very much. Thus, the suspension force is not varied even when the iron core is replaced by the air within the PM. Therefore, the suspension forces in the rotor with iron core are almost equal to those in the coreless rotor in the short-pitch winding model"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001517_j.mechmachtheory.2009.12.004-Figure3-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001517_j.mechmachtheory.2009.12.004-Figure3-1.png",
        "caption": "Fig. 3. A six-bar linkage with two sliders.",
        "texts": [
            " (18) represent the boundaries of the JRS, where the five-bar chain is at singularities and the three non-input joints E, F and G are collinear. For a given pair of h2 and h3 values, two solutions, i.e. x5\u00bd1 \u00bc Q 2 ffiffiffiffiffiffi D2 p 2P2 and h5 \u00bc 2 arctan\u00f0x5\u00de; \u00f020a\u00de x5\u00bd2 \u00bc Q 2 \u00fe ffiffiffiffiffiffi D2 p 2P2 and h5 \u00bc 2 arctan\u00f0x5\u00de \u00f020b\u00de may be obtained from Eq. (15). Each real solution corresponds to a linkage configuration. It is noted that h5 2 ( p, p] in a full cycle. Since arctan(x5) 2 ( p/2, p/2], each real solution of x5 corresponds to a unique h5 if h5 2 ( p, p]. 4R1P Loop: The Stephenson type linkage in Fig. 3 contains a 4R1P loop. The loop equation for the 4R1P chain ABEFGA in Fig. 3 can be expressed as a2eih2 \u00fe a4ei\u00f0h3\u00feb\u00de a1 a7 a5eih5 s2eia1 \u00bc 0: \u00f021\u00de Eliminating s2 and using the tangent half-angle formula of Eq. (14), Eq. (21) can be written in the quadratic form as Eq. (15) and from which equations in the form of Eqs. (17) and (20) can be derived. 3R2P Loop: The Stephenson type linkage in Fig. 4 contains a 3R2P loop. The loop equation for the 3R2P chain ABEFGA in Fig. 4 can be expressed as a2eih2 \u00fe a4ei\u00f0h3\u00feb\u00de a1 a7 s3ei\u00f0a1\u00fec\u00de s2eia1 \u00bc 0: \u00f022\u00de Eliminating s2, Eq. (22) can be written as s3 sin c a2 sin\u00f0h2 a1\u00de a7 sin a1 a4 sin\u00f0h3 \u00fe b a1\u00de \u00bc 0: \u00f023\u00de The discriminant of Eq",
            " Sub-branch g-l-h: Segment g-l-h with h2 2 [45.3 , 87.3 ] and Eqs. (9a) and (20a); Sub-branch g-k-h: Segment g-k-h with h2 2 [45.3 , 87.3 ] and Eqs. (9b) and (20a). Branch h-g: This branch contains the branch h-g of the four-bar chain with h2 2 [45.3 , 87.3 ] and satisfies Eqs. (4) and (20b). Sub-branch h-l-g: Segment h-l-g with h2 2 [45.3 , 87.3 ] and Eqs. (9a) and (20b); Sub-branch h-k-g: Segment g-k-h with h2 2 [45.3 , 87.3 ] and Eqs. (9b) and (20b). Example 3. Given the dimensions for the single-DOF double-loop linkage in Fig. 3 as follows, a1 = 2.4, a2 = 5.5, a3 = 4.8, a4 = 4.2, a5 = 2.5, a7 = 2.4, b = 45.0 , a1 = 30.0 , a2 = 50.0 . With the above given dimensions, the plot for the linkage is shown in Fig. 12, where the 3R1P I/O curve is drawn from Eq. (12) and the motion domain (shade area) is drawn from Eq. (17). The mobility analysis of the single-DOF double-loop linkage with the proposed method above can be carried out as follows, (1) Branch identification of 3R1P chain: From Eq. (8) with D = 0, there are two dead center positions u-( 162",
            " (9a) and (20a); Sub-branch 4-3: Segment 4-3 with h2 2 [ 105.2 , 15.6 ] in Eqs. (9a) and (20b). Branch 6-5: This branch contains segment 6-5. Sub-branch 6-5: Segment 6-5 with h2 2 [7.9 , 72.6 ] in Eqs. (9b) and (20a); Sub-branch 5-6: Segment 5-6 with h2 2 [7.9 , 72.6 ] in Eqs. (9b) and (20b). Example 4. Given the same dimensions for the single-DOF double-loop linkage in Fig. 4 as in the above example and sin c \u2013 0. It may be noted that with the same dimensions for the single-DOF double-loop linkage in Fig. 4 as in Fig. 3, the branches of the linkage are determined by the 3R1P chain and each branch of the 3R1P chain represents one branch of the whole linkage since Eq. (24) is always satisfied and there is only one solution to s3 for a given pair of h2 and h3 and sin c \u2013 0. Thus, there is only one branch for the linkage in Fig. 4 in this example. This paper presents the first successful attempt that extends the discriminant method to the mobility identification of any type of single degree-of-freedom (DOF) double-loop planar linkages under any input and output condition"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002129_s0263574709005426-Figure5-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002129_s0263574709005426-Figure5-1.png",
        "caption": "Fig. 5. Screw movement or twist.",
        "texts": [
            " (9) The general spatial differential movement of a rigid body consists of a differential rotation about, and a differential rotation along an axis named the instantaneous screw axis. In this way, the velocities of the points of a rigid body with respect to an inertial reference frame O-xyz may be represented by a differential rotation \u03c9 about the instantaneous screw axis and a simultaneously differential translation \u03c4 about this axis. The complete movement of the rigid body, combining rotation and translation, is called screw movement or twist and is here denoted by$. Fig. 5 shows a body \u2018twisting\u2019 around the instantaneous screw axis. The ratio of the linear velocity to the angular velocity is called the pitch of the screw h =\u2016 \u03c4 \u2016 / \u2016 \u03c9 \u2016. The twist may be expressed by a pair of vectors, i.e. $ = [\u03c9T; V T p ]T, where \u03c9 represents the angular velocity of the body with respect to the inertial frame, and Vp represents the linear velocity of a point P attached to the body which is instantaneously coincident with the origin O of the reference frame. A twist may be decomposed into its magnitude and its corresponding normalized screw"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003803_j.conbuildmat.2015.07.152-Figure4-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003803_j.conbuildmat.2015.07.152-Figure4-1.png",
        "caption": "Fig. 4. Schematic drawing of the James Machine prescribed in ASTM D 2047.",
        "texts": [
            " Schematic drawing of the slider and 10 \u00b1 0.5 mm in width. The front edge of the slider was beveled at 35 \u00b1 5 . Two of the three sliders, placed at the front, were leather sliders with a density of 1.0 \u00b1 0.1 g/cm3, constructed using Shore D hardness 60 \u00b1 10 tanned leather. The back slider is a shoe-rubber slider constructed using Shore A hardness 95 styrene-butadiene rubber. EN 13893 prescribes that the test be performed on a cleaned floor. 3.2. Test prescribed in ASTM D 2047 The ASTM D 2047 test uses a James Machine, shown in Fig. 4, to obtain the static coefficient of friction. A specimen is fixed to the test table, and a shoe, which has a load applied to it via a strut, is placed on the top of the sample. The test table is then moved assembly prescribed in EN 13893. horizontally at a constant speed while the shoe also moves, causing the angle of the strut to gradually increase. Hence, the vertical load on the shoe gradually decreases while the horizontal load increases. The test table is moved farther to obtain the shoe\u2019s traveling distance at the point in time at which the shoe slides and the vertical column drops"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000973_978-3-642-04466-3_5-Figure14-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000973_978-3-642-04466-3_5-Figure14-1.png",
        "caption": "Fig. 14 Direction of drag (D) and side force (S) relative to wind direction (top view)",
        "texts": [
            " The AFL ball and the Rugby ball were tested at 60, 80, 100, 120, and 140 km h 1 speeds under \u02d990\u0131 yaw angles with an increment of 10\u0131. The balls were yawed relative to the force sensor (which was fixed with its resolving axis along the mean flow direction) thus the wind axis system was employed. Flow visualizations by wool tufts at 0\u0131 and 90\u0131 yaw angles for the AFL ball and the Rugby ball are shown in Figs. 12 and 13 respectively. The direction of wind with respect to the direction of drag force .D/ and side force .S/ is shown in Fig. 14. The aerodynamic forces were converted to non-dimensional parameters (drag coefficient, CD and side force coefficient, CS) and tare forces were removed by measuring the forces on the sting in isolation and removing them from the force of the ball and sting. The influence of the sting on the balls was checked and found to be negligible. Only drag force and side force coefficients are presented in this work and they are plotted against yaw angles. The CD and CS were defined as CD D D 1=2 V 2 A and CS D S 1=2 V 2 A , where A is the projected frontal area of the ball defined as A D D2 4 "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003362_j.proeng.2013.08.190-Figure6-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003362_j.proeng.2013.08.190-Figure6-1.png",
        "caption": "Fig. 6. Experimental setup for Active Target tets.",
        "texts": [
            " To see the influence of these factors tests have been developed to measure changes in position of the reflector at its rotation. The rotation about the vertical axis is simple, just focus the spotlight on the table of roundness and align with an interferometer. Rotating the roundness table, the variation on the interferometer measurement corresponds to the axis eccentricity. The rotation of the horizontal shows more problems and a 45\u00ba support has been proposed so it will produce the simultaneous rotation of the two axes as shown in Figure 6. Previously knowing the vertical axis error, the corresponding to the horizontal axis can be calculated. It has been shown the beginning of a process which will lead to with the development of a simple procedure for the calibration of laser tracker systems. The work done so far include the tests on SMR and planning of them with Active Target. Also the definition of LT models to study and geometric errors is at an advanced stage along with the definition of an automatic generator of synthetic data with simulated errors"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000597_978-3-540-73129-0-Figure6.149-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000597_978-3-540-73129-0-Figure6.149-1.png",
        "caption": "Fig. 6.149. Actuator with electrorheological fluid. a Physical arrangement, b equivalent circuit diagram",
        "texts": [
            " The response time of electrorheological fluids lies in the range of milliseconds. From an electrical point of view, they are capacitive loads with a parallel conductance. Capacitance and conductance are determined by the geometry of the assembly and by the physical properties of the employed fluids, such as the specific electrical resistance \u03c1 and the permittivity \u03b5. When computing the time constant for self-discharge \u03c4 , the geometrical parameters cancel each other, so that the time constant is a property of the fluid. Figure 6.149a illustrates a capacitor with the fluid between its parallel plates. Figure 6.149b displays its equivalent circuit diagram. If the actuator discharges itself too slowly such that it cannot be deactivated at a certain operating frequency, the system requires an additional discharging circuit (two-quadrant operation). If the driving frequency is so low that the actuator can discharge itself sufficiently fast, then the charging circuit will suffice (one-quadrant operation). Retroactive Effects on the Power Electronics. The operating modes of electrorheological actuators are classified into three types: flow mode (or valve mode), shear mode and squeeze mode"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001065_13506501jet487-Figure21-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001065_13506501jet487-Figure21-1.png",
        "caption": "Fig. 21 Engine drive train layout with fitted pulley torque transducers for measuring camshaft friction",
        "texts": [],
        "surrounding_texts": [
            "1 Craig, R. C., King, W. H., and Appeldoorn, J. K. Oil film thickness in bearing of a fired engine, part II: the bearing as a capacitor. SAE paper 821250, 1982. 2 Choi, J. K., Lee, J. H., and Han, D. C. Oil film thickness in engine main bearings: comparison between calculation and experiment by total capacitance method. SAE paper 922345, 1992. 3 Choi, J. K., Min, B. S., and Han, D. C. Effect of oil aeration rate on the minimum oil film thickness and reliability of engine bearing. SAE paper 932785, 1993. 4 Bates, T. W. and Benwell, S. Effect of oil Rheology on journal bearing performance, part 3: Newtonian oils in the connecting rod bearing of an operating engine. SAE paper 880679, 1988. 5 Filowitz, M. S., King, W. H., and Appeldoorn, J. K. Oil film thickness in a bearing of a fired engine. SAE paper 820511, 1982. 6 Bates, T. W., Willamson, J. A., Spearot, J. A., and Murphy, C. K. A correlation between engine oil rheology and oil film thickness in engine journal bearing. SAE paper 860376, 1986. 7 Spearot, J. A. and Murphy, C. K. A comparison of the total capacitance and total resistance techniques for measuring the thickness of journal bearing oil films in an operating engine. SAE paper 880680, 1988. 8 Suzuki, S., Ozasa,T.,Yamamoto, M., Nozawa,Y., Noda,T., and Ohori, M. Temperature distribution and lubrication characteristics of connecting rod big end bearings. SAE paper 953550, 1995. 9 Choi, J. K., Min, B. S., and Han, D. C. Effect of oil aeration rate on the minimum oil film thickness and reliability of engine bearing. SAE paper 932785, 1993. 10 Choi, J. K., Min, B. S., and Oh, D. Y. A study on the friction characteristics of engine bearing and cam/tappet contacts from the measurement of temperature and oil film thickness. SAE paper 952472, 1995. 11 Cho, M. R., Han, D. C., and Choi, J. K. Oil film thickness in engine connecting-rod bearing with consideration of thermal effects: comparison between theory and experiment. ASME J. Tribol. 1999, 121, 901\u2013907. 12 Irani, K., Pekkari, M., and Angstrom, H. E. Oil film thickness measurement in the middle main bearing of a six-cylinder supercharged 9 litre diesel engine using capacitive transducers. Wear, 1997, 207, 29\u201333. 13 Masuda, T. A measurement of oil film pressure distribution in connecting rod bearing with test rig. Tribol. Trans., 1992, 35(1), 71\u201376. 14 Schilowitz, A. M. and Waters, J. L. Oil film thickness in a bearing of a fired engine, part IV: measurements in a vehicle on the road. SAE paper 861561, 1986. 15 Mihara, Y., Hayashi, T., Nakamura, M., and Someya, T. Development of measuring method for oil film pressure of engine main bearing by thin film sensor. JSAE, 1995, 16, 125\u2013130. 16 Helena, R., Antti, V., Simo, V., Ari, H., Markku, K., Ingmar, S., Juha, V., Ari, L., and Juhani, M. Oil film pressure measurement in journal bearing tests by using an optical sensor. In Proceedings of the 35th Leeds\u2013Lyon Symposium on Tribology, Leeds, 2008. 17 Warrens, C., Jefferies, A. C., Mufti, R. A., Lamb, G. D., Guiducci, A. E., and Smith, A. G. Effect of oil rheology and chemistry on journal-bearing friction and wear. Proc. IMechE, Part J: J. Engineering Tribology, 2008, 222(J3), 441\u2013450. DOI: 10.1243/13506501JET348. 18 Mufti, R. A.,Warrens, C., Lamb, G. D., and Jefferies, A. C. Effect of oil formulation on engine journal bearing performance. In Proceedings of Additive 2007, Application for future transport, London, UK, 2007. 19 Mufti, R. A., Warrens, C., Lamb, G. D., Guiducci, A. E., and Jefferies, A. C. Effect of viscosity modifiers on the bearing load carrying capacity of crankcase lubricants. In Proceedings of the IMechE Tribology Conference on Surface Engineering and Tribology for Future Engines and Drivelines, London, UK, 2006. 20 Tanaka, M. Journal bearing performance under starved lubrication. Tribol. Int., 2000, 33, 259\u2013264. 21 Lu, X. and Khonsari, M. M. On the lift-off speed in journal bearing. Tribol. Lett., 2005, 20(3\u20134), 299\u2013305. 22 Cerrato, R., Gozzelino, R., and Ricci, R. A single cylinder engine for crankshaft bearings and piston friction losses measurement. SAE paper 841295, 1984. 23 Syverud, T. Experimental investigation of the temperature fade in the cavitation zone of full journal bearings. Tribol. Int., 2001, 34, 859\u2013870. 24 Sinanoglu, C., Nair, F., and Karamis, M. B. Effects of shaft surface texture on journal bearing pressure distribution. J. Mater. Process. Technol., 2005, 168, 344\u2013353. 25 Sorab, J., Korce, S., Brower, C. L., and Hammer, W. G. Friction reducing potential of low viscosity engine oils in bearings. SAE paper 962033, 1996. 26 Tseregounis, S. I.,Viola, M. B., and Paranjpe, R. S. Determination of bearing oil film thickness (BOFT) for various engine oils in an automotive gasoline engine using capacitance measurements and analytical predictions. SAE paper 982661, 1998. 27 Unlu, B. S. and Atik, E. Determination of friction coefficient in journal bearings. J. Mater. Des., 2007, 28, 973\u2013977. 28 Mufti, R. A. and Priest, M. Experimental and theoretical study of instantaneous engine valve train friction. ASME J. Tribol., 2003, 125, 628\u2013637. 29 Mufti, R. A. and Priest, M. Experimental evaluation of piston assembly friction under motored and fired conditions in a gasoline engine. Am. Soc. Mech. Eng., J. Tribol., 2005, 127, 826\u2013836. 30 Mufti, R. A. and Priest, M. Experimental evaluation of engine valve train friction under motored and fired conditions. Tribol. Res. Des. Eng. Syst., Tribol. Ser. (Elsevier), 2003, 41, 767\u2013778. 31 Cameron, A. Basic lubrication theory, 3rd edition, 1981 (Ellis Horwood Ltd). 32 Taylor, C. M. Engine tribology, 1993, Tribology Series 26 (Elsevier Science). Proc. IMechE Vol. 223 Part J: J. Engineering Tribology JET487 \u00a9 IMechE 2009 at University of Ulster Library on March 31, 2015pij.sagepub.comDownloaded from 33 Booker,J. F. Dynamically loaded journal bearings: mobility method of solution. Trans. ASME, D, J. Basic Eng., 1965, 187, 537\u2013546. 34 Booker, J. F. A table of the journal bearing integral. Trans. ASME, D, J. Basic Eng., 1965, 187, 533\u2013535. 35 Yang, L. S. Friction modelling for internal combustion engines. PhD Thesis, Department of Mechanical Engineering, University of Leeds, UK, 1992, p. 230. 36 Goenka, P. K. Dynamically loaded journal bearings: finite element method analysis. Trans. ASME, J. Tribol.,1984, 106(1), 429\u2013439. 37 Dickenson, A. N. Engine friction modelling considering lubricant tribological characteristics. PhD Thesis, School of Mechanical Engineering, University of Leeds, UK, 2000, p. 232. 38 Martin, F. A. Friction in internal combustion engines. In Proceedings of the Institution of Mechanical Engineers Conference on Reduction of Friction and Wear in Combustion Engines, 1985, pp. 1\u201317, paper C67/85. 39 Booker, J. F. Dynamically loaded journal bearings: numerical application of the mobility method. Trans. ASME, F, J. Lubr. Technol., 1971, 93\u201394, 168\u2013176. 40 Frene, J., Nicolas, D., Degueurce, B., Berthe, D., and Godet, M. Hydrodynamic lubrication, bearing and thrust bearings, 1997 (Elsevier Science). 41 Gulwadi, S. D. and Shrimpling, G. Journal bearing analysis in engines using simulation techniques. SAE paper 2003-01-0245, 2003. 42 Bates, T. W., Fantino, B., Launay, L., and Frene, J. Oil film thickness in an elastic connecting rod bearing: comparison between theory and experiment. STLE, 1990, 33, 254\u2013266. APPENDIX 1 Notation ao piston acceleration A area b bearing length c bearing radial clearance C out of balance load d bearing diameter D cylinder liner diameter e eccentricity F bearing load Fm main bearing load Frec reciprocating inertia force Frot rotating inertia force FB big-end bearing load FG gas force FT side thrust force h film thickness H bearing friction power loss mc connecting rod mass mp piston assembly mass M \u03b5, M \u03c6 mobility components p pressure r bearing journal radius Ra crank radius S piston displacement V journal centre velocity wb bearing bush angular velocity wj bearing journal angular velocity wl bearing load angular velocity Z ,W Cartesian co-ordinate \u03b1 angle measured form hmax in the direction of rotation \u03b2 angle between \u2212\u2192 V and \u2212\u2192 F \u03b5 eccentricity ratio \u03b7 dynamic viscosity \u03b8 crank angle \u03bb crank length/connecting rod length \u03be , \u03bc rectangular co-ordinates \u03c6 attitude angle \u03c8 angle between connecting rod axis and cylinder liner axis \u03c9 angular velocity APPENDIX 2 Ricardo hydra gasoline engine data Engine bore 85.99 mm Engine stroke 86 mm No. of cylinders 1 No. of valves/cylinder Two inlet and two exhaust Piston rings Two compression rings and one oil control ring Connecting rod length 143.5 mm Crank radius 43 mm Stroke 4 Main bearings 1 and 2, diameter 57 mm Main bearings 1 and 2, clearance 0.03 mm Main bearing 1, length 35 mm Main bearing 2, length 32 mm Big-end bearing diameter 49 mm Big-end bearing length 20 mm Big-end bearing clearance 0.025 mm Reciprocating mass 0.64 kg Rotating mass 0.48 kg For reference lubricant viscosities Parameters SAE 0W20 SAE 5W40 Vk40 (cSt) 41.37 57.4 Vk100 (cSt) 8.07 9.9 JET487 \u00a9 IMechE 2009 Proc. IMechE Vol. 223 Part J: J. Engineering Tribology at University of Ulster Library on March 31, 2015pij.sagepub.comDownloaded from Fig. 22 Instrumented connecting rod for measuring piston assembly friction Proc. IMechE Vol. 223 Part J: J. Engineering Tribology JET487 \u00a9 IMechE 2009 at University of Ulster Library on March 31, 2015pij.sagepub.comDownloaded from"
        ]
    },
    {
        "image_filename": "designv11_12_0003159_978-1-4419-9985-6-Figure1.2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003159_978-1-4419-9985-6-Figure1.2-1.png",
        "caption": "Fig. 1.2 The transducer design with one rigid cross and one cross including springs for the connection of the MEMS structure to the transducer. The left figure shows the whole transducer and the right figure the central part where the MEMS structure is mounted",
        "texts": [
            " A silicon wafer with the surface oriented in [1 1 0] crystal direction was used, and the sensor piezoresistors were applied in the same orientation, which resulted in a high piezoresistive effect. For the low cost transducer, a sheet of steel was used and the spring system, which adapts the MEMS sensor element to the applications, was manufactured by laser cutting. A critical problem was then how to design a laser cut transducer in order to make the mounting of the MEMS sensor on the transducer as simple as possible. Figure 1.2 illustrates the solution that was found to this problem. According to Fig. 1.2 on the left the transducer is divided into two cross like shapes, where one cross is rigid and connects to the inner part of the sensor element while the other cross is provided with springs that connect to four spots on the outer part of the sensor element. Figure 1.2 shows on the right a detail of the transducer where the MEMS structure is mounted. Notice that the transducer is designed in such a way that it can be cut from a single sheet of metal. A FEM model including both the mechanical properties of the crystalline silicon MEMS structure and the electrical properties of the piezoresistors was developed. This model was used to find appropriate design parameters to obtain the targeted sensitivity without overloading the silicon structure. A maximum stress level smax of 300 MPa was adopted, which can be compared with the fracture strength of Silicon which is between 1 and 5 GPa dependent on the geometry"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002342_j.mechmachtheory.2011.06.006-Figure6-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002342_j.mechmachtheory.2011.06.006-Figure6-1.png",
        "caption": "Fig. 6. A gymnast on a trampoline.",
        "texts": [
            " ; if joint i is revolute \u0398i; p\u00f0 \u00de e \u2192 i ; if joint i is prismatic: 8>>< >>: \u00f057\u00de Now, if the ith joint is revolute and locked, then the stiction torque in the joint is determined by the expression M \u2192st \u03bci = \u03bbi\u00f0 \u00deqi =q i =q i =0 e\u2192i; \u00f058\u00de while if the ith joint is prismatic and locked, then the stiction force in the joint is determined by the expression F \u2192st \u03bc i = \u03bbi\u00f0 \u00deqi =q i =q i =0 e\u2192i: \u00f059\u00de 6. Examples 6.1. A gymnast on a trampoline Let us consider a gymnast on a trampoline as a first example of the application of the algorithm. This example has been considered in [17,39]. In [17,39], the gymnast was modelled as a planar multibody system depicted in Fig. 6. See [39] for the assumptions on which this multibody model is based. In order to illustrate the proposed method for the determination of joint reaction forces, the multibody system shown in Fig. 6 is replaced by a multibody system shown in Fig. 7. In Fig. 7, body (V0) is a fixed base body and bodies (V1) and (V2) represent fictitious bodies. Joints 1, 2, 3, and 6 are unactuated, while the driving torques in the remaining joints are given by \u2212\u03c42 e\u21924, \u03c43 e\u21925, \u03c46 e\u21927, \u03c44 e\u21928, and \u03c45 e\u21929, where \u03c4i(i=1, \u2026, 6) represent the torques of muscle forces in the gymnast's joints. Note that bodies (V6) and (V3) act to each other by the couples with torques M \u2192 3 = \u03c41 i \u2192 and M \u2192 2 = \u2212\u03c41 i \u2192 , respectively, and that the axes of joints 3 and 6 coincide"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002742_icarcv.2010.5707422-Figure7-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002742_icarcv.2010.5707422-Figure7-1.png",
        "caption": "Fig. 7. Illustration of the anti-collision condition",
        "texts": [
            " The RUTF and the potential function with removal (RUTF-RR) algorithm derives a robot to reach a goal by removing a repulsion force since the movement back to previous local minima after RUTF is from the repulsive force existence. The attractive force exertion of (10) continues until a robot arrives at a goal. When the RUTF-RR is used, the robot collision issue should be considered as well. Since the robot moves with only attractive force, it is possible for the robot to collide with an obstacle. The collision occurs more frequently when the size of a robot and/or an obstacle increases, and the distance between a goal and an obstacle becomes shorter. In order to investigate the anti-collision condition, Fig. 7(a) illustrates the placement and the size of a robot, a goal and an obstacle. Given the robot position Pr, the RUTF should be generated in the collision free area. Then, after the RUTF exertion to the collision free area, the collision is prevented even with the attraction force only. Fig. 7(a) is re-illustrated in Fig. 7(b) to find the anti-collision condition. The angles \u03b81 and \u03b82 are formulated as \u03b81 = \u03b82 = csc ( rr+ro dv ) Then, the shortest distance dw between a robot and the shaded region boundary is expressed as dw = du \u00b7 sin \u03b81. In addition, the RUTF angle \u03b8(Fc tot) is \u03b8(Fc tot) = \u03c0 \u2212 \u03b81. (11) In order to move inside the shaded region with speed sr, the robot should keep moving with the force direction \u00b1(\u03c0 \u2212 \u03b81) for \u03c4 time-instants, where \u03c4 = \u2308 dw sr \u2309 . The sign \u00b1 of the force direction is randomly assigned"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001637_s11044-010-9215-x-Figure3-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001637_s11044-010-9215-x-Figure3-1.png",
        "caption": "Fig. 3 Two-point contact configuration",
        "texts": [
            " The second point of contact might be on the wheel flange or the back of flange in the case of special track work. In order to discuss the contact geometry of the wheelset subjected to two-point contact, the contact condition associated with the second point of contact needs to be imposed [11]. To this end, surface parameters associated with the two-point contact are defined as swr3 and the coordinates used in the analysis of two-point contact are given as follows: p = [ qwT swr1T swr2T swr3T ]T (14) where swr3 is a set of surface parameters for the second point of contact of the right wheel as shown in Fig. 3. The preceding 18 parameters are determined by solving the following 18 nonlinear algebraic equations: C(i) ( qw, swr1, swr2, swr3 ) = \u23a1 \u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a3 Cwr1(qw, swr1) Cwr2(qw, swr2) Cwr3(qw, swr3) \u03c8w \u2212 \u03b4 (i) \u03c8 Rw X \u03b8w \u23a4 \u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a6 = 0 (15) It should be noted at this point that the lateral displacement Rw Y and the yaw displacement \u03c8w cannot be changed independently since the additional contact condition associated with the second point of contact (Cwr3(qw, swr3) = 0) is imposed. In the preceding equation, the yaw angle is prescribed by \u03b4 (i) \u03c8 and, therefore, the lateral displacement is determined as a result of the change in the yaw displacement when the wheelset is subjected to two-point contact. Furthermore, since the contact condition associated with the second point of contact is imposed as a non-conformal contact, the change in the location of contact points along the rail as well as the circumferential point on wheel flange can be accurately predicted for two-point contact configurations. In other words, the location of the lead/lag contact point on the flange or the back of flange in special track work can be determined using (15) as shown in Fig. 3. Note that it is assumed that both wheels are in contact with rails and loss of contact is not considered in the contact geometry analysis. However, a transition from one to two point contacts can be modeled by switching the equations (see (13) and (15)) to be solved. More details on the detection of the second point of contact are found in the literature [9]. For given lateral displacement and yaw angle of a wheelset defined with respect to the track trajectory coordinate system, the four surface parameters of right and left contact points can be obtained using the look-up contact table"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002806_s12206-012-0216-y-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002806_s12206-012-0216-y-Figure1-1.png",
        "caption": "Fig. 1. Geometry and coordinate systems of hermetic capsule.",
        "texts": [
            " The hemispherical shell and the cylindrical shell were assumed to have free boundary condition and the simply supported boundary condition, respectively, at the junction and the effects of the spherical shell was assumed to be smaller than the cylindrical shell to the free vibration of joined structure. Redekop [7] also reported a study on the free vibration of toruscylinder shell assembly using differential quadrature method. 2. Hemispherical shells A schematic diagram of a hermetic capsule is given in Fig. 1, which is divided into three subregions. 1\u03a9 and 2\u03a9 represent the hemispherical shells, and 3\u03a9 is the cylindrical sec*Corresponding author. Tel.: +82 10 7321 2589, Fax.: +82 41 862 2664 E-mail address: jinhlee@hongik.ac.kr \u2020 Recommended by Editor Yeon June Kang \u00a9 KSME & Springer 2012 tion. The subregions are connected through the junctions 1 1 3\u0393 = \u03a9 \u2229 \u03a9 and 2 2 3\u0393 = \u03a9 \u2229 \u03a9 . Assuming simple harmonic motions, the equations of motion of the axisymmetric vibration of hemispherical regions 1\u03a9 and 2\u03a9 with the effects of transverse shear and rotary inertia taken into account are derived from Soedel [8] as ( ) 2 2 2 2 2 1 1 1 tan tan 1 , C U U WU R Gh U W hU R R R \u03bd \u03bd \u03c6 \u03c6 \u03c6 \u03c6 \u03c6 \u03ba \u03c9 \u03c1 \u03c6 \u23a7 \u23ab\u239b \u239e\u2202 \u2202 \u2202\u23aa \u23aa+ \u2212 + + +\u23a8 \u23ac\u239c \u239f\u2202 \u2202 \u2202\u23aa \u23aa\u239d \u23a0\u23a9 \u23ad \u239b \u239e\u2202 + \u03a8 \u2212 + = \u2212\u239c \u239f\u2202\u239d \u23a0 (1a) ( ) 2 2 2 2 2 tan 1 tan tan 1 2 , tan Gh R Gh U U W W R C U U W hW R \u03ba \u03c6 \u03c6 \u03ba \u03c6 \u03c6 \u03c6 \u03c6 \u03c6 \u03bd \u03c9 \u03c1 \u03c6 \u03c6 \u239b \u239e\u2202\u03a8 \u03a8 +\u239c \u239f\u2202\u239d \u23a0 \u239b \u239e\u2202 \u2202 \u2202 + \u2212 \u2212 + +\u239c \u239f\u2202 \u2202 \u2202\u239d \u23a0 \u239b \u239e\u2202 \u2212 + + + = \u2212\u239c \u239f\u2202\u239d \u23a0 (1b) 2 2 2 2 3 2 1 1 tan tan 1 "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003160_1.57112-Figure3-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003160_1.57112-Figure3-1.png",
        "caption": "Fig. 3 Noncoherent initial ECAV position.",
        "texts": [
            " Theorem 3: The optimal phantom target heading for minimizing the lateral acceleration induced penalty cost function J P n i 1 vt sin t i 2 is given by t 1 2 tan 1 P n i 1 sin 2 iP n i 1 cos 2 i Proof: Consider the following minimization criterion: J Xn i 1 vt sin t i 2 v2t sin2 t Xn i 1 cos2 i sin2 i Xn i 1 sin2 i sin 2 t 2 Xn i 1 sin 2 i (30) Differentiating J partially with respect to the phantom target heading and equating to zero, yields @J @ t v2t sin 2 t Xn i 1 cos2 i sin2 i cos 2 t Xn i 1 sin 2 i 0 which on solving yields t 1 2 tan 1 P n i 1 sin 2 iP n i 1 cos 2 i (31) \u25a1 Equation (31) gives the optimal value of phantom target heading that the ECAVs should generate for the minimum control effort criterion J. C. ECAV Deceptive Formation from Initial Noncoherent Configuration The ECAVs may not necessarily be in an initial position such that all the LOS intersect at the same point as considered in Fig. 1. In such conditions the ECAVs first need to maneuver so as to satisfy the initial position requirements for generating a single coherent desired phantom target. Consider the ith Radar-ECAVpair as shown in Fig. 3 where the ECAV is off the desired phantom target LOS angle ti and is flying at an LOS angle ai with respect to the radar. The ECAV needs to correct its position error, ei, which can be approximated as ei \u2019 ri ti ai (32) Using Eq. (32) a proportional controller can be derived as aai Kei ri ti ai (33) To avoid an oscillatory response, the proportional controller command of Eq. (33) can be compensated as aai KG s ri ti ai (34) with D ow nl oa de d by U N IV E R SI T Y O F SY D N E Y o n M ay 2 0, 2 01 3 | h ttp :// ar c"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001325_elan.200804415-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001325_elan.200804415-Figure2-1.png",
        "caption": "Fig. 2. The complete fabricated electrochemical microdevice.",
        "texts": [
            " 1, each substrate comprised of two flow cell base plates which was sawn into two separate parts by trimming the runner (polymer flow path) surrounding the plates. Each part was then overmolded with Zeonor forming a flow channel of 100 mm depth. In parallel to that, the top plate containing the fluidic connectors was also injection molded from Zeonor. An energy-directing ridge of 0.25 mm deep was molded around the channel for the two halves to be sealed using an ultrasonic welder (Model 2000aed, Branson Ultrasonic Corp., Slough, UK). A picture of the complete device with the integrated electrodes is shown in Figure 2. All electrochemical and ECL measurements were carried out using an Autolab PGSTS 12 potentiostat (Eco Chemie BV, Netherlands) interfaced to a PC for data acquisition via the General Purpose Electrochemical System Software package (GPES version 4.9). The ECL emission was detected using a miniaturized photomultiplier tube (PMT) (H5784-04, HAMAMATSU Photonics, Hertfordshire, UK) and recorded on a PC using an in-house program written with LabVIEW software version 8.2 (National Instruments, Austin, TX, USA)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003520_j.euromechsol.2011.11.003-Figure7-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003520_j.euromechsol.2011.11.003-Figure7-1.png",
        "caption": "Fig. 7. Maximally regular T2R1-type parallel manipulator with bifurcated planarspatial motion of the moving platform: constraint singularity (a), branch with planar motion (b) and branch with spatial motion (c); limb topology PkRkRkRtRPaPattPtkPtRtRSRkR.",
        "texts": [],
        "surrounding_texts": [
            "The term of constraint singularity (CS) has been recently coined (Zlatanov et al., 2002) to characterize the configuration of lower mobility parallel manipulators in which both the connectivity of the moving platform and the mobility of the parallel mechanism increase their instantaneous values. From a constraint singularity, the mechanism can get out with or without branching (Gogu, 2008b). When branching occurs in a constraint singularity, the mechanism can reach different configurations, called branches, and have different independent motions of the moving platform. In this case, the constraint singularity is also called branching or bifurcation singularity. A branch refers to the free-of-singularity configurations of the mechanism inwhich each structural parameter keeps its value. For this reason, this value is called global of full-cycle value for a branch. Two types of branching in constraint singularity (BCS) have been defined in Gogu (2008b). Branching of type BCS1 occurs when a parallel mechanism F ) G1 . Gj . Gk gets out from a constraint singularity in different branches characterized by the same mobility and the same connectivity of the moving platform but with different bases of the vector space of relative velocities between the moving and fixed platforms. Branching of type BCS2 occurs when a parallel mechanism F ) G1 . Gj . Gk gets out from a constraint singularity in different branches characterized by distinct values of mobility and connectivity of the moving platform. The parallel manipulators presented in this paper have constraint singularities with branching of type BCS1. Branching of type BCS2 occurs in kinematotropic mechanisms (Gogu, 2009c). The finite displacements and the velocities in the actuated joints are denoted by qi and _qi, the linear velocities of the characteristic point H of the moving platform, by v1 \u00bc _x; v2 \u00bc _y and the angular velocity of the moving platform by u \u00bc ua \u00bc _a or u \u00bc ud \u00bc _d: In both branches, the moving platform undergoes two planar translations and one rotation but the rotation axis is different in the two branches. In the first case, the rotation velocity u \u00bc ua \u00bc _a is parallel to the x-axis and the moving platform undergoes a spatial motion. In the second case, the rotation velocity u \u00bc ud \u00bc _d is parallel to the z-axis and the moving platform undergoes a planar motion. With these notations, the linear mapping between the actuated joint velocity space and the end-effector velocity space for a branch the T2R1-type PMs is defined by: 2 4 v1 v2 u 3 5 \u00bc \u00bdJ 2 4 _q1 _q2 _q3 3 5 (8) where J is the Jacobian matrix. In the constraint singularity, the T2R1-type parallel manipulators with bifurcated motions has instantaneously iM \u00bc iSF \u00bc 4 and (iRF)\u00bc(v1, v2, ua, ud). The bifurcation in this constraint singularity can be used to change motion type of the moving platform. In the two distinct branches, the parallel mechanism is characterized by M\u00bc SF \u00bc 3 and (RF)\u00bc(v1, v2,ua) or (RF)\u00bc(v1, v2,ud). In the constraint singularity, both connectivity of moving platform and mobility of parallel manipulator increase their instantaneous values. More details about the behaviour of parallel manipulators in a constraint singularity and the associated structural parameters can be found in Gogu (2009c)."
        ]
    },
    {
        "image_filename": "designv11_12_0000387_s0022112080001784-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000387_s0022112080001784-Figure2-1.png",
        "caption": "FIGURE 2. The inaccessible regions.",
        "texts": [
            " From physical intuition and from equation (3a), which gives a = - 8 at a = 0, it is clear that there is a possibility with this shape that in time a(t) = R(0, t ) will decrease to zero and thus the drop will burst into two. At a = 0 we have 1 = ( - b/c)l/(m-n) and then from the volume normalization (m-n)/(%n+l) ( - b)(Zm+1)/(2%+1). 1 3(m - n)2 ( 2 m + l ) ( m + n + 1)(2n+ 1) c = [ For a to be positive b must be greater than the value which gives the equality in (5). The inaccessible region in the bc plane of nonsensical shapes is thus the shaded region in the second quadrant of figure 2. To find the motion near this forbidden region we note that any straight line from the origin meets the curve (5) once in the quadrant, because the index of the power laws exceeds unity (m > n) . By the global stability result (4) the solution trajectories move across this line away from the c axis, and so must eventually meet the curve. (They cannot escape to the origin because b and c grow when they are small.) Hence any drop shape with b < 0 at the initial instant must burst. When b > 0 and c c 0, there is a possibility that the shape of the drop does not close, the limiting case having a cusped end"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003411_elan.201100616-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003411_elan.201100616-Figure1-1.png",
        "caption": "Fig. 1. (A) The detailed scheme of BDD disk electrode used in wall-jet arrangement: electrode body made of Teflon (1), stainless steel (2), screw attachment made of Teflon (3), small metal spring (4), brassy sheet (5), boron-doped diamond thin film electrode on silica support (6), Viton gasket (7) and access for solution (8). (B) Amperometric thin-layer detection cell: screw clamp (1), Kel-F body, bottom piece (2), copper contact (3), boron-doped diamond thin film electrode on silica support, cell volume ~10 mL (4), Kel-F body, top piece (5), counter electrode (6), reference electrode (7), Viton gasket with an ellipsoid groove defining the flow channel (8). Figure (A) adapted with permission from [2]. Copyright (2007) Wiley-VCH. Figure (B) adapted with permission from [14]. Copyright (1997) American Chemical Society.",
        "texts": [
            "0 : acetonitrile :methanol (40 :30 :30; v/v/v) was degassed by ultrasonication using PS 02000A ultrasonic bath (Powersonic, USA) followed by passing helium for 5 min. The measurements were carried out at laboratory temperature. The spectrophotometric detector was set at the wavelength l of 290 nm. The amperometric detector with three electrode arrangement was controlled by potentiostat ADLC 2 (Laboratorn pr\u030c stroje, Prague, CZ). The wall-jet arrangement was realized using laboratory-made BDD disk electrode [2] of an active surface of 12.6 mm2 (i.e. diameter 4.0 mm; Figure 1A), saturated calomel reference electrode (SCE) and platinum indicator electrode (both Monokrystaly, Turnov, CZ). The electrode surface\u2013capillary outlet distance for the wall-jet arrangement was kept at the distance of 0.5 mm; the jet diameter was 0.15 mm. The thin layer detection cell is depicted at Figure 1B; it was described in detail earlier [14]. Our arrangement differed in the Ag/AgCl/3 molL 1 KCl reference electrode (SSCE) of the type 66-EE009 (Cypress Systems, Chelmsword, MA, USA). Further, a 0.1 cm thick Viton (fluoropolymer elastomer, G schu, CZ) gasket separated the surface of the working electrode from the top piece of the cell. An ellipsoid groove (a=3.2 mm, b=1.3 mm) was cut in the gasket and defined the flow channel. Assuming a 25% compression of the gasket when the two pieces of the cell were clamped together, the cell volume was estimated to be ~10 mL",
            " The adsorbed aminobiphenyls were eluted by 1 mL of methanole without the use of vacuum and 20 mL of the eluate was directly injected into the HPLC system. The separation of 2-AB, 3-AB and 4-AB using the optimum separation and detection conditions is for all detection modes, i.e. , electrochemical detection with BDD electrode in thin-layer and wall-jet arrangements and for the spectrophotometric detection depicted at Figure 2. The electrode surface\u2013capillary outlet distance for the wall-jet arrangement was kept at the minimum distance of 0.5 mm given by the construction of the disk electrode (Figure 1A). Higher distances led firstly to 20% increase of peak height and area with maxima at the distance of about 2.0 mm followed by a slight continuous drop. Further, for electrochemical detection the influence of the Edet imposed on BDD electrode on its signal was investigated. The hydrodynamic voltammograms were recorded point by point by making 100 mV changes in the applied potential over the +600 mV to +2250 mV (vs. SHE) range, while the height of the peak Ip of the studied ABs simultaneously with the absolute value of the background current Ib and peak-to-peak noise In were recorded",
            " Further the state of the electrode surface influencing the kinetic of the electron transfer or usage of peak height instead of the limiting current iL in Equation 1 is often questionable [50]. Higher values of x than these derived theoretically for laminar flow (Table 2) may be in general caused by local turbulences [50,55] \u2013 in our case turbulences caused by (i) the nonplanar placement of BDD indicator electrode in both arrangements when considering other parts of the electrode body (for the wall-jet arrangement, the Si wafer covered by BDD is in fact placed in a cavern with depth of 0.5 mm defined by the width of the Teflon electrode body and the viton gasket (Figure 1A); similarly, the width of the gasket defines the depth of the ellipsoid groove housing the BDD electrode in the thin-layer arrangement (Figure 1B)), and (ii) turbulences caused by the roughness of the BDD surface itself. The importance of the latter effect increases with increasing flow rate and is probably more relevant for the wall-jet arrangement, where the intact jet impinges directly the electrode surface due to the short electrode surface\u2013capillary outlet distance of 0.5 mm leading in fact to a wall-jet thin layer cell and causing higher noise of the wall jet compared to the thin-layer arrangement of the detection cell. Parameters of calibration dependencies measured under optimized conditions are listed in Table 3"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002607_j.asoc.2012.07.032-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002607_j.asoc.2012.07.032-Figure1-1.png",
        "caption": "Fig. 1. The position of the fighter in a 3D environment.",
        "texts": [
            " This paper is divided into five sections, the first section is the introduction, the second contains the flight dynamic equation, the design and analysis of the fighter\u2019s variable feedback gain controller, the third is the structure of the PSO-based VFGC algorithm, the fourth contains the simulation results, and the fifth contains the conclusion. 2. Design of variable feedback gain controller for AFT problems 2.1. Flight dynamic equation of fighter The posture and position of the fighter in a three-dimensional (3D) environment is shown in Fig. 1. The fighter\u2019s dynamic behaviors are listed below: (a) x\u0307 = V cos sin (b) y\u0307 = V cos cos (c) z\u0307 = V sin (1) where x, y and z are the fighter\u2019s positions in the inertia coordinate system with units in meters, and there first order differential are denoted as x\u0307, y\u0307 and z\u0307. The path angle with unit in degrees represents the angle between the velocity vector and the horizon. The heading angle with units in degrees represents the angle between the velocity vector and the local north. The fighter\u2019s velocity V is with unit in m/s"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001135_s11661-009-9870-9-Figure3-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001135_s11661-009-9870-9-Figure3-1.png",
        "caption": "Fig. 3\u2014Finite element model for the LPD process.",
        "texts": [
            " For this purpose, four cases with the same deposition area are studied for different deposition patterns, as shown in Figure 1. In the LPD process, as shown in Figure 2, the deposition region can be divided into two regions: the \u2018\u2018leading half,\u2019\u2019 which moves in front of the laser beam and melts the substrate; and the \u2018\u2018trailing half,\u2019\u2019 which follows the leading half and contains the molten material. This concept is used for modeling the deposition process by the finite element method, as shown in Figure 3. In this figure, four regions are distinguished: the white elements represent the substrate or the layer underneath, the light-gray elements represent the deposited elements that were activated in the previous timesteps, the dark-gray elements are the activated elements Table II. Assumed Parameters Values in the Modeling Equations [11], [12], [14], [16], and [18] Parameter Description Value a surface absorption 0.2 e surface emissivity 0.6 Ac1 austenization start temperature ( C) 850 Ac3 austenization end temperature ( C) 925 Ca ferrite carbon content 0"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003859_978-3-642-36365-8_6-Figure12-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003859_978-3-642-36365-8_6-Figure12-1.png",
        "caption": "Fig. 12. Buck converter",
        "texts": [],
        "surrounding_texts": [
            "As noted above the highest current-draw of a wireless sensor node will often occur at start-up, which can cause significant negative effects on the output voltage of the MFC system. In order to avoid these issues, when a buffering capacitor of sufficient capacity is used, this can be pre-charged and used to start the node up independent of the MFCs. Once this has occurred the MFC system only needs to supply enough energy to operate during transmission, which is often significantly less than start-up."
        ]
    },
    {
        "image_filename": "designv11_12_0003903_tdei.2015.005053-Figure20-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003903_tdei.2015.005053-Figure20-1.png",
        "caption": "Figure 20. The gap between semi-conductive rubber sleeve and zinc sleeve.",
        "texts": [
            " The surface electric field intensity can decrease to 4.8 kV/cm when installing semi-conductive rubber sleeve. The nephograms of electric field intensity shown in Figure 19 also illustrate that, the installation of conductor rubber sleeve has better electric field distribution characteristic than the other two kinds of sleeves. If the bonding property between organic material sleeve and zinc sleeve is bad in its production process or the organic material aging is generated by high electric field intensity, a gap can be generated, which is shown in Figure 20. In terms of semi-conductor rubber sleeve, the electric field intensity of samples with 0.5, 1.0 and 1.5 mm width gaps are calculated, respectively. The results are shown in Figure 21. Figure 21 shows that, the electric field intensity near pin reduces with the increase of gap size, which does not mean the larger the gap size the better. Because the gap converge water in it under the condition of raining, condensation and heavy fog, which will corrode the pin and zinc sleeve inside the semi-conductive rubber sleeve"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002116_elan.200804439-Figure4-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002116_elan.200804439-Figure4-1.png",
        "caption": "Fig. 4. Stripping voltammetry of SPE/DDAB-AuNP. Cathodic peak height dependence on oxidation conditions (potential and time), scanning from 0.6 to 0.6 V, 50 mV/s, PBS, pH 7.4",
        "texts": [
            " At low scan rates (from 5 to 100 mV/s) the peak current increased linearly with the scan rate (vE), but not with vE 1/2 (Ip (A) vE (V/s), A\u00bc 7.82 10 7, B\u00bc 8.51 10 4, R\u00bc 0.999), indicating thin layer electrochemistry [35]. The peak potential is linearly shifted in the more negative range with lg vE (Ep (V) lg vE (V/s), A\u00bc 0.06, B\u00bc 0.11, R\u00bc 0.998), for the 3e irreversible process the slope of this curve is 2.3RT/anF\u00bc 0.059/an, so that a can be estimated as 0.18 [36]. As expected, the reduction peak area was found to correspond to AuNP oxidation degree (Fig. 4). The reproducibility of SPE/DDAB-AuNP was assessed by analyzing signals of 6 different sensors and RSD was 12%. AuNP signal is internal characteristic of the sensor and was changed with deposition of biolayers because of differences in electron transfer efficacy and availability of buffer oxygen. The array approach was suggested for evaluation of changes of gold nanoparticles status during biorecognition events. Detection of binding between anti-HMb immobilized onto AuNP modified electrode and HMb was performed by the measuring the difference between AuNP cathodic peak areas in PBS for SPE/DDAB-AuNP/antiHMb and SPE/DDAB-AuNP/anti-HMb/HMb"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002748_s12206-012-0408-5-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002748_s12206-012-0408-5-Figure1-1.png",
        "caption": "Fig. 1. The proposed series-parallel manipulator.",
        "texts": [
            " On the other hand, it is interesting to note that a seriesparallel manipulator with different performance can be obtained by changing the sequence of its kinematic pairs. In that way Lu et al. [18] approached the kinematics and statics of an inversion of the 2(3-RPS) robot. This work introduces a new six-degrees-of-freedom nonredundant spatial series-parallel manipulator, in other words a robot where the number of degrees of freedom is equal to the six-dimensional task space. Naturally the number of available motors or generalized coordinates is also equal to the degrees of freedom of the proposed robot. The proposed robot, see Fig. 1, consists of a lower parallel manipulator (LPM) and an upper parallel manipulator (UPM) connected to each other through \u2018compound joints\u2019 attached to a coupler platform (body 2). A compound joint results from the combined action of a revolute joint covering a spherical joint where the revolute axis intersects the center of the spherical joint. The LPM is a 3-PPS parallel manipulator where the kinematic pairs connecting the limbs to the fixed platform (body *Corresponding author. Tel.: +52 461 6117575, Fax",
            " The forward position analysis of the proposed mechanism is formulated as follows: given the generalized coordinates { , }ii q q , in the remainder of the contribution i = 1, 2, 3; compute the pose of the output platform with respect to the fixed platform. A strategy to approach this analysis consists of determining firstly the coordinates of the centers Ci of the spherical joints attached at the coupler platform. To this end, consider that the position vectors Ci of such points can be obtained, see Fig. 1, as \u02c6 i i i i d= + +C A h u (1) where \u02c6\u02c6sin( ) sin( )i i i i r q i r q k= = +rA is the position vector of the nominal point Ai of the i-th lower link, \u02c6hj=h and ) \u02c6\u02c6 \u02c6\u02c6 cos( )sin( sin( ) cos( )cos( )i i i \u0398 q i \u0398 j \u0398 q k\u2212= + \u2212u is the unit vector along the i-th PS kinematic chain. Clearly, nine linear equations in twelve unknowns can be obtained upon expressions (1). In order to complete the number of equations, consider that the equilateral triangle \u0394C1C2C3 brings three nonlinear equations as follows: 2( ) ( ) , 1, 2, 3 mod(3)i j i j e i j\u2022\u2212 \u2212 = =C C C C (2) where the dot (\u2022) denotes the inner product of the usual threedimensional vectorial algebra",
            " Screw theory has been proved to be an excellent resource to investigate the kinematics of the 3-RPS parallel manipulator, see for instance [22-26], and therefore it is chosen as the mathematical tool to approach the velocity, acceleration and singularity analyses of the series-parallel manipulator. Let 1 1 1;[ ] ( 1,2)k k k k k k T O O k+ + += =\u03c9V v be the velocity state of the platform k+1 with respect to the platform k, where 1k k+\u03c9 and 1k k O +v are the angular and linear velocities of the platform k+1 taking point O, which is instantaneously coincident with the origin O1 of the fixed reference frame XYZ, as the reference pole. Moreover, the six-dimensional vector 1k k O +V can be written in screw form, the infinitesimal screws are depicted in Fig. 1, as 1 1 1k k k k k k i i O\u2126+ + +=J V (12) where 1k k i +J and 1k k i\u2126 + are, respectively, the screwcoordinate Jacobian matrix and the joint-rate velocity matrix of the indicated limb. For the LPM k= 1 and 1 2 0 1 1 2 5 6[ $ $ $ ]i i i i =J while 1 2 0 1 1 2 5 6[ ] .Ti i i i\u2126 \u03c9 \u03c9 \u03c9= Similarly, for the UPM k = 2 and 0 1 1 2 5 62 3 [ $ $ $ ]i i i i =J whereas 2 3 0 1 1 2 5 6 [ ] .i i i T i\u2126 \u03c9 \u03c9 \u03c9= It should be noted that the joint rates 0 1 i i q\u03c9 = and 2 3 i i q\u03c9 = have the privilege to be chosen as the generalized speeds of the robot"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002127_j.pocean.2010.01.002-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002127_j.pocean.2010.01.002-Figure1-1.png",
        "caption": "Fig. 1. Shapes of the perceptive volume considered in this work. In (a) we have the simple spherical reference model with only one parameter R, while (b) shows a more general conical volume, with two parameters R and h.",
        "texts": [
            " Consequences of finite capture probabilities will not be discussed here. The effects of finite inertia and finite sizes have been discussed elsewhere for simple cases without predator motion (P\u00e9cseli and Trulsen, 2007, 2008). Model (1) also ignores the time spent by the predator for handling prey (Visser, 2007). The analysis giving (1) is essentially based on dimensional arguments (Buckingham, 1914), and the resulting flux-scaling with n0 e1/3R7/3 is universal and does not depend on the shape of the encounter region, see Fig. 1. The surface of the encounter region can be spherical or conical as long as the surface changes self-similarly with a certain length scale, R. A conical surface will serve just as well as the spherical one, only with a new coefficient, replacing CM. There are no other adjustable parameters. By following conical surfaces in the flow, we have empirically determined these coefficients (Mann et al., 2006; P\u00e9cseli and Trulsen, 2007). The results are best accounted for by keeping CM for the spherical surface and introduce a correction factor in the range {0:1}",
            " More generally it can be argued that since the suggested analytical results (1) and (2) are well satisfied for a wide range of parameters R and h, the results can be applied with confidence also in other cases. The expression (2) is then amenable for a test with the given numerical database (Biferale et al., 2005, 2006; Boffetta et al., 2006). We interpolate the flow field from the database of point particle velocities, and let the centre of the virtual detective sphere (or the apex for a conical volume, see Fig. 1) move with a velocity Vc relative to the local fluid volume. The flux over the surface of the perceptive volume can be determined for different velocities Vc and radii R. Here we use R/g = 25, 50, and 75. All R-values mentioned are in the universal subrange of the turbulence. The analysis is carried out for two different simulations, as indicated by full and dashed lines in Fig. 4. The two simulations differ by the value chosen for the kinematic viscosity and the energy dissipation rate, and also the numerical resolution is different by a factor of two",
            " The results are illustrated by the time-integrated capture rates, where the integration is over one pause-travel period. The results shown in Fig. 8a and b are normalized by analytical results obtained via the expression (6). We find that the jump length is an important parameter. In order to have a high probability of entering fluid volumes that have not been depleted before, it seems necessary to have a jump length exceeding 3R in terms of the radius R in the sphere of interception. This result remains qualitatively accurate also for hemispherical or conical surfaces, where R is defined in Fig. 1b for these cases. In the limit of large jump lengths we find good agreement between analytical estimates and our numerical results. The analysis takes into account the time variation of the turbulent encounter rate. We have obtained simple analytical expressions (9) and (10) for the relative importance of the direct and the integrated captures, these expressions being valid for different limiting cases. Our reference case is for simple spherical detection volumes, but our analysis contains an analytical prescription for obtaining the correction for finite opening angles in a conical volume"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001936_iembs.2010.5627660-Figure3-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001936_iembs.2010.5627660-Figure3-1.png",
        "caption": "Fig. 3. Calibration of robot to laserscanner. (a) Instead of a direct calibration, a reference image of the calibration phantom is used for calibration. (b) When we lay the reference image into the laserscanner coordinate system, we obtain the needed transform RTL.",
        "texts": [
            " While the algorithms proposed by Tsai and Lenz [9] or Daniilidis [10] provide accurate hand-eye-calibration, the results are not necessarily optimal: the algorithms always produce orthonormal matrices. Since we may safely assume that no tracking camera is calibrated perfectly, i.e., all such systems suffer from distortion, an orthonormal calibration matrix may not be optimal. Furthermore, in the presented case of a laserscanner, the ICP method brings additional distortion to the tracked data. A na\u0131\u0308ve approach is to look at the general relation: RTE ETM = RTMref Mref TM . (1) Equation 1 is illustrated in Figure 3(a). Here, the matrices ETM , the transform from the robot\u2019s effector to the calibration phantom, and RTMref , the transform from the robot\u2019s base to the reference image, are unknown. Instead of the calibration of robot to reference image, we are interested in the calibration from robot to laserscanner, RTL, because we want to track the head position in laserscanner coordinates. So that we can easily transform them into the robot coordinate system with the calculated transform. When we lay the reference image in the laserscanner coordinate system, we have the same behavior, after using ICP, of the laserscanner like a standard tracking system for the robot hand-eye-calibration. This means, the presented method results in the needed transform RTL. This is illustrated in Figure 3(b). Therefore, we define the origin and axes manually in the reference image by selecting three points (orign, x-axis, y-axis) that span the coordinate system. The registration of the coil C has to be done the same way presented above. But for the TMS coil there are different ways to obtain a high quality reference image. One could make a scan with a medical imaging technique like CT or one uses the specific CAD data that is provided by the coil manufacturer. For simplicity, we use a high resolution laser scan of the coil"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002346_jf60158a001-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002346_jf60158a001-Figure2-1.png",
        "caption": "Figure 2. Ultraviolet absorption spectra of MA in diethyl ether-cyclohexane mixtures Concentration of M A 4.5 X 10-j M",
        "texts": [
            " Thus, the prsviously reported high molar absorptivity may be mainly attributable to the aggregation of M A (Kwon and Olcott: 1966b) caused by the high concentration and high temperature employed during hydrolysis of M A acetal. When this aggregated M A solution was diluted, the extrapolated absorbance at zero concentration was always higher than zero, suggesting some irreversible modifications. Ultraviolet Absorption Spectra of MA in Organic Solvents. In diethyl ether, two bands at 234 and 271 mp are observed, and the absorbance at 234 mp is about three times higher than 271 mp (Figure 2). Furthermore, with increasing cyclohexane content in diethyl ether-cyclohexane mixtures, the absorbance at 234 mp gradually diminishes while that at 271 mp increases, and at 99% cyclohexane essentially only the latter band remains. The spectra in such organic solvents are very different from those in aqueous solution, indicating that the solvent has very striking effects on position and intensity of the band. The possibility that these effects are due to molecular aggregation was eliminated by dilution studies",
            " x* band observed in aqueous solution; and this is the only other reasonable conformation which would give absorption in this wavelength region. The apparent molar absorptivities of P-dicarbonyl compounds are linear functions of the enol content in solvents, and the absorption characteristics are apparently almost independent of solvent (Hammond et al., 1959). If one assumes the molar absorptivities of the cis and trans enols are similar to those of the chelated enol and the enolate anion, respectively, from the geometry considerations, then the calculated M A concentration from the spectra (Figure 2) gradually decreases with increasing cyclohexane content in the solvent mixtures, while the total MA concentration (4.5 x lO--jM) determined by the TBA reaction is constant. Furthermore, the solvent-dependent spectral changes are reversible, so there is no irreversible loss of MA by changing proportions in the solvent mixtures. These observations suggest that in cyclohexane the enols are converted to the diketo form (I), which is not expected to absorb in this ultraviolet region. Surprisingly, the calculated enolic content in dichloromethane was less than 2 x of the total MA, and thus more than 9 8 z was present as the diketo form"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003142_1.3616922-Figure20-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003142_1.3616922-Figure20-1.png",
        "caption": "Fig. 20 Profile radii of curvature at pitch point, P, and at point of load concentration K '",
        "texts": [
            "2Arg + 4.0iVP N0 - NP (22a) (22 b) distance along line of action in mean normal section from gear addendum circle (outside circle) to pitch point number of pinion teeth number of gear teeth \u00b1 FZr^ _ Apf l (22) On curved-tooth bevel and hypoid gears use the plus ( + ) sign for the concave side of the pinion tooth (convex side of the gear tooth) and the minus ( \u2014 ) sign for the convex side of the pinion tooth (concave side of the gear tooth) in equation (22). Relative Radius of Profile Curvature From Fig. 20 it can be seen that at point K' the relative radius of profile curvature po of the two contacting tooth profiles is Po P1P2 Pi + Pi (23) 1 2 0 / A P R I L 1 9 6 7 Transactions of the AS ME Downloaded From: http://tribology.asmedigitalcollection.asme.org/ on 01/29/2016 Terms of Use: http://www.asme.org/about-asme/terms-of-use Pi = pP + zo radius of profile curvature on the pinion p2 ~ PQ \u2014 zo radius of profile curvature on the gear pP = radius of profile curvature at pitch point on pinion in the mean normal section p0 = radius of profile curvature at pitch point on gear in the mean normal section (23 Cont"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000762_s0003-2670(01)82507-1-Figure4-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000762_s0003-2670(01)82507-1-Figure4-1.png",
        "caption": "Fig. 4. Dependence of the limiting cell potential on the concentrat ion of NEBA.CIOa in the membrane. solution. Curves (a) and (b) refer to the NEBA. ClO,,/fl,fl'-dichlorodiethyl ether and NEBA.CIO4./I,2-. dichlorobenzene systems, respectively. The dotted line (c) represents the limiting potent ial for the sol idstate electrode after appropriate adjustments to the observed potentials. The agreement between this line and the limiting potential for the saturated NEBA. CIO4/fl,~'-dichlorodiethyl ether electrode indicates that the electrodes show the same response range.",
        "texts": [
            " F igu re 4 shows the var ia t ion in l imi t ing cell po ten t i a l (which defines the lowest m e a s u r e a b l e act iv i ty level) as a func t ion of the c o n c e n t r a t i o n o f N E B A - C I 0 4 in the m e m b r a n e . Obse rva t i ons for bo th solvents are i nc luded ; an increase in po ten t i a l c o r r e s p o n d s wi th an increase in measu r ing range. M a r k e d l y different behav iou r was s h o w n b y the two-e lec t rode systems. B o t h showed the expec ted i m p r o v e m e n t in measu r ing range d u r i n g the initial s tages o f d i lu t ion f rom sa tu r a t ed so lu t ions ( the po in t s for the h ighes t c o n c e n t r a t i o n s in each case in Fig. 4 r ep resen t sa tu ra ted so lu t ions) bu t whereas an u p p e r l imit was a p p r o a c h e d by the 1 ,2 -d ich lorobenzene system, the f l , f l ' -d ichlorodie thyl e the r e lec t rode exh ib i ted a r ap id d e t e r i o r a t i o n in response range for m e m b r a n e c o n c e n t r a t i o n s o f N E B A . CIO4 be low ca. 0.1 g 1- ~. The b e h a v i o u r of t he la t ter is cons i s t en t wi th the b r e a k d o w n of D o n n a n exclusion, w h e r e u p o n co- ion t r a n s p o r t ac ross the m e m b r a n e is p e r m i t t e d ~ 2"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003903_tdei.2015.005053-Figure13-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003903_tdei.2015.005053-Figure13-1.png",
        "caption": "Figure 13. Test results of samples with different kinds of organic sleeves.",
        "texts": [
            " The test electric charge quantity is set as 45000 C, according to the maximum average annual amount of corrosion charge (1500 C/year) obtained from Chusui transmission line and the 30-year service life of porcelain insulator. After tests, 1000 kg of cement block has been hung on these insulators for six months, as shown in Figure 11. Six months later, the tested insulators are carried out tensile load test in accordance with Chinese National Standard GB/T19443 [23]. The tensile failure load and pins are shown in Figure 12 and Figure 13 respectively. Note: A stands for without organic sleeve insulators, B stands for Flush type sleeve insulators, C stands for whole package type sleeve insulators, D stands for half packed type sleeve insulators. Every value in Figure 12 is the average of 3 pieces of insulators' test results. Figure 13b shows that, the installation of half package type organic material sleeve has altered the corrosion area of zinc sleeve. The original corrosion part cement-zinc sleeve interface bears radial stress, present corrosion area is the exposed part of zinc sleeve that withstands no hoop stress. Thus, the expansion of corrosion by-products can not pose adverse impact on the mechanical strength of insulator. The test results illustrated in Figure 12 have verified the effectiveness of half package type organic material sleeve. However, the corrosion parts are pins under installing flush type sleeve as well as whole package type sleeve, which can reduce the cross-section of pins and decrease the mechanical strength. Moreover, the test results illustrated in Figure 13b indicate that the part of zinc sleeve between the two lines is packed by organic sleeve and cannot be used to protection pin as sacrificial electrode. Namely, the installation of half package type organic sleeve can reduce the effective size of zinc sleeve and lead to the pin corrosion, as the blue circle area shown in Figure 13b. In order to solve this problem, It is suggested to increase 1-1.5cm to the height of exposed part of zinc sleeve based on the existing height. For organic materials, if its surface electric field intensity is higher than 5 kV/cm, surface discharge phenomena occur. It poses adverse impact on the mechanical and electrical characteristics of organic material and accelerates its aging. So, the calculation of electric field intensity is significant to the material selectionand structure optimize design of organic sleeve"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003142_1.3616922-Figure8-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003142_1.3616922-Figure8-1.png",
        "caption": "Fig. 8 Mean normal section of a spiral bevel gear",
        "texts": [],
        "surrounding_texts": [
            "Then the maximum temperatures of each surface are identical and equal to\n\\Ztt qoHV, \u2022.Vd\nAT = C\\ -x/wi + C2 s/vi\n(15)\nV, = !)! \u2014 t>2 Y / I ' I 2 + I>22 \u2014 2VIV2 COS R (16)\nwhere r is the angle between the velocity vectors Vi and t\u00bb2. Fig. 7 shows the distances di and cfe through which the area of contact will move over a particular point on each body. The total fractional energy being dissipated as heat in the contact area is\n/\nAc rA.\n0 dAc = I W>Mc (in-lb/sec) (17)\nAgain, assuming that the division of flux between the two bodies is\n(18)\n02 = (19)\nThen the approximate maximum temperature rise at the surface is\n\\Ar qonv. AT = (20)\nIn this formula the absolute magnitude of all velocities should be used. It is assumed that the cylinders are cooled to the initial ambient temperature between successive contacts (one revolution of each cylinder). Time variations in pressure, coefficient of friction, and velocity have been neglected.\nContact Velocities in Opposite Direction If the two contacting cylinders rotate in the same sense, then the velocities of the point of contact relative to each cylinder are oppositely directed. The temperature distribution is no longer that of Fig. 4. Methods of superposition at the surface of contact cannot be used because the trailing side of the contacting surface of one member preheats the leaving surface of the other. For the case of two cylinders having sliding velocities of equal magnitude and opposite direction, only 80 percent of the input energy can be accounted for by superposition at the surface. Graphically the maximum temperature can be shown to be the same as that given by assuming a flux distribution according to equations (13) and (14) and a temperature rise according to equation (15), if the energy equation below the surface of contact is plotted.\nTemperature Rise in Most General Case Fig. 6 shows the contact area between two bodies which are theoretically in contact at point P, but which are actually in contact over the shaded elliptical area, fi is the velocity of the point of contact P on body 1; Vj is the velocity of the point of contact on body 2. The absolute relative sliding between the two bodies\nIn the foregoing formula the absolute magnitude of all velocities should be used. It can be seen that this formula will reduce down to the Blok equation [3, 4] for spur gears.\nAmbient Temperature Care has been taken to express the temperature rise at points of contact in terms of a rise above ambient. In this case the ambient temperature 1\\ is the temperature of the surface just prior to the contact. This is the gear blank temperature. It may be argued that although the time between successive contacts at a point is short, this period of rotation is long compared with the actual time of contact. Hence it is reasonable to assume that with proper forced lubrication the ambient temperature at the start of contact will be only slightly higher than the temperature of the oil bath.\nApplication of the Foregoing Equation to Gear Teeth In order to apply equation (20) to determine the temperature rise on a gear pair it is first necessary to evaluate the quantities in this equation. At first glance it would appear that this presents no problem. However, on accurately made gear teeth one should expect load sharing between teeth and therefore one must ask what value to use for surface pressure. Velocities and Vi as well as the sliding velocity v, can all be determined with reasonable accuracy, but the question arises as to where on the tooth surface should the analysis be made. The distances di and ch through which the point of contact moves are dependent upon certain assumptions concerning the instantaneous tooth contact. And finally the coefficient of friction is a very important variable which must be accurately evaluated if one wishes to apply this formula with any degree of certainty.\nThe method of treating gear teeth as cylinders and of calculating the contact pressure on a rectangular band at the pitch line, which is commonly done with spur gears, does not hold for bevel and hypoid gears, since the tooth surfaces are modified (mismatched) in both the lengthwise (longitudinal) and profilewise (lateral) directions. In order to locate approximately the position of load application, the sharing of the load, and the shape of the instantaneous contact pattern, the following assumptions are made:\n1 The surface of action is taken as a plane and the lines of contact in this plane will be straight.\nJournal of Lubrication Technology A P R I L 1 9 6 7 / 1 1 7\nDownloaded From: http://tribology.asmedigitalcollection.asme.org/ on 01/29/2016 Terms of Use: http://www.asme.org/about-asme/terms-of-use",
            "NORMAL SECTION OF SPIRAL TOOTH\n2 The contact stresses are considered to follow the theory arrived at by Hertz for a semi-infinite solid and, therefore, the teeth which theoretically would have line contact if there were no elastic deformation and no tooth bearing modification, will, because of deformation and tooth bearing modification, have an area of contact in the form of an elongated ellipse.\n3 The teeth under maximum load will have actual contact in the plane of action only within an ellipse tangent to the four sides of the zone of action owing to their being modified (mismatched) both longitudinally and laterally.\n4 The stresses will not exceed the elastic limit of the material.\nIn a previous paper by the author [15] it was shown that when load sharing and deflection are not taken into account the instantaneous ellipses of contact will frequently spill over the edges of the tooth making a mathematical determination of contact pressures highly involved and of questionable accuracy. By the foregoing assumptions a mathematical simplification is possible, which makes the solution of the problem less arduous. Since the basic approach to load sharing and the determination of contact pressure at any point of the tooth surface has been covered in another paper by the author [16], the subject will be reviewed only briefly here.\nJ\nUsually the theoretical contact between a pair of gear teeth is a line. However, when the tooth surfaces are modified, the line becomes a point. If a load is now applied to the gears, resulting from elastic deformation of the teeth, the point becomes an area of contact, and the shape of this area will be dependent upon mismatch contour lines, Fig. 10. These contour lines represent constant separation between a modified tooth surface and a theoretically conjugate tooth surface under the existing conditions of load and elastic deformation. In like manner, contour lines may be drawn to represent the area of tooth bearing, which is the summation of all the areas of contact as the gears rotate, Fig. 11.\nSince the tooth bearing represents all points of contact on the tooth surface, and since contact takes place in the surface of action, there is a point-to-point correspondence between conditions of contact on the tooth surface and those existing in the zone of action. Because the zone of action shows the contact on all teeth in action simultaneously, it is convenient to use diagrams in this plane.\nAssuming that the tooth bearing has been \"developed\" to take into account elastic deformation of the gear teeth and deflection of the gear mountings under maximum load, the contact will be considered to lie within the elliptical boundary tangent to the four sides of the theoretical zone of action as shown in Fig. 12, and the corners outside of this ellipse play no active part in carrying the load.\n1 1 8 / A P R I L 1 9 6 7 Transactions of the AS ME\nDownloaded From: http://tribology.asmedigitalcollection.asme.org/ on 01/29/2016 Terms of Use: http://www.asme.org/about-asme/terms-of-use",
            "Fig. 13(a) Lines of contact\u2014(a) shows the lines of contact in a pair of gears when only one tooth is in actual contact\nAmount of Load Carried by One Tooth In Fig. 13(a) the theoretical lines of contact AB, CD, and EF are drawn in the position when line AB is tangent to the ellipse and is therefore outside of the ellipse. (Note that lines AB, CD, and EF are three successive parallel lines of contact on three successive teeth.) EF is also totally outside of the ellipse. Therefore, the total load is carried on the portion of line CD between V and W ; that is, all on one tooth. Fig. 13(6) shows the case of another pair of gears where there are always at least two teeth in contact. In the particular position shown, three teeth are sharing the load.\nAlong these lines of contact the load will vary from a maximum near the center to zero at the boundaries of the ellipse. These lines of contact, according to the simplified theory used here, will be ellipses of contact. This is shown by the smaller ellipse V X W Y in Fig. 14. The load distribution along the major axis, VW, of the ellipse of contact is shown in Fig. 15 by the semiellipse VUW, and along the minor axis, X Y , by the semiellipse X U ' Y . The representation of the resultant load distribution over the ellipse of contact is, therefore, semiellipsoidal. (Unit load O'U = O'U' . )\nIn Fig. 16 three lines of contact ST, VW, and QR are shown within the boundaries of the ellipse. Two secondary ellipses of constant separation shown dotted in the diagram are drawn tangent to chords ST and QR. The respective points of tangency are U and P. The distance between successive lines of contact in the plane of action is the normal base pitch, p\u201e, Fig. 17. It is assumed that the load distribution along chords ST and QR is semiellipsoidal. It is also assumed that the contact stress at P is equal to the contact stress at L and the contact stress at U is equal to the contact stress at Ar, because these points are on the respective ellipses of constant separation.\nIt can then be shown that [16]:\nFig. 15 Assumed load distribution along line of contact VW. This is an oversimplification.\nFig. 16 The two small dotted ellipses here are secondary ellipses of contact. The three ellipses, QR, VW, and ST, are on three adjacent teeth.\nmN -cr (21) Fig. 17 ration The dotted ellipses are the secondary ellipses of constant \u00bbepa.\nJournal of Lubrication Technology A P R I L 1 9 6 7 / 1 1 9\nDownloaded From: http://tribology.asmedigitalcollection.asme.org/ on 01/29/2016 Terms of Use: http://www.asme.org/about-asme/terms-of-use"
        ]
    },
    {
        "image_filename": "designv11_12_0002335_detc2010-28926-Figure10-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002335_detc2010-28926-Figure10-1.png",
        "caption": "Figure 10. The result of applying fundamental and regular extensions (a1,b1) on ground and inner edges (a,b),",
        "texts": [
            " The first extension, termed fundamental extension, produces all the Assur Graphs in the first row, called also the fundamental Assur Graphs. This operation is done by replacing a ground edge by a triangle and two new ground edges, as shown in 10(a,a1). The fundamental AGs can also be related as representatives, since from each one of them it is possible to derive an infinite number of different AGs. This is done by applying a second extension, termed regular extension, that divides, splits, one of the edges (x,y) by a new vertex, z, and adding a new edge (z,t) for some vertex t/=x,y, as shown in Figure 10(b,b1). Figure 11 depicts example of AGs that are the result of applying a sequence of extensions, starting from the basic AG \u2013 the dyad (Figure 11a). The first row presents the fundamental AGs, called also the representatives, all derived from the basic dyad through applying the fundamental extensions. For example, the fundamental AG in Figure 11b known also as Triad, is obtained by replacing the ground edge (A,O2) with the triangle <A,B,C> with the two new ground edges (C,O2) and (B,O3). All the other infinite fundamental AGs are obtained in the same way; each time a ground edge is replaced with a triangle and two new ground edges"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001724_tmag.2010.2072910-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001724_tmag.2010.2072910-Figure2-1.png",
        "caption": "Fig. 2. Three-phase induction motor model core.",
        "texts": [
            " In (1), the magnetic field strength Hz was calculated with the first and second term by the static integration-type E&S modeling, the third term represents that the magnetic field strength generated by the eddy current due to the fundamental frequency changes Hz , and the fourth term represents that the magnetic field strength generated by the eddy currents in higher harmonic components Hz . 0018-9464/$26.00 \u00a9 2011 IEEE The governing equation of the two-dimensional magnetic field considering the two-dimensional vector magnetic properties expressed by the dynamic E&S modeling can be written as (3) where is the magnetic vector potential, the exciting current density, the electric scalar potential, and the conductivity. Fig. 2 shows the three-phase induction motor model core used in the analysis. This is the provided model by the Institute of Electrical Engineers of Japan (IEEJ). The numbers of slots in the stator and rotor core were 24 and 34, respectively. The exciting voltage and frequency were 50 V and 50 Hz. The electrical steel sheet was assumed to be 50A470, and the rolling direction was parallel to 0 (horizontal direction of this figure). As shown in Fig. 2, three-phase stator windings were distributed. The model core was divided into the linear triangular elements. The number of nodes was 22 936, and the number of elements was 44 263. In this mesh, the air gap was divided into eight layers. In the analysis, the periodic boundary condition of 180 was used, and the fixed boundary condition was employed at the outer air region. Fig. 3(a)\u2013(c) shows the distributions of the maximum magnetic flux density at slip 0, 0.2, and 0.5, respectively. At the no-load condition slip as shown in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000880_tmag.2007.891399-Figure9-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000880_tmag.2007.891399-Figure9-1.png",
        "caption": "Fig. 9. Measurement system of EMI drive mechanism. (a) Overview; (b) core and conductor; (c) ring magnet.",
        "texts": [
            " Usually, the eddy current torque of Lorentz force is negative torque when the rotation direction of rotor is constant, however, the calculated results show negative and positive values. Therefore, the eddy current torque of Lorentz force and the magnetic attractive torque between the ring magnet and stator core are separated. Fig. 8 shows the calculated time variations of eddy current torque and magnetic attractive torque, respectively. It is found that the magnetic attractive torque is generated when the eddy current torque becomes minimum. Fig. 9 shows the EMI torque measurement system. In this measurement, the inner and outer cores with conductors are rotated instead of the ring magnet, which are connected to the motor. The ring magnet connected to the torque meter is fixed. The EMI torque amplified by the mechanical resonance is measured using torque meter when the cores and conductors are rotated by motor. Fig. 10 shows the measured time variations of EMI torque when the motor rotates at the rotation speed of 2000 rpm. From this figure, it is found that the measured EMI torque has same period with the calculated transient torque"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000400_1.2709514-Figure3-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000400_1.2709514-Figure3-1.png",
        "caption": "FIG. 3. Color online Optical images of droplet movement using dynamic changes of ferrofluid pattern. Droplet size is 8 l.",
        "texts": [
            " Dynamic variations of the stirrer field through its continuous rotation changes the field configuration accordingly. We investigated the transport speed of various sized water droplets 5, 8, and 16 l under different rotation speeds 100, 200, 300, 400, and 500 rpm of the magnetic stirrer. For the measurement of the transport speed of water droplets, a single water droplet was dropped on top of the dynamic ferrofluid pattern and was driven to the target spot which is 2 cm away from the starting point. Figure 3 shows two 8 l water droplets dispensed by a micropipette driven to the other corner of the silicon wafer from the dispensing point. Since the magnetic stirrer rotates clockwise at 200 rpm, the droplets also move in the same direction in a controlled manner. The measurement results of various sized droplets are shown in Fig. 4. Due to the limited accuracy in the measurement, our results at measuring droplet speed for rpm higher than 600 rpm are not included in this figure. It was found that the transport speed of droplet at the same rotation speed of the magnetic stirrer does not depend on the droplet size as long as the volume remains less than 16 l"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002732_isie.2011.5984383-Figure4-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002732_isie.2011.5984383-Figure4-1.png",
        "caption": "Fig. 4. Vision subsystem.",
        "texts": [
            " In this work, it is proposed to improve this performance by using FPGAs as processing elements, taking advantage of their operation speed and abundance of resources, including embedded processors [7]. The flexible architecture of FPGAs and their ability to perform tasks in parallel, has made them to be increasingly used in industry for implementing signal and image processing and control systems [8]-[11]. A. Vision subsystem The cladding system in the Laser Application Center includes a vision subsystem (Fig. 4) in a coaxial arrangement with the laser beam. It consists of a CMOS camera with 8-bit (256 grayscale levels) CameraLink interface and dynamic range up to 120dB, and auxiliary optical parts: mirrors, a lateral augmentation telescopic system, and two filters, one to prevent the laser beam to damage the camera, and another to adjust the saturation of the camera by the light generated in the fusing process. In this work, the configurable resolution of the camera has been set to 800x600 pixels. B. Image processing According to the previous experience of the authors [12], and as also assumed in [13], the geometric property to be extracted from the images is the width of the melt pool, which is directly related to dilution"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000597_978-3-540-73129-0-Figure6.19-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000597_978-3-540-73129-0-Figure6.19-1.png",
        "caption": "Fig. 6.19. Mechanical displacement amplification. a Implementation with elastic hinges, b moonie transducer, c amplified piezo actuator, APA (derived from [5])",
        "texts": [
            " In piezoelectric transducers with displacement amplification the achieved deflection is increased by constructive means. The stiffness of such a design decreases with the square of the displacement amplification ratio and is therefore much smaller than in the stack design. This kind of transducer used for displacements of up to 1 mm with forces of several tens of Newtons is achieved, for instance, with elastic joints or hinges. These elastic hinges transform small angular alterations into parallel movements free of backlash. Figure 6.19 illustrates the principle. The highly elastic material region of the displacement amplifier in Fig. 6.19a is locally concentrated, while the designs in Fig. 6.19b and c make use of the global elastic behaviour of metallic materials. The so-called moonie transducer in Fig. 6.19b consists of a piezoelectric disk sandwiched between two metal end caps. The ceramic is poled in the thickness direction and uses the d31 mode. In this way the small radial displacement of the disk is transformed into a much longer axial displacement normal to the surface of the caps. The moonie design is very simple and its manufacture can easily be automated. It generates moderate forces and displacements, filling the gap between bimorph and multilayer actuators [9]. Figure 6.19c shows a related design in which the piezo stack and subsequently the d33-mode are used. The advantages of these APAs (amplified piezo actuator) are their relatively high displacements combined with its large forces and compact size along the active axis [10]. Figure 6.20 shows an entirely different solution. A hydraulic force-displacement transformer functions according to the two-piston hydraulic principle. Leak-free operation is achieved in the presented design through the use of two folding bellows of different effective diameters"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003520_j.euromechsol.2011.11.003-Figure5-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003520_j.euromechsol.2011.11.003-Figure5-1.png",
        "caption": "Fig. 5. T2R1-type parallel manipulator with uncoupled and bifurcated planar-spatial motion of the moving platform: constraint singularity (a), branch with planar motion (b) and branch with spatial motion (c); limb topology PkRkRkRtRPtPttRtRSRkRkP.",
        "texts": [],
        "surrounding_texts": [
            "The term of constraint singularity (CS) has been recently coined (Zlatanov et al., 2002) to characterize the configuration of lower mobility parallel manipulators in which both the connectivity of the moving platform and the mobility of the parallel mechanism increase their instantaneous values. From a constraint singularity, the mechanism can get out with or without branching (Gogu, 2008b). When branching occurs in a constraint singularity, the mechanism can reach different configurations, called branches, and have different independent motions of the moving platform. In this case, the constraint singularity is also called branching or bifurcation singularity. A branch refers to the free-of-singularity configurations of the mechanism inwhich each structural parameter keeps its value. For this reason, this value is called global of full-cycle value for a branch. Two types of branching in constraint singularity (BCS) have been defined in Gogu (2008b). Branching of type BCS1 occurs when a parallel mechanism F ) G1 . Gj . Gk gets out from a constraint singularity in different branches characterized by the same mobility and the same connectivity of the moving platform but with different bases of the vector space of relative velocities between the moving and fixed platforms. Branching of type BCS2 occurs when a parallel mechanism F ) G1 . Gj . Gk gets out from a constraint singularity in different branches characterized by distinct values of mobility and connectivity of the moving platform. The parallel manipulators presented in this paper have constraint singularities with branching of type BCS1. Branching of type BCS2 occurs in kinematotropic mechanisms (Gogu, 2009c). The finite displacements and the velocities in the actuated joints are denoted by qi and _qi, the linear velocities of the characteristic point H of the moving platform, by v1 \u00bc _x; v2 \u00bc _y and the angular velocity of the moving platform by u \u00bc ua \u00bc _a or u \u00bc ud \u00bc _d: In both branches, the moving platform undergoes two planar translations and one rotation but the rotation axis is different in the two branches. In the first case, the rotation velocity u \u00bc ua \u00bc _a is parallel to the x-axis and the moving platform undergoes a spatial motion. In the second case, the rotation velocity u \u00bc ud \u00bc _d is parallel to the z-axis and the moving platform undergoes a planar motion. With these notations, the linear mapping between the actuated joint velocity space and the end-effector velocity space for a branch the T2R1-type PMs is defined by: 2 4 v1 v2 u 3 5 \u00bc \u00bdJ 2 4 _q1 _q2 _q3 3 5 (8) where J is the Jacobian matrix. In the constraint singularity, the T2R1-type parallel manipulators with bifurcated motions has instantaneously iM \u00bc iSF \u00bc 4 and (iRF)\u00bc(v1, v2, ua, ud). The bifurcation in this constraint singularity can be used to change motion type of the moving platform. In the two distinct branches, the parallel mechanism is characterized by M\u00bc SF \u00bc 3 and (RF)\u00bc(v1, v2,ua) or (RF)\u00bc(v1, v2,ud). In the constraint singularity, both connectivity of moving platform and mobility of parallel manipulator increase their instantaneous values. More details about the behaviour of parallel manipulators in a constraint singularity and the associated structural parameters can be found in Gogu (2009c)."
        ]
    },
    {
        "image_filename": "designv11_12_0002207_s10955-009-9681-9-Figure5-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002207_s10955-009-9681-9-Figure5-1.png",
        "caption": "Fig. 5 (A) Two non-retractable loops on a torus. (B) Coordinate system for torus. \u2212\u03c0 < \u03b8 \u2264 \u03c0 , 0 \u2264 \u03c6 < 2\u03c0",
        "texts": [
            " The first Betti number of cylinder is one; hence only one nontrivial cycle, i.e., the loop \u03b3\u03d5 winding around the cylinder once, and only one nontrivial cocycle, i.e., \u03c8 = d\u03d5. The 1-form (23) is the linear superposition of this cocycle d\u03d5 and an exact form a dz. 2.4 Smectic Order on Torus Let us apply the cohomology theory to defects-free smectic packing on a torus. The first Betti number of a torus is b1 = 2 [13, 34], hence there are two independent 1-cycles, \u03b3\u03b8 and \u03b3\u03c6 , wrapping around the torus in two different directions, as shown in Fig. 5A. Correspondingly there are two independent 1-cocycles, naturally chosen as d\u03b8 and d\u03c6, where \u03b8 and \u03c6 are the angular coordinates wrapping around the loops \u03b3\u03b8 and \u03b3\u03c6 , see Fig. 5B. We easily see that {\u03b3\u03b8 , \u03b3\u03c6} is indeed the dual basis of {d\u03b8, d\u03c6}: \u222e \u03b3\u03b8 d\u03b8 = \u222e \u03b3\u03c6 d\u03c6 = 2\u03c0, \u222e \u03b3\u03b8 d\u03c6 = \u222e \u03b3\u03c6 d\u03b8 = 0. d\u03b8 and d\u03c6 are not exact, since \u03b8 and \u03c6 are not single valued function. An arbitrary defects-free smectic state \u03c8 on a torus is therefore characterized two global dislocation charges N\u03b8 and N\u03c6 : \u03c8N\u03b8 ,N\u03c6 = N\u03b8 d\u03b8 + N\u03c6 d\u03c6 + d (\u03b8,\u03c6) = (N\u03b8 + \u2202\u03b8 )d\u03b8 + (N\u03c6 + \u2202\u03c6 )d\u03c6, (29) where \u222e \u03b3\u03b8 \u03c8 = 2\u03c0 N\u03b8, \u222e \u03b3\u03c6 \u03c8 = 2\u03c0 N\u03c6. Two integers (N\u03b8 ,N\u03c6) classify all topologically distinct, defects-free smectic states on a torus, and are the minimal numbers of smectic layers intersecting the non-retractable loops \u03b3\u03b8 and \u03b3\u03b8 respectively",
            " (31) Since translation of \u03c6 by \u03c0 amounts to rotation of the torus by \u03c0 around the z-axis, we find that a state (N\u03b8 ,N\u03c6) is transformed into (\u2212N\u03b8,N\u03c6) by the space inversion. Consequently states with N\u03b8 = 0 and N\u03c6 = 0 are not invariant under spatial inversion, i.e. they are chiral. We shall call states with N\u03b8 > 0 left handed, while states with N\u03b8 < 0 right handed. Clearly a left handed state is transformed into a right hand state under spatial inversion, and vice versa. To understand the geometric properties of the states (29) let us calculate the layer spacing for the special case with = 0. With the coordinate system shown in Fig. 5B, the torus is parameterized by two angles (\u03b8,\u03c6) as r(\u03b8,\u03c6) = \u239b \u239d (R\u03c6 + R\u03b8 cos \u03b8) cos\u03c6 (R\u03c6 + R\u03b8 cos \u03b8) sin\u03c6 R\u03b8 sin \u03b8 \u239e \u23a0 . (32) The corresponding metric tensor can be easily calculated using (9): g\u03b8\u03b8 = R2 \u03b8 , g\u03c6\u03c6 = (R\u03c6 + R\u03b8 cos \u03b8)2, g\u03b8\u03c6 = g\u03c6\u03b8 = 0. (33) We shall also define the aspect ratio r = R\u03b8 R\u03c6 . (34) Note that 0 < r < 1 if the cylinder does not self-intersect. We shall be especially interested in the thin torus limit r 1, where r provides a natural small parameter that greatly simplifies the analysis"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003851_icuas.2015.7152312-Figure3-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003851_icuas.2015.7152312-Figure3-1.png",
        "caption": "Fig. 3: Engagement geometries for the three guidance laws (PP, LOS, and PPN)",
        "texts": [
            " Guidance in general refers to the determination of the heading from the vehicle\u2019s current location to a designated target. There are three main classical guidance laws [24]. Figure 2 gives a basic comparison between the velocity commands of the three guidance laws. In PP, velocity vector is directed towards the target, in LOS it is directed so as to make \u03c7u=\u03b8 and in PPN velocity vector is such that the target is always perceived at a constant bearing from the UAV. Pure pursuit (PP) guidance belongs to the two-point guidance schemes. Figure 3a shows the engagement geometry between the UAV and the target. The tracking vehicle is supposed to align its velocity along the line of sight between the UAV and the target. This strategy is equivalent to a predator chasing a prey, and very often results in a tail chase as can be seen in figure 4a. Equation (3) is the expression for guidance command which further gives the change in missile velocity vector angle. R\u0307 = {Vt cos(\u03c7t \u2212 \u03b8)\u2212 Vu cos(\u03c7u \u2212 \u03b8)}, (1) \u03b8\u0307 = {Vt sin(\u03c7t \u2212 \u03b8)\u2212 Vu sin(\u03c7u \u2212 \u03b8)}/R, (2) au = Vu\u03b8\u0307 \u2212Ka(\u03c7u \u2212 \u03b8), (3) \u03c7\u0307u = au/Vu",
            " (4) Line of sight (LOS) guidance is classified as a three-point guidance scheme since it involves a (typically stationary) reference point in addition to the tracking vehicle (UAV) and the target. Here, the vehicle is supposed to reach the target by constraining its motion along the line of sight between the reference point and the target. LOS guidance has typically been employed for surface-to-air missiles, often mechanized by a ground station which illuminates the target with a beam that the guided missile is supposed to ride, also known as beam-rider guidance. Figure 3b shows the engagement geometry for the vehicle for line of sight guidance. Equation (7) gives the expression for acceleration command whereas equations (5) and (6) are the equations of motion for LOS guidance law. Figure 4b shows the trajectory for LOS guidance law in matlab 2D. \u03b8\u0307 = Vt sin(\u03c7t \u2212 \u03b8t)/Rt, (5) R\u0307 = Vu cos(\u03c7u \u2212 \u03b8u), (6) au = KbRu(\u03b8t \u2212 \u03b8u) +Ru\u03b8\u0308 + 2Ru\u03b8\u0307, (7) \u03c7\u0307u = au/Vu. (8) Proportional Navigation Guidance law is based on Constant Bearing (CB) principle. In CB engagement, the UAV is supposed to align the relative UAV-target velocity vector along the LOS between the UAV and the target. This is equivalent to reducing the LOS rotation rate to zero such that the UAV sees the target at a constant bearing, on a direct collision course. The most common method of doing this is to make the rate of turn of UAV velocity vector proportional to the LOS rate. Figure 3c shows the engagement geometry for the vehicle for proportional navigation guidance. Equation 8 gives the expression for the acceleration command for PPN. Figure 4c shows the trajectory for PPN guidance law in matlab 2D. au = NVu\u03b8\u0307, (9) \u03c7\u0307u = au/Vu. (10) where expression for \u03b8\u0307 and R\u0307 is as given in equation (1) and (2). Figure 2 and equations (1)-(9) show the velocity vector details and engagement equations of the three guidance principles respectively, in a 2D scenario. In a 3D guidance scenario the parameters to be considered are the course and the flight path angles of the UAV in xy and yz plane"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003463_1.4004588-Figure4-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003463_1.4004588-Figure4-1.png",
        "caption": "Fig. 4 Theoretical profile of the driven rotor",
        "texts": [
            " Manuscript received March 8, 2011; final manuscript received July 4, 2011; published online September 7, 2011. Assoc. Editor: Prof. Philippe Velex. Journal of Mechanical Design SEPTEMBER 2011, Vol. 133 / 094501-1Copyright VC 2011 by ASME Downloaded From: http://mechanicaldesign.asmedigitalcollection.asme.org/ on 09/04/2013 Terms of Use: http://asme.org/terms 2.2 Edge Blunting of the Driven Rotor. The theoretical profile of the driven rotor consists of three segments: a dedendum circle, , a cycloidal curve, , and an addendum circle, , as shown in Fig. 4. The cycloidal curve is an extended epicycloid, generated as the trajectory of point M rigidly connected to the circle O1 that rolls over circle O2, N is the intersection point of the epicycloid, , with the addendum circle, , and the angle at this point between the two limiting rays of the two arcs is 61.87 deg. Meanwhile, N is the generating point of the driving rotor profile. Edge blunting is carried out for two reasons: the meshing point is moved from the outer circumference to be more durable and angle b at the transition point is enlarged to make the profile less sharp"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001247_0954406jmes321-Figure16-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001247_0954406jmes321-Figure16-1.png",
        "caption": "Fig. 16 Face gear, casting mold, and helical pinion of the fishing reel",
        "texts": [
            " This example shows that the bias angle of the contact path, the shape of the motion curve, and the bearing ratio can be controlled independently with the aid of the proposed double-crowning methodology. Fig. 13 Transmission error results for the five example cases This example validates the use of the proposed methodology in production. The face-gear set applied here to fishing reels is shown in Fig. 15, while the design parameters are listed in Table 3. This face-gear set includes a machined helical pinion, cut by a standard hobbing process without crowning, and a face gear made by mould casting as shown in Fig. 16. The three-dimensional tooth surface and the CNC code of the face gear were produced by the computer software developed for the proposed double-crowning methodology. Since the tooth number of the pinion is small but the helical angle is large, it is very difficult for Proc. IMechE Vol. 221 Part C: J. Mechanical Engineering Science JMES321 \u00a9 IMechE 2007 at UNIV OF CONNECTICUT on April 13, 2015pic.sagepub.comDownloaded from Fig. 14 Relation between the contact ratio and coefficient \u00b5 Fig. 15 Three-dimensional CAD model of the face-gear driver set Table 3 Design parameters applied in example 2 Design parameters Pinion tooth number N1 7 Shaper tooth number Ns 7 Face-gear tooth number N2 34 Normal pressure angle \u03b1n (\u25e6) 20 Helical angle \u03b2 (\u25e6) \u221240 Module m (mm) 0"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003713_978-1-4614-2419-2-Figure50.6-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003713_978-1-4614-2419-2-Figure50.6-1.png",
        "caption": "Fig. 50.6 Ice-skate contact forces and strain gauges position",
        "texts": [
            " Measured signals have been recorded on a MTA Digitek Cobra data-logger also placed on board the vehicle and subsequently transferred to a PC for post-processing via an Ethernet link. All the signals have been acquired at a frequency of 100 Hz and low-pass filtered at 10 Hz. Ice-skate contact forces reconstruction is based on the measurement of the deformations of the axle attached to the runner carriers, which are induced by the contact forces themselves. Ice-skate contact forces and the position of the strain gauges on the axle are represented in Fig. 50.6. The number and the position of the strain gauges was bounded by the limited space available between the connections of the axle with the skate and the bobsleigh frame (about 100 mm). Strain gauges 1 have two measuring grids perpendicular one to the other and they are connected to form a full bridge allowing the axial force (Fy) measurement and the compensation of thermal effects and bending. Fig. 50.4 Rotary potentiometer for steer angle measurement 883.5mm G D 130mm 207mm P Fig. 50.5 Position of the inertial gyroscopic platform (P) and of the optical sensor with respect of the bobsleigh cog (G) The pair of strain gauges 2\u20133 (which is connected in a half bridge configuration balanced by means of to resistors) allows the measurement of the bending moment about the x axes in sections AA (MxAA) and BB (MxBB)",
            " The measurement vector f \u00bc {Fy MxAA MxBB MzAA MzBB} T can be evaluated from the vector s of the output signals of the strain gauge bridges by means of a proper calibration matrix [C] (the definition of the calibration matrix will be discussed later on): f \u00bc C\u00bd s (50.1) The lateral component of the ice-skate contact force is already included into vector f, while components Fz and Fx can be determined by applying the following relations: Fz \u00bc MxBB MxAA b ; Fx \u00bc MzBB MzAA b (50.2) being b the distance between sections AA and BB (see Fig. 50.6). It must be in fact considered that: \u2022 The lateral force Fy provides a constant contribution to the bending moments about both the x and the z axis; \u2022 The vertical force Fz originates a bending moment about the x axes, which increases proportionally to the distance from the skate (triangular distribution, Fig. 50.7); \u2022 The longitudinal force Fx originates a bending moment about the z axes, which increases proportionally to the distance from the skate (triangular distribution, Fig. 50.8). Fx, Fy, Fz"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003761_9781118818428-Figure20.2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003761_9781118818428-Figure20.2-1.png",
        "caption": "Figure 20.2 Notation for instantaneous, local effective temperature difference \u03b5T = 1\u20442(Twd + Twc) \u2013 T of the regenerative annulus",
        "texts": [
            " The \u2018hot air\u2019 engine 223 The Schmidt, the \u2018adiabatic\u2019 and other \u2018ideal\u2019 cycle analyses rely on the assumption that gas and wall temperature distributions are linear and identical between cycle temperature limits TE and TC. Applied to the \u2018serious\u2019 Stirling engine the assumption is justifiable on the basis that design aims for that ideal. It is inappropriate to the hot air type, where the temperature gradients on inner and outer surfaces of the regenerative annulus not only differ, but are subject to substantial relative change during each cycle. This page establishes a more realistic picture, and one which serves better for first principles design. Figure 20.2 is a simplified representation of cylinder and displacer, and indicates a notional lengthwise temperature distribution Tw for both. The axial length of displacer is Ld; the stroke is Sd. The annular gap is formed between displacer body of length Ld and the identical length of the enclosing cylinder. However, the latter length is a fraction Ld\u2215(Ld + Sd) of total axial length Ld + Sd of the enclosing cylinder. Moreover, that fraction lies at a different axial location at each instant, so that, relative to the displacer, a different enclosure temperature distribution applies. 224 Stirling Cycle Engines End temperatures of the enclosing cylinder are nominally TE and TC. The extremes of displacer temperature TdE and TdC are not known, although it is beyond doubt that TdE < TE and that TdC > TC. It is worth proceeding despite this uncertainty, because the result will be a formulation allowing ready substitution of an improved picture as and when available. Figure 20.2 focuses on the pair of adjacent wall temperatures Twd and Twc lying at axial distance x, where the datum for x is the head of the displacer. The variable axial distance xd between this latter datum and cylinder head is acquired by inverting the expression for instantaneous expansion-space volume Ve(\u03c6) = 1\u20444\u03c0D2xd, viz: xd = 4Ve(\u03c6)\u2215\u03c0D2 (20.6) With \u03b6 to denote fractional distance x\u2215Ld: Twd = TdE + (TdC \u2212 TdE)\u03b6 (20.7) Twc = TE + (TC \u2212 TE)(\u03bce(\u03c6)\u03bbSd\u2215Ld + \u03b6)\u2215(Sd\u2215Ld + 1) \u03bce(\u03c6) = Ve(\u03c6)\u2215Vsw \u03bb = Sp\u2215Sd = Finkelstein\u2019s kinematic volume ratio (20"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000586_978-1-4020-8889-6_6-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000586_978-1-4020-8889-6_6-Figure2-1.png",
        "caption": "Fig. 2 General scheme of helicopter, load attachment",
        "texts": [
            " Figure 1a shows the helicopter during hovering, where the force required to lift the helicopter FLift is equal to the force generated by the main rotor FMR 3 . During forward flight (see Fig. 1b) the force FMR 3 is split into the lifting force FLift and the force FAcc used for the acceleration of the helicopter. To keep the helicopter at the same height, the magnitude of FLift has to be preserved and therefore the magnitude of FMR 3 has to be increased. To sum up: The coupling between rotation and translation (for a helicopter without slung load) can be expressed in the following relation: \u201cthe orientation has effect on the translation\u201d. In Fig. 2 the attachment of a slung load to a helicopter fuselage is schematized. The rope connecting helicopter and load is attached to the fuselage at point r . The force caused by the load in point r is given by the vector Fr. The vector pr cm connects the mounting point r and the center of mass cm of the helicopter. The load angle \u03b8 relative to the fuselage is defined as the angle between the extension of the vector pr cm and the rope. Because of constructional limitations it is normally not possible to attach the rope directly in the center of mass cm of the helicopter. For that reason the vector pr cm connecting r and cm is assumed not to be zero during the following considerations. During hovering of the helicopter, with the load at rest and no perturbations, the angle \u03b8 is zero and no torque is caused by the load. If helicopter and load are considered to be in the state as depicted in Fig. 2, the load angle \u03b8 is not zero and the load causes a torque that will change the helicopters orientation. A change of the helicopters orientation leads to acceleration/deceleration of the helicopter (see Fig. 1b). The acceleration/deceleration of the helicopter leads to a change of the load angle \u03b8 and to acceleration/deceleration of the load. This again changes Fr and the torques acting on the helicopter. Therefore the following relation is true for a helicopter with a slung load attached: \u201corientation has effect on translation and vice versa\u201d"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003761_9781118818428-Figure11.6-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003761_9781118818428-Figure11.6-1.png",
        "caption": "Figure 11.6 Notation for algebra of thermal bottleneck and Availability calculations",
        "texts": [
            " The simple algebra covers thermal conduction through heat exchanger walls (where, however, the temperature difference driving heat flow through the solid is more appropriately denoted \u0394T). There is no catch! Numerical values of \u03b5T (and\u0394T) resulting from application of Availability algebra may, however, come as a surprise. It will be easier to \u2018sell\u2019 this elegant, labour-saving approach after discussing a simple case showing that it yields a numerical result identical to that returned by traditional arithmetic. Availability Theory elegantly quantifies lost available work: The upper-left diagram of Figure 11.6 indicates flow of heat by conduction at rate q\u2032 between a source at TE and a sink at TC. Entropy rate as q\u2032 [W] enters at TE is (by definition) q\u2032\u2215TE [W/K]. As q\u2032 leaves, entropy rate has increased to q\u2032\u2215TC W/K \u2013 increased because, although q\u2032 is unchanged, TC < TE. A process of thermal conduction has generated entropy at rate s\u2032: s\u2032 = q\u2032{1\u2215TC \u2212 1\u2215TE}W\u2215K. Allowing heat to be downgraded by conduction has forfeit an opportunity for partial conversion to work. The rate of loss of potential work is given in Availability notation as T0s \u2032 [W], in which T0 is the lowest temperature at which q\u2032 could realistically be rejected",
            " The probes would have responded to radiation from the inside of the tubes, suggesting an over-indication. With little else to go on, an \u2018educated guess\u2019 in the style of the legendary, late Professor G (Joe)Walker suggests mean temperature difference between heat source and point of heat reception by the gas can be 1800 \u25e6C\u2013600 \u25e6C = 1200 \u25e6C! It puts the principal thermal bottleneck at the expansion end. The GPU-3 tests noted compressionspace temperatures of 247 \u25e6F, or 119.4 \u25e6C \u2013 a penalty to be reckoned with (but after the main culprit has been addressed). The lower-most image of Figure 11.6 indicates schematically the thermal conduction path between combustion reaction products at Tcomb and the working gas at the notional cycle mean Tgxe of its fluctuating temperature. The path may be seen as 106 Stirling Cycle Engines three \u2018thermal bottlenecks\u2019 in series: (a) between products at Tcomb and external surface of the expansion exchanger tubes at Txo, (b) between Txo and the inner surface of the tubes at Txi, (c) between Txi and working fluid at Tgxe. Figures in Table 11.1 are for General Motors\u2019 GPU-3 engine at the peak power point",
            " The focus to this point has been on the high-temperature end of the cycle. It should nevertheless be obvious that achieving high efficiency requires that the working fluid (rather than the heat exchangers) should receive and reject heat at temperatures as close as possible to those of source and sink respectively. The criterion is consistent with minimizing temperature differences along the heat-flow path. In engineering terms, cutting down on such losses amounts to minimizing \u2018thermal bottlenecks\u2019. Figure 11.6 identifies three bottlenecks (or \u2018thermal impedances\u2019) in series at the expansion end alone: Finite thermal conductivity k [W/mK] of tube walls. The 8.9 kW output of the GPU-3 at 22.5% efficiency equates to a heat input rate of 8.9\u22150.225 = 39.5 kW. Steady conduction through the walls of the 40 exchanger tubes is governed by the equation q\u2032 = \u2013kA\u0394T\u2215tw, where k is coefficient of thermal conductivity [W/mK], A is total surface area [m2] of the Getting started 107 \u0394Tcond = 39.5E + 03[W] \u00d7 0.905E \u2212 03[m] 20 [W\u2215mK] \u00d7 40 \u00d7 0"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002550_s0219455412500186-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002550_s0219455412500186-Figure2-1.png",
        "caption": "Fig. 2. The stress and strain components act on the cylindrical shell.",
        "texts": [
            " D ow nl oa de d fr om w w w .w or ld sc ie nt if ic .c om by J O H N S H O PK IN S U N IV E R SI T Y o n 01 /0 3/ 15 . F or p er so na l u se o nl y. which has a solution of the form p\u0302 \u00bc anr n \u00fe bnr n; \u00f011\u00de where an and bn are two constants to be determined using the \u00b0uid boundary conditions. 2.2. Structural dynamics equations The strain energy U for a cylindrical shell can be written as: U \u00bc 1 2 Z l 0 Z 2 0 Z h=2 h=2 \u00f0 x\"x \u00fe \" \u00fe x \"x \u00derdzd dx; \u00f012\u00de where x, , and x are the stress components act on the shell as shown in Fig. 2, and \"x, \" , and \"x are the strain components. The kinetic energy of the rotating cylindrical shell is given by K \u00bc 1 2 sh Z 2 0 Z l 0 \u00f0u: 2 \u00fe v :2 \u00fe w : 2 \u00fe 2 \u00f0vw: wv :\u00de \u00fe 2\u00f0v2 \u00fe w2\u00de\u00derdxd ; \u00f013\u00de where u, v, and w are the axial, tangential, and radial displacement component of the shell deformation, respectively. According to the \u00afrst-order shear deformation theory, the displacement \u00afelds of the shell can be written as u\u00f0x; ; r\u00de \u00bc u0\u00f0x; \u00de \u00fe z ; v\u00f0x; ; r\u00de \u00bc v0\u00f0x; \u00de \u00fe z x; w\u00f0x; ; r\u00de \u00bc w0\u00f0x; \u00de; \u00f014\u00de where u0, v0, and w0 are the middle surface displacements, and x respectively show the rotations of the middle surface about the and x directions and z is the radial distance from the middle surface"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003761_9781118818428-Figure5.1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003761_9781118818428-Figure5.1-1.png",
        "caption": "Figure 5.1 Skeleton of drive mechanism of Karabulut engine. Not to scale! \u2013 the figure serves to establish notation",
        "texts": [
            " On removal of the heat source, no-load running continued down to TE = 75 \u25e6C. Here is a prototype which evidently wants to run \u2013 and to do so between un-promising temperature limits \u2013 in stark contrast to the majority of prototypes which manifestly prefer not to run \u2013 even when roasted to red heat. The evident free-running is the more remarkable for having been achieved with displacer driven by a mechanism not known for low friction \u2013 a slider similar to the \u2018rapid return\u2019 of a machine-shop shaping machine (Figure 5.1). In general, intermediate temperature operation promises benefits: the range of available \u2018fuels\u2019 widens to include waste heat and geothermal sources. Operation from solar energy can be achieved despite relatively imprecise focusing and tracking. The requirement for sophisticated materials and fabrication techniques for hot-end components is relaxed. It would be useful to be able to decipher the Karabulut account for the secret of the success. The technical specification (Table 1 of the original account) yields nothing remarkable: Displacer L\u2215D ratio is a healthy 3"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001540_s11431-010-4176-0-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001540_s11431-010-4176-0-Figure1-1.png",
        "caption": "Figure 1 Structure and theory indication for the pump. 1, Vibrating amplifier; 2, piezoelectric ceramics; 3, clamped plane.",
        "texts": [
            " This paper proposed a fishtailing type of valveless piezoelectric pumps, which is based on the fish swimming principle that the water will flow backward relative to the fish if the fish is swimming. The relationship between flow rates and vibrating frequencies was derived from the interaction between the vibrator and fluid. Numerical simulations with FEM software were conducted to study the first and second vibration modes for the piezoelectric vibrator. The experiments were also conducted with a Doppler laser vibration measurement system to test the flow rates with different driving frequencies. The experimental results verified the dynamic analysis and matched the simulation results. Figure 1 shows both the structure and working principle for the fishtailing type of the piezoelectric pump (PZT pump for short). The two pieces of piezoelectric ceramics were glued in the front end of the vibrating amplifier. When the piezoelectric ceramics was driven by the alternating current (AC), the piezoelectric ceramics outside will be prolonged in the direction of AA\u2032 while the one inside the paper will be shortened in the DD\u2032 direction, assumed that the voltage outside the paper is positive and inside negative"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002764_speedam.2010.5544902-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002764_speedam.2010.5544902-Figure2-1.png",
        "caption": "Fig. 2. Unequal air gap width due to static eccentricity",
        "texts": [
            "00 \u00a92010 IEEE SPEEDAM 2010 International Symposium on Power Electronics, Electrical Drives, Automation and Motion Static eccentricity between rotor and stator causes an unequal air gap width ( ), which can be mathematically described by the constant air gap width 0, the relative eccentricity and the angle of the eccentricity e [2, 3]: e0 cos1)( (1) The relative eccentricity is described by the static eccentricity e divided by the constant air gap width 0: 0/ e (2) The static eccentricity, leading to an unequal air gap, is shown in Fig. 2. The calculation of the magnetic force Fr, acting with double supply frequency is based on [2, 3], neglecting damping effects due equalizing currents in massive conducting parts, the short circuit rings and the rotor bars. Additionally the magnetic resistance in the rotor shaft, bearings, end shields and stator housing is assumed to be infinitely small here, so that a maximum homopolar flux occurs. Also no saturation of the iron parts and no influence of the slots are considered in the simplified analytical calculation",
            " This static magnetic force does not cause any vibrations. For a small eccentricity ( < 0.2) it can be described by the product of the static magnetic spring constant cm,stat and the eccentricity e [6]. ecF statm,statr, (8) With the assumption of maximum homopolar flux and without considering saturation effects, influence of the slots and field damping effects, the static magnetic spring constant can be described by: 2 p 00 stat m, \u02c6 2 BlRc (9) Due to the excitation of the oscillating magnetic force Fr the centre of the rotor core W (Fig 2), will be forced to move on an orbit. Therefore, the air gap width ( ) will no longer be independent of time ( ) ( , t). This results in a third magnetic force which acts at the rotor, changing its direction with the movement of the rotor, and which tries to magnetize the orbit. This effect can be described by the dynamic magnetic spring constant cm. Without considering field damping effects this dynamic magnetic spring constant is equal to the static magnetic filed constant [1, 6]. stat m,m cc (10) This dynamic magnetic spring constant cm can be subsequently implemented in the rotor dynamic model to take into consideration this dynamic magnetic force"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002504_amc.2012.6197100-Figure6-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002504_amc.2012.6197100-Figure6-1.png",
        "caption": "Fig. 6. Experimental setup.",
        "texts": [
            " Elements of hybrid matrix of this proposed method become (16), (17), (18), and (19), H11 = C2 p ( e\u2212sT \u2212 1 ) \u2212 s2(s2 + 2Cp)( 1 M GH + GLCf ) (s2 + Cp) (16) H12 = \u2212e\u2212sT2Cp s2 + Cp (17) H21 = e\u2212sT1Cp s2 + Cp (18) H22 = \u2212CfGL + GH M s2 + Cp (19) and gain characteristics and phase characteristics of H12(H21) and H11 are shown in Fig. 5. From Fig. 5(b), it turns out that the operationality becomes higher than 2ch control. Additionally, from (13) and Fig. 5(a), it turns out that the stability of control system is better than 4ch control system. In this section, experiments using linear motor are conducted to show the availability of the proposed method. In the experiment, same linear motors are controlled as master motor and slave motor by each methods. Fig. 6 shows the overview of linear motors used in the experiments. The environment is located in the slave motor. A hard aluminum block is used as environment. The control parameters used in the experiment are listed as TABLE II, respectively. In experiments, the pseudo communication delay is made in PC and used, and T1 and T2 are set same. T1 = T2 = 1.0[sec] (20) And jitter of communication delay is set in anywhere from -0.015 sec to +0.015 sec following normal distribution. The communication delay T1 and T2 include jitter used in experiments are shown in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001366_secon.2007.342901-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001366_secon.2007.342901-Figure1-1.png",
        "caption": "Figure 1. In (a,b,c) the camera is fixed to the UAV and motion of the UAV changes the camera orientation. In (d,e,f) the movable camera mount can compensate for UAV movements to keep the camera positioned at a desired orientation.",
        "texts": [
            " The six-degree of freedom rigid body craft is positioned by changing the relative speed of the four rotors. These speed differences can produce torques about the roll, pitch, and yaw axes in addition to the thrust produced as the sum of the four rotating blades. Since the system is inherently under-actuated, there are only four control inputs to directly control the six degrees of freedom, makes performing surveillance and inspection tasks challenging. The inherent problem of attaching a fixed camera to the quad-rotor is illustrated in Figure 1.a, b, and c where it can be seen that the motion of the camera is directly tied to the motion of the UAV. That is, it is not possible to simultaneously specify the attitude of the camera and the attitude of the aircraft. In contrast, the moveable camera depicted in Figure 1.d, e, and f maintains the same orientation regardless of the UAV position. The UAV and camera motions must be coordinated in order to keep the camera pointed in a particular direction. The traditional approach has been to have a pilot position the aircraft about a target and have a camera operator position the camera with the subtask of compensating for motion of the aircraft. A new approach to this control problem was presented in [1] where the UAV and the camera positioning unit are considered to be a single robotic unit"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002400_1.4006324-Figure21-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002400_1.4006324-Figure21-1.png",
        "caption": "Fig. 21 Deformation of the center web gears",
        "texts": [],
        "surrounding_texts": [
            "7.1 Gear Deformation and Discussion. To obtain a good understanding of the tooth root stresses of thin-rimmed gears, the deformation of the gear is increased by 2000 times, and the images of the deformed gears are given in Figs. 20\u201322. Figures 20(a)\u201320(d) show images of deformed left web gears with web angles of 0, 15, 30, and 45 deg, respectively. Figures 21(a)\u201321(d) show images of the deformed center web gears with web angles of 0, 15, 30, and 45 deg, respectively. Figures 22(a)\u201322(d) show images of the deformed right web gears with web angles of 0, 15, 30, and 45 deg, respectively. To understand the deformation characteristics of the loaded tooth of the thin-rimmed gear well, an image of the deformed right web gear with a web angle of 0 deg is shown in Fig. 23(a), and an enlarged view of the loaded tooth is also shown in Fig. 23(b). Figure 23 aids in explaining the deformation characteristics of the loaded tooth of the thin-rimmed right web gear with a web angle of 0 deg in the following discussion. Based on Fig. 23, the tooth positions before deformation and after deformation are sketched in Fig. 24(a). From Fig. 24(a), the deformation of the loaded tooth can be roughly divided into two types of deformation: one type is an upward and downward deformation of the loaded tooth as shown in Fig. 24(b), and the other type is a rotation deformation of the loaded tooth as shown in Fig. 24(c). In Fig. 24(b), end A of the loaded tooth has a downward deformation, and end B has an upward deformation. This is because end A is further away from the web than end B, so end A is more flexible than end B. Thus, end A can be deformed more easily than end B. Therefore, when the tooth is very rigid, an inclined deformation (end A is down and end B is up) of the loaded tooth, as shown in Fig. 24(b), occurs. In Fig. 24(c), when the tooth is loaded, end A moves toward the right and end B moves toward the left. It appears that the loaded tooth rotates around the web (the web is the axis). This deformation is called a rotation deformation of the loaded tooth in this paper. The asymmetrical web position (the web center is offset from Journal of Mechanical Design MAY 2012, Vol. 134 / 051001-9 Downloaded From: http://mechanicaldesign.asmedigitalcollection.asme.org/ on 01/28/2016 Terms of Use: http://www.asme.org/about-asme/terms-of-use the tooth center) can be considered to be the main reason for the tooth rotation deformation. As stated above, because end A is further away from the web than end B, the rigidity of end B is greater than that of end A. Therefore, end B shares most of the tooth\u2019s load as shown in Fig. 8(a). When the tooth is very rigid, the greater tooth load on end B will cause the tooth to rotate around the web, as shown in Fig. 24(c). When the rotation deformation of the loaded tooth occurs, end A of the loaded tooth will approach the neighboring tooth, and end B will move far away from the neighboring tooth on the right side, as shown in Fig. 24(a). This allows the tooth root of the loaded tooth to experience a counterintuitive compressive stress on end A and a tensile stress on end B on the side of the loaded tooth surface. When the web is inclined, the supporting rigidity of the web to the teeth will become smaller. In this case, the deformation of the loaded tooth will be affected by the web\u2019s supporting rigidity, and the deformation of the loaded tooth will become more 051001-10 / Vol. 134, MAY 2012 Transactions of the ASME Downloaded From: http://mechanicaldesign.asmedigitalcollection.asme.org/ on 01/28/2016 Terms of Use: http://www.asme.org/about-asme/terms-of-use complicated than the statements mentioned above. It is necessary to consider the web deformation on the loaded tooth to understand the root stress of thin-rimmed gears with an inclined web. 7.2 Deformation-Sharing Ratios of the Tooth, the Rim, and the Web. A tooth\u2019s relative deformation of a pair of contact gears along the line of action can be calculated by LTCA under the application of a torque load. Table 4 shows the results. In Table 4, the tooth\u2019s relative deformation of a pair of solid gears (the mating gears in Fig. 2(d)) is calculated at the worst load position, and the result is shown in the first column (denoted by \u201csolid gears\u201d). The relative tooth deformations of thin-rimmed left web gears with different web angles are also calculated when the left web gears are engaged with the solid mating gear. The calculation results are shown in the columns denoted by \u201cThin-rimmed left web gear\u201d in Table 4. In Table 4, the deformations are divided into the gear deformation (the total deformation of a pair of gears), the tooth deformation (the deformation resulting only from the contact teeth), and the rim and web deformation (the deformation resulting only from the rim and the web). Therefore, the deformation of the gear is equal to the deformation of the teeth plus the deformation of the rim and the web. In Table 4, the tooth\u2019s relative deformation of the pair of solid gears is 3.8 lm. Because it can be considered that the solid gears have no rim or web deformations, 3.8 lm represents the deformation resulting from the contact teeth. Therefore, in Table 4, the deformation of the teeth is also 3.8 lm. This deformation can be regarded as the tooth deformation of all of the thin-rimmed gears used in this paper when they are engaged with the solid mating gear. The tooth relative deformations of the thin-rimmed left inclined web gears with different web angles are also calculated when they are engaged with the solid mating gear at the worst load positions of the tooth contact. The calculation results are given in Table 4. The deformations of the rim and the web of the thin-rimmed left inclined web gears are obtained by calculating the value of the gear deformations minus the tooth deformation. The deformation-sharing ratio of the teeth can be obtained by dividing the tooth deformation by the gear deformation. The deformation-sharing ratio of the rim and the web can be calculated by dividing the rim and the web deformation by the gear deformation. The calculation results are shown in Table 4. From Table 4, it can be seen that the teeth share only 29.6% of the total deformation of the gears while the rim and the web share 70.4% of the total deformation for the thin-rimmed straight web gear. The deformation-sharing ratio of the teeth decreases while the deformation-sharing ratio of the rim and the web increases when the web angle is increased. When the web angle is increased to 60 deg, the rim and the web share 93.4% of the total deformation of the gears while the teeth only share 6.6% of the total deformation. These findings indicate that the tooth deformation can be neglected in the engineering calculations for thin-rimmed left web gears with large web angles. The deformation-sharing ratios of the center web and right web gears are calculated in the same way, and the calculation results are given in Tables 5 and 6, respectively. From Tables 5 and 6, we find that the same conclusions can be obtained."
        ]
    },
    {
        "image_filename": "designv11_12_0003761_9781118818428-Figure11.2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003761_9781118818428-Figure11.2-1.png",
        "caption": "Figure 11.2 Schematic of sequence for production of internally-slotted exchanger achieving slot width b narrower than that of the tool which originally cut it",
        "texts": [
            " The potential mismatch is more tractable in the case of air- or N2-based designs charged to modest pressure. This permits a choice between fins and tubes which will probably be decided on the basis of manufacturing convenience. Neither is an obvious candidate for mass manufacture, but there is an approach to the production of multiple slots which might change prospects. It follows from observation of production methods for the races of rolling element bearings, some of which are now manufactured by wrapping from strip and butt-welding before finish- machining. Figure 11.2 suggests a sequence for adapting to the generation of internal slots. The machining phase anticipates the final stage which is to extrude the wrapped and welded ring through a tapered die to achieve final external diameter simultaneously with 100 Stirling Cycle Engines predetermined reduction in slot width w. The over-length strip has width equal to eventual slot length and thickness equal to target fin height h plus target pressure-wall thickness t minus a percentage for the increase which will occur during extrusion (expression below)",
            " Regardless of final shape, the slot will have a calculable value of hydraulic radius for use in evaluation of Reynolds number Re. Further deformation occurring during subsequent extrusion may be expected to be confined to circumferential compression between the fin roots, and not to include the fins themselves. This brings about further reduction in effective b and in increase t. A little algebra predicts the percentage reduction in diameter D required to achieve a given target reduction in b. With reference to the notation of Figure 11.2: ns(w + b) = \u03c0D b = \u03c0D\u2215ns \u2212 w Differentiating: bdt\u2215dD + tdb\u2215dD = 0, or dt\u2215t [%] = \u2212db\u2215b [%] (11.2) Some hundreds of tubes per gas path are needed to reconcile wetted-area and hydraulic radius requirements at the compression end. At the assembly stage the sheer number is a problem in itself. Over and above this, failure of a single tube to braze or solder and to form a leak-free assembly renders the exchanger useless. This problem is not inherent to the slotted exchanger. Moreover, a single slot can substitute a number of tubes"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003702_cjme.2013.04.801-Figure5-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003702_cjme.2013.04.801-Figure5-1.png",
        "caption": "Fig. 5. One effective pose of support phase for heavy-duty six-legged robot",
        "texts": [
            " Therefore, when the crab-type tripod gait is used to pass over a slope under condition N, the maximum articulated torques can be respectively obtained for the hip joint and knee joint of the rearmost leg. Similarly, when the ant-type tripod gait is selected to traverse the slope under condition N, one of the abductor joints can be found to experience the maximum articulated torque. joint at rotation angle \u03b8 of 0\u00b0 The crab-type tripod gait is selected for the heavy-duty six-legged robot to pass over the slope. Then, the support phase contains legs 2, 4, and 6 (as shown in Fig. 5), and the transfer phase includes legs 1, 3, and 5. The legs of the transfer phase are ignored in the process of analyzing the articulated torques, but their weights are added to the center of gravity of the bearing platform. Under condition N, only the legs of the support phase are considered when solving the toques for the hip joint and knee joint. Some parameters are defined in Fig. 5. The angle of the slope is defined as \u03b1i, whose value is assumed to range from 0\u00b0 to 34.8\u00b0. Points o2, o4, and o6 are viewed as footholds for legs 2, 4, and 6, respectively; they lie in plane A of the slope. Fz2, Fz4, and Fz6 are the forces for footholds o2, o4, and o6, respectively. Based on the analysis results of section 2, it can be found that the modulus of Fz2 is higher than the modular values of Fz4 and Fz6. Fx2, Fx4, and Fx6 denote the forces in the direction of x; they respectively affect footholds o2, o4, and o6",
            " 2 r l l l l\u03b2 \u03b2 \u03b8       And G is the weight for the robot and rated load, rbp is the effective radius of the bearing platform, cl  is the effective length of the equivalent coxa, tl  is the effective length of the equivalent thigh, sl  is the effective length of the equivalent shin, l is the projection length of the equivalent leg in the direction of x, lGx is the distance from the vector of the G component along the direction of x to foothold o, lGz is the distance from the vector of the G component in the vertical direction from the plane of the slope to foothold o, and lz2 is the distance from the vector of Fz2 to foothold o. According to Fig. 5, when the value of rotation angle \u03b8  is zero under condition N, the foot end forces can be roughly viewed as zero in the direction of y for legs 2, 4, and 6. Based on Fig. 4, the crab-type tripod gait is used by the heavy-duty six-legged robot to pass over a slope of 34.8\u00b0. Then, the maximum static torques, which refer to Mh2 and Mk2, are respectively obtained for the hip joint of leg 2 and the knee joint of leg 2. Based on Fz2 and the right-handed rule for judging the sign of the torque of the joint, a mathematical expression for the articulated torque can be written for the hip joint of leg 2 and the knee joint of leg 2",
            " Meanwhile, the results of a static simulation analysis related to the static torques for the hip joint and knee joint are obtained. The 3D model of the support phase is shown in Fig. 19, and the simulation curves of the static articulated torques are shown in Fig. 20. Fig. 20 shows that the maximum static torque is 992.2 N \u2022 m for the hip joint when the heavy-duty six-legged robot traverses a 34.8\u00b0 slope using the crab-type tripod gait. In Fig. 20, it is also found that the maximum static torque is 471.4 N \u2022 m for the knee joint. In addition, the rearmost leg, which is called leg 2 in Fig. 5, includes the maximum articulated torques on the hip joint and knee joint. Based on the theoretical calculation values for the maximum static torques for the hip joint and knee joint, it is deduced that the theoretical calculation value of the torque of the hip joint is 1.013 0 times its static simulation value. In addition, the theoretical calculation value of the torque of the knee joint is 1.099 9 times its static simulation value. This verifies the reasonableness of the theoretical analysis of the maximum static torques for the hip joint and knee joint"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000664_00423117308968432-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000664_00423117308968432-Figure1-1.png",
        "caption": "Fig. 1. Force and moment acting from road upon tire.",
        "texts": [
            " Usually, these vibrations are most notable near the natural frequencies of the front wheel suspension and steering system. The wheel is connected to the chassis by means of the flexible suspension and steering system. These systems which consist of a complex three-dimensional configuration of links and compliant members are difficult to describe mathematically. However, the pneumatic tire, which forms the connecting medium between wheel-rim and road, probably constitutes the most complicated part to handle. Figure 1 shows the tire in top view. The wheel has a forward speed, V, and the swivel angle is denoted by J / . As a result of the swivel motion, a lateral force, F, and a moment, M , arise which act from the ground via the elastic tire upon the wheel. In vibratory problems, the non-stationary properties of the rolling tire must be considered. A number of classical theories exist which explain and predict the * Extended and revised version of lecture delivered in Sopron (Hungary) October 1971. ** Delft University of Technology (The Netherlands)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003903_tdei.2015.005053-Figure8-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003903_tdei.2015.005053-Figure8-1.png",
        "caption": "Figure 8. Three kinds of organic material sleeves and insulator samples.",
        "texts": [
            " The organic sleeve can alter the corrosion part of zinc sleeve, which will shift the corrosion area from the interface of cement-zinc sleeve to the exposed portion of zinc sleeve where does not bear hoop stress. This method can avoid the decrease of bonding strength between cement and zinc sleeve. The improved disc suspension insulators that installed organic material sleeves are mainly used in the regions with high humidity and serious pollution. Based on the XZP2-300 type porcelain insulators, the samples installed half package type, flush type and whole package type organic material sleeves are designed and manufactured, as shown in Figure 8. The spray water method is used to verify the effect of the organic material sleeve upon change of the corrosion part of the zinc sleeve. The voltage applied on insulator is 1.0 kV and the conductivity and spray velocity are 2-3 mS/cm and 2-3 L/h respectively. During the test process, NaCl solution is sprayed to the surface of insulator to form electrolyte. The experimental set-up is shown in Figure 9. Before the test, the copper electrode should be pasted and fixed onto the lower surface of the insulator"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001460_s12555-010-0211-y-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001460_s12555-010-0211-y-Figure2-1.png",
        "caption": "Fig. 2. Two link robot.",
        "texts": [
            " \u03b3 \u03be \u03be \u03b3 \u03be \u03be \u2212 \u0398 <   \u0398 = \u2265 \u0398 =    \u0398 =   <  i i i i i i i i T f f T f f f T f f f T x B P e M M x B P e P x B P e M x B P e (33) For ,\u0398gij we use: ( ) { } ( ){ } ( ) { } ( ){ } 2 2 2 2 2 2 2 2 2 2 \u02c6 \u02c6 if \u02c6 \u02c6or and 0 , \u02c6 \u02c6 if \u02c6 \u02c6and 0 . \u03b3 \u03be \u03be \u03b3 \u03be \u03be \u2212 \u0398 <   \u0398 = \u2265 \u0398 =    \u0398 =   < ij ij ij ij ij T j gij g T gij g j g T g j g gij T j x B P eu M M x B P e u P x B P e u M x B P e u (34) where the projection operator, [ ]* if P and [ ]* ijg P are In this section, the validity and effectiveness of the proposed controller scheme are examined through the simulation of tracking control for two-link robot manipulator shown in Fig. 2. The objective of the adaptive tracking control design Slim Frikha, Mohamed Djemel, and Nabil Derbel 262 of a robot manipulator is to drive an adaptive control law for the actuator torque u to make the actual trajectories of the robot manipulator with system uncertainties to track the given desired trajectories qd (t) of the joint position and velocity with desired accuracy and stability. In the simulation, we examine the effects of parametric variation on behaviors of the closed-loop control system with the neural networks based adaptive state feedback control, and the proposed observer based adaptive neurosliding controller, respectively"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001678_ac60316a044-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001678_ac60316a044-Figure1-1.png",
        "caption": "Figure 1. Absorbance of iron(I1)-ferrozine formed as a function of pH",
        "texts": [
            " Interference of formaldehyde, nitrous and nitric oxides, hydrogen sulfide, and mercaptans was checked by adding small amounts of these substances in the flow stream or directly into the reaction solution. Reduction of iron(I1I) by sulfur dioxide in the presence of ferrozine in acetic acid buffer depends on pH, the concentration of acetate ion, time, and the molar ratios of iron(II1) to ferrozine. The reduction of iron(II1) by sulfur dioxide in the presence of ferrozine is highly dependent on p H as shown in Figure 1. The pH of solution must be controlled between 3.7 and 4.0, quite contrary to the pH dependence of a similar reduction in the presence of 1 ,lo-phenanthroline. Furthermore, the final acetate ion concentration must be kept in the range of 0.04 and 0.06M as shown in Figure 2. At concentrations of total acetate lower than O.O4M, the p H control is rather poor while a t concentrations higher than 0.06M, the reduction of iron(II1) is incomplete. This is apparently due to the stabilization of iron(II1) by the acetate ion, which is present in approximately a hundredfold excess of iron(II1)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000018_14763140601058409-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000018_14763140601058409-Figure2-1.png",
        "caption": "Figure 2. Representative examples of the trajectory of the bat head in the X\u2013Y plane. (a) The trajectory showing the highest coefficient of cross-correlation (Batter 4). (b) The trajectory showing the lowest coefficient of crosscorrelation (Batter 2) (see Table I). Solid lines show the bat trajectories when hitting a stationary ball. Dotted lines show those when hitting a pitched ball. As the location of the pitched ball varied slightly spatially, these were moved parallel so as to overlap impact points. A circle shows the overlapped impact point.",
        "texts": [
            " For every combination of time and speed in Conditions A and B (3 trials \u00a3 3 trials), the coefficient of cross-correlation was calculated. Then, these nine coefficients for each participant were averaged after Fisher\u2019s z-transformation. The results of the additional experiments suggested that the kinematics of the bat head in Condition B were not significantly different from those in Condition A, when the impact points were close together. The trajectories of the bat head under Condition A and B in the horizontal plane are shown in Figure 2 and those in the sagittal plane are shown in Figure 3. Figures 2a and 3a show the highest coefficient of cross-correlation between both conditions for the speed and each velocity component (Table I). Figures 2b and 3b show the lowest coefficient of cross-correlation (Table I). All coefficients of cross-correlation were significant (P , 0.01). In addition, the speeds of the bat head immediately before impact were not significantly different (32.2 ^ 0.9 m/s under Condition A and 32.7 ^ 1.8 m/s under Condition B)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002522_mesa.2012.6275544-Figure3-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002522_mesa.2012.6275544-Figure3-1.png",
        "caption": "Figure 3. A tripod robot with a tool gripping a cylindrical object.",
        "texts": [
            " It could be found from [30, 32] that each criterion has its own characteristics resulting in some advantages and disadvantages. Accordingly, the provided example in this section is brought to demonstrate and challenge the presented criterion and show its own characteristics which make it efficient for use in an on-line and real-time controller. One of general applications of legged robots is manipulation while walking. Assume there is a hexapod robot which walks based on tripod gait. The robot is equipped with a gripper tool to move objects. Here, a cylindrical rod is chosen to be manipulated as shown in Fig. 3. The robot is commanded to reduce its height, pick the rod up, increase the height, and move it forward. During all these steps, the robot is subjected to a dynamic motion considering velocity and acceleration of CG. The robot is analyzed under such a situation to investigate when the robot loses its stability and tumbles. This scenario is called \u201cLoaded\u201d scenario since the robot picks a load within its motion. Also, there will be another scenario called \u201cNonLoaded\u201d scenario in which the robot has the same motion as the first scenario but without picking a load"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001546_j.jmatprotec.2010.12.002-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001546_j.jmatprotec.2010.12.002-Figure2-1.png",
        "caption": "Fig. 2. The spatial heat distribution in the volumetric heat input model.",
        "texts": [
            " Goldak proposed a double-ellipsoidal heat source model, which has the capability of analyzing the thermal fields of deep penetration welds given above. The power densities of the ellipsoidal heat source, q(x, y, z) describing the heat flux distributions inside heat source and it can be expressed as: q(x, y, z) = 6 \u221a 3fQ abc \u221a e(\u22123(x2/a2))e(\u22123(y2/b2))e(\u22123(z2/c2)) (4) where Q is the energy input rate, f is the fractional factor of the heat deposited in the front and rear quadrant, and; a, b, and c are the heat source parameters. Goldak\u2019s model has been modified in the present work to model laser beam as heat source as shown in Fig. 2. The spatial heat distri- A.F.M. Arif et al. / Journal of Materials Proces b q w l a d d p t d v h r 2 c t C i c m s [ linearly with sin . Microphotonics digital microhardness tester (MP-100TC) was used to obtain microhardness across the weld cross-section. The standard test method for Vickers indentation hardness of advanced ceramics (ASTM C1327-99) was adopted. Microhardness measure- ution in a moving frame of reference can be calculated as follows: = 6 \u221a 3 \u221a \u02dbQ br20 e\u22123(x/r0)2 e\u22123(y/b)2 e\u22123(z\u2212Vt/r0)2 (5) here \u02db is the absorption coefficient; Q is laser power (W); r0 is aser beam radius at focused surface, V is laser welding speed; nd b is heat source parameter which depends on the focal spot ispersion. Fig. 2 shows 3D heat source model, where the power density eposited region is maximum on the top surface along the weld ath, and is minimum at the inner surface. Along the wall thickness, he diameter of the power density distribution region is linearly ecreased. Results, presented by Arif (2010), show that varying arious laser parameters the above volumetric heat source model as the ability to depict actual heating phenomenon in the heating egion. .3. Stress analysis From the principle of virtual work (PVW), a virtual (very small) hange in the internal strain energy (\u0131U) must be offset by an idenical change in the external work due to the applied loads (\u0131V)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001763_robot.2009.5152405-Figure6-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001763_robot.2009.5152405-Figure6-1.png",
        "caption": "Fig. 6. Stability of the proposed model",
        "texts": [
            " If 1 1 )(2 < \u2212 + \u03be \u03b7\u03b6 , the state of the system is unexpected because no pheromone potential field is constructed. The expected condition is achieved when 1 1 )(2 = \u2212 + \u03be \u03b7\u03b6 . Under this condition, the system is stable, and the pheromone potential field is constructed. We verify the parameters given in Fig. 4 by means of simulations based on the results shown in Fig. 5. An artificial pheromone system has to generate pheromone potential field like an expected condition showed in Fig. 5(b) to guarantee the stability of the system. From Fig. 6, we can see that the system is stable and goes on to a steady state if we use the parameters from Fig. 4 that follows 1 1 )(2 = \u2212 + \u03be \u03b7\u03b6 . 5 EXPERIMENT By using an artificial pheromone system composed of data carriers and autonomous robots, as explained in section 3, we conducted an experiment to verify our model by using real hardware. In the experiment, we used data carriers in a random manner. We used one autonomous robot and 25 data carriers for this purpose. Figure 1(a) shows the experimental field employed to realize an artificial pheromone system"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001901_1.3591479-Figure4-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001901_1.3591479-Figure4-1.png",
        "caption": "Fig. 4 Spherical four-bar mechanism with four real cusp positions",
        "texts": [
            " If all 4 equations subsist together, we obtain the maximum number of cusps, which is equal to 4. For example, take k = r = R = 90 deg. This makes cos r = cos R = 0 and A = 0. We then need only to make D in equation (21) equal to zero. This we can do, for example, by taking 0 = 6 = 30 deg and cos c = 2/3, which makes c about 48.2 deg. I11 each case the angle at .1/ is cos\"1 ( \u2014 ' A ) and we obtain 4 real cusps. The coupler curve, moreover, does not degenerate (P3 becomes a multiple of xy and P4 does not contain a factor x or 1/ or {1 + a:2 + i/2J). The cusp positions are shown in Fig. 4. The mechanism is a spherical double rocker. Spherical cardanic motion may also be envisaged as the motion of a link, two points of which are constrained by spherical guides, which are great circular arcs (equator and meridian). This corresponds to the case r = R \u2014 k = 90 deg. The coupler curve is a good deal simpler than in the general case, but remains of order 8 and genus 1. Skew Four-Bar Mechanisms (2 Turning Joints, 2 Spherical Jo ints ) Case (i) : Two Turning Joints F ixed. T h e re lat ion b e t w e e n the crank rotations is well known [4, 15, 16], The mechanism is shown in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001095_j.patrec.2009.08.004-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001095_j.patrec.2009.08.004-Figure1-1.png",
        "caption": "Fig. 1. A view of the adidas_1 shoe, depicting the cushioning element and motor unit. The indicated magnet induces a magnetic field for compression measurement.",
        "texts": [
            " This can thus also be regarded as a decision for either a \u2018control\u2019 or a \u2018cushioning\u2019 condition and a corresponding complete adaptation of the shoe. This automatic adaptation ideally takes into account the athlete\u2019s weight, speed, fatigue level and furthermore the current surface condition, elevation profile and shoe condition. To facilitate this adaptation, the shoe features a cushioning element, whose ability to give way in vertical direction (hereafter defined as z-axis) can be regulated by a motor-driven cable system. The cushioning element is depicted in Fig. 1. The regulating cable is visible in the z-axis X-ray image of the adidas_1 in Fig. 2. It is running from the motor through the middle of the cushioning element to its opposite end and is fixated there. The motor shown in Fig. 1 can adjust the attenuation setting by turning a screw which lengthens or shortens the cable. When the cable is shortened, the cushioning element is tensed and compresses very little when external forces are applied. When the cable is longer, it allows the cushioning element to compress further by giving it more room to expand in the x-axis direction (forward\u2013backward direction), effectively making the shoe softer. Changes to the softness setting are gradual. The attenuation setting from one extreme to the other is made in 15 increments",
            " when running over a small stretch of grass while being mainly on a hard sidewalk surface. Using this approach, we can ensure that the battery (see also next Section 2.2) holds for the complete life-time of a running shoe, which is about 100 h. For more details on the shoe design the reader is referred to DiBenedetto et al. (2004) and DiBenedetto et al. (2005). The compression measurement of the adidas_1 shoe is made by a hall sensor that is mounted at the top of the cushioning element. It detects the magnetic field strength induced by a small magnet, see Fig. 1, and can be sampled with a rate fs of up to 1 kHz. The sensor-magnet distance dm can then be computed from the magnetic field strength with an accuracy of 0:1 mm. A decision whether the attenuation of the shoe has to be adapted is made based on the measured sensor data, see Section 2.3. The sensor-magnet distance is sampled by the built-in microprocessor that is mounted on a flexible circuit board on the motor element. Currently, a Cypress Semiconductor Corporation controller CY8C21634 is used. However, the methodology that is presented below does not make any special requirements to the employed Microprocessor"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003713_978-1-4614-2419-2-Figure39.7-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003713_978-1-4614-2419-2-Figure39.7-1.png",
        "caption": "Fig. 39.7 (a) FE model of the fully bonded (reference state), (b) FE of first Damage scenario of FRP beam",
        "texts": [
            " Damaged beams are modelled with different progressive damage at the 0.5 L (middle of the beam). The damage in FE model was done by merging two adjacent meshed regions that have coincident nodes and keypoints, always merging nodes before merging keypoints. Merging keypoints before nodes can result in some of the nodes becoming disconnected and modelling damage or crack; that is, the nodes lose their association with the plates attached to the shell model indicated as blue part of the model, as shown in Fig. 39.7. The keypoints are merged and any higher order shell model entities (e.g. area) attached to those keypoints are considered for merging. Initial FE Model Results. Table 39.7 shows the finite element model of the undamaged state or reference state. The damage in the FE shell model was simulated by merging two adjacent meshed regions (as indicated in Fig. 39.7b) that have coincident nodes and keypoints. The comparison of experimental and numerical approaches is shown in Table 39.7, where the differences of calculated frequencies from both cases are less for lower modes. The comparison of the natural frequencies between FE model and the experimental result demonstrated the needed of the model updating process. 39 Application of FE Model Updating for Damage Assessment of FRP Composite Beam Structure 391 The model updating method used in this study is classified into two manual updating (MU) and automatic updating (AU)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002773_inista.2011.5946063-Figure3-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002773_inista.2011.5946063-Figure3-1.png",
        "caption": "Figure 3. ANN used in the observer schemes.",
        "texts": [
            " However this observer has some drawbacks caused by the discontinuous term ( )iesign ~ and the high value of the correction gain \u03b4i which can create a very large amplification in measurement noise and therefore a serious damage in global system. Several solutions in literature have been proposed to solve this problem where the basis idea is to minimize the observer gain (\u03b4i) whenever the observation error bring closer to zero [8, 9] or to use a smoother functions as sigmoid function to replace the sign function [10]. However, these methods have limitations especially if the system requires a very important value of gain observer (\u03b4i) as the case of quadrotor helicopter. To avoid this problem the ANN (Fig. 3) is used to replace the ( )ii esign ~\u03b4 in the sliding mode observer. The ideal ANN observer proposed in this paper is given by: ( ) ( ) piECeeFEBAE iiiiiiiiiii :1,\u02c6\u02c6,~\u02c6\u02c6 1 0 ==\u0398\u0394+\u03a6\u2212= \u2217\u2212 (25) where ( )ieF ~\u2217 is as follows: ( ) ( ) ( )[ ] pieobseobseF ir iii :1,~~~ =\u211c\u2208= \u2217\u2217\u2217 (26) and: ( ) ( ) piTeobs Oii T iii :1~ , =+\u0395\u0391=\u2217 \u03b5\u03bc (27) with : [ ]Tiii 2,1, \u03b1\u03b1=\u0391 , [ ]Tiii ee \u02c6=\u0395 and Oi,\u03b5 is the reconstruction error, the activation function T(.) is a saturation or a shifted sigmoid function, which can be chosen as the following function: ( ) z z e ezT 2 2 1 1 \u2212 \u2212 + \u2212= (28) Assumption 5 : Ai and \u03bci are bounded : mii ,\u0391\u2264\u0391 and mii ,\u03bc\u03bc \u2264 , with mi,\u0391 and mi,\u03bc are unknown positive constants"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002129_s0263574709005426-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002129_s0263574709005426-Figure2-1.png",
        "caption": "Fig. 2. Closure error in a four bar chain.",
        "texts": [
            " The simplest technique is based on the Euler integration method; given an integration interval t , if the joint positions and velocities at time tk\u22121 are known, the joint positions at time tk = tk\u22121+ t can be computed as qs(tk) = qs ( tk\u22121 ) \u2212 N\u22121 s ( tk\u22121 ) Np ( tk\u22121 ) q\u0307p(tk) t. (5) By calculating the secondary joint positions using Eq. (5), a cumulative error in qs is introduced. Therefore, Eq. (1) is not satisfied and an opening in the closed-loop chain is introduced. To illustrate this, consider the four-bar mechanism in Fig. 1. The cumulative error in q\u0307s opens the closed chain, as depicted in Fig. 2. To solve this problem, we present in Section 4 a new method to model the robot differential kinematic equation, where the position is obtained using numerical techniques in which the closure error converges exponentially to zero. To describe it, first we present the fundamental kinematics tools used in this study. Our approach is based on the method of successive screw displacement,8 on the screw representation of differential kinematics, on the Davies method and on the Assur virtual chain concept, which is briefly presented in this section",
            " (13)), and if the joint has the opposite direction to the circuit, the sign will be negative. In the example, the twist R$j , associated with the joint j, will have a positive sign in the circuit a equation and a negative sign in the circuit b equation. An integration algorithm is necessary to integrate the differential kinematics equation to obtain the joint positions.19 The algorithm proposed in this paper has two steps. The first step is to introduce a virtual chain to represent the closure error resulting from the integration error as shown in Fig. 2. For the same example of a four bar planar mechanism, the resulting closed chain is shown in Fig. 10. The constraint equation of this closed-loop chain results in Npq\u0307p + Nsq\u0307s + Neq\u0307e = 0, (18) where Np and Ns are the primary and secondary network matrices obtained by integration, q\u0307p and q\u0307s are the primary and secondary magnitude vectors, respectively, Ne is the error network matrix and q\u0307e is the error magnitude vector. The second step is to replace Eq. (3) by q\u0307s = \u2212N\u22121 s Npq\u0307p + N\u22121 s NeKeqe, (19) where the gain matrix Ke is chosen to be positive definite and qe is the position error vector",
            " (22) As in a closed-loop chain, the first and the last links are the same, and the orientation and position of a link with respect to itself are given by a homogeneous matrix equal to the fourth-order identity matrix. In a closed-loop chain with np primary joints and ns secondary joints (Eq. (22)), the closed-loop equation results in np\u220f i=1 [Ap]i ns\u220f i=1 [As]i = I, (23) where [Ap]i , i = 1. . .np are the homogeneous matrices corresponding to the primary joints, and [As]i , i = 1 . . .ns are the homogeneous matrices corresponding to the secondary joints. Consider a closed-loop that has an error chain as shown in Fig. 2. As in ref. [6], we represent the closure error with a homogeneous matrix E, and the closed-loop equation becomes { np\u220f i=1 [Ap]i ns\u220f i=1 [As]i } E = I. (24) The closure error is calculated by E = { np\u220f i=1 [Ap]i ns\u220f i=1 [As]i }\u22121 = [ Re pe 0 1 ] , (25) where pe = [pex pey pez]T is the position error vector and Re is the rotation matrix error. The matrix Re corresponds to errors measured in rex , rey and rez virtual rotative joints considering their structural conception. The \u2018position\u2019 error (which is a posture error involving position and orientation) is given by the position error vector qe = [rex rey rez pex pey pez]T "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000470_j.ijfatigue.2007.12.003-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000470_j.ijfatigue.2007.12.003-Figure2-1.png",
        "caption": "Fig. 2. The configuration of test specimen.",
        "texts": [
            " In order to evaluate the fatigue life according to the depth of surface removal, a set of finite element analysis was performed by sequentially removing elements along the depth direction. Resulting strain and stress data obtained from this analysis was used for the corresponding fatigue analysis. The residual stress and micro hardness were measured by applying the hole drilling method and compared with those from finite element analysis results. In order to evaluate the fatigue life for various surface removal depths, a set of contact fatigue test were conducted. Fig. 2 shows the configuration of test specimen simulating the contact between wheel and rail [9,10]. The wheel specimen has 90 mm in diameter and 15 mm in thickness, and the rail specimen has 110 mm in diameter and 15 mm in thickness. The contact surface thickness of rail specimen was set to 5 mm to maintain a constant contact stress regardless of wear. In order to keep the constant hardness of specimen on contact surface, the heat treatment process was applied. The contact surface was grinded to simulate the surface roughness of railway wheel"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000264_3-540-36224-x_4-Figure3-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000264_3-540-36224-x_4-Figure3-1.png",
        "caption": "Fig. 3. A three-revolute-joints device. It can be proven [9] that any two-actuator configuration of this system is kinematically controllable, i.e., one can always find two decoupling vector fields whose involutive closure is fullrank.",
        "texts": [
            " Both methods are consistent, complete and constructive (consistent planners recover the known solutions available for linear and nilpotent systems, and complete planners are guaranteed to find a local solution for any nonlinearly controllable system). The following decoupling methodology was proposed in [9] to reduce the complexity of the motion planning problem. The method is constructive (only quadratic equations and no PDEs are involved) and physically intuitive. We consider as a motivating example a common pick-&-place manipulator: Fig. 3 shows a vertical view of a three-revolute-joints device. We investigate planning schemes for this system when one of its three motors is either failed or missing. We present a decoupling idea to reduce the complexity of the problem: instead of searching for feasible trajectories of a dynamic system in R6, we show how it suffices to search for paths of a simpler, kinematic (i.e., driftless) system in R3. A curve \u03b3 : [0, T ] V\u2192 Q is a controlled solution to equation (1) if there exist inputs ua : [0, T ]\u2192 R for which \u03b3 solves (1)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000597_978-3-540-73129-0-Figure6.42-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000597_978-3-540-73129-0-Figure6.42-1.png",
        "caption": "Fig. 6.42. Multi-mode magnetostrictive FLEX-M1 motor",
        "texts": [
            " Therefore, unlike piezoelectric motors, this motor cannot operate at resonance. As a consequence and in relation to the previous analysis of power (Fig. 6.34), the efficiency is comparatively weak. Its other characteristics are a speed of 40\u25e6/s and a torque of 1.8Nm [62]. It is difficult to convert existing piezo-motors to magnetostrictive versions; new designs have to be found. A first magnetostrictive motor using the mechanical resonance of two vibration modes has been built and tested by Cedrat Recherche [63] (Fig. 6.42). Its stator modules are made of a ring and two Terfenol-D linear actuators. The translation mode of the stator produces a vibration that is tangential to the contact zone (Fig. 6.45a). The flexure mode produces a vibration that is normal to the contact zone (Fig. 6.45b). These modes are coupled using a 90\u25e6 phase shift, in order to produce ellipti- cal vibrations (Fig. 6.43) that are used to transmit a motion to two rotors by friction. A low rotating speed of 100\u25e6/s, and a torque of 2.1Nm are achieved (Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002218_j.ymssp.2010.12.001-Figure4-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002218_j.ymssp.2010.12.001-Figure4-1.png",
        "caption": "Fig. 4. Schematic of the whole-machine balancing of horizontal decanter centrifuge: (1) process spindle, (2) photoelectric sensor, (3) reflective flake, (4) vibration sensor, (5) photoelectric sensor, (6) reflective flake, (7) conveyor(inner rotor), (8) bowl(outer rotor), (9) vibration sensor, (10) differential mechanism and (11) pulley.",
        "texts": [
            " Therefore, an accurate result without error can be obtained. According to the Whole-Beat Correlation Theory, the whole-machine balancing instrument has been successfully developed based on computer for the dual-rotor system with a slight speed difference. With the application of this device, the field balancing experiment was conducted on the horizontal decanter centrifuge, which is a typical dual-rotor machine in industry, and the satisfied vibration results validated the effectiveness of this instrument. Fig. 4 is the schematic of the wholemachine balancing experiment. The photoelectric sensor 2 and reflective flake 3 are used to sample the key-phase signal of the conveyor (inner rotor). Similarly, the photoelectric sensor 5 and reflective flake 6 are combined to collect the key-phase of the bowl (outer rotor), and the vibration sensors 4 and 9 can acquire the vibration signals of left and right bearing, respectively. The sampled key-phase signals and vibration signals are transmitted into the computer via a data sampling box, and then are processed according to the whole-beat correlation algorithm"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002720_00423114.2011.602419-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002720_00423114.2011.602419-Figure1-1.png",
        "caption": "Figure 1. The scheme of wheel\u2013rail contact.",
        "texts": [
            " According to the algorithm, we assume that the tangential stress depends linearly on the tangential strain; such \u2018tangential Winkler\u2019 model is correct enough for the wheel\u2013rail contact. The main disadvantage of the deterministic model is its very large calculation time. In the following sections, both methods for finding contact forces, relative position of rail and wheel and a value of slippage at the contact zones are used to study the two competitive fracture mechanisms in wheel\u2013rail contact, namely wear and damage accumulation. 2.2. Formulation of the contact problem Let us consider the wheel\u2013rail interaction (Figure 1). The first coordinate system O(x0, y0, z0) is coupled to a wheel. It is usually accepted that the axis Ox0 is parallel to the axis of the wheelset. The second (moving) coordinate system is connected with the initial contact point. It is obtained from the first system by rotating in the plane Ox0, z0 (the angle of rotation \u03b1 is the angle of attack for the \u2018wheel system\u2019), parallel displacement in the plane Ox0y0 and by rotating in the plane Ox0y0 by the angle, which is determined by the position of the point of contact. The axis Oy of the second coordinate system coincides with the common normal to the contacting surfaces at the point O of the initial contact; the direction of the axis Oz coincides with the direction of the motion and Ox is directed inside the tread (see Figure 1). The initial shape of the contacting surfaces at the moment t is determined by the functions y = f1(x0, t) and y = f2(x0, t) of the wheel and rail profiles, respectively. One-point and two-point rail/wheel contacts are considered: the one-point contact occurs at the tread (point A, Figure 1), and the two-point contact realises at the tread (point A, Figure 1) and at the wheel flange (point B, Figure 1). For the determination of the contact pressure distribution at any fixed moment of time t = t\u0303, we have a system of equations, which consists of contact conditions, relation between the elastic displacements and the contact pressure and the equilibrium condition. The following contact condition in the local system of coordinates is considered: w1(x, z, t) + w2(x, z, t) + h(x, z, t) = \u03b4, (1) where w1(x, z, t) and w2(x, z, t) are the elastic displacements of the surfaces of the wheel and rail, respectively, \u03b4 is the approach of solids under the load and h(x, z, t) the gap between the surfaces before loading"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001866_bf00251592-Figure15-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001866_bf00251592-Figure15-1.png",
        "caption": "Fig. 15. Cusp position of Stephenson-2 mechanism with a symmetrical six-bar curve generated by point G",
        "texts": [
            " However, this is possible in the special case, because the equation corresponding to (2), with ~0' instead of ~, becomes independent of t, so there are only three equations for X, Y, u, instead of four. 9. Cusps The general condition for a cusp is that G lies at the intersection of the fines through A F and ED. As in the earlier analysis of the Watt-1 and Stepheuson-1 six-bar curves, we confine our discussion to the case of a symmetrical Stephenson-2 Six-Bar Motion. II 67 curve (sin c~=sin ct'=0). For the symmetrical Stephenson-2 curve, the cusp configuration is shown in Fig. 15. This figure differs from the corresponding figure (Fig. 8) for the Stephenson-1 six-bar curve essentially only in that the triangular link on the right leg of triangle A GE, is floating. The cusp configuration in Fig. 15 is defined by the collinearity of A B F G and E D G C . Algebraically, it is defined by equating cos 2 in triangles G B C and GAE. The four cusp configurations which, when subsisting together, yield an eight-cusped, symmetrical Stephenson-2 curve are defined by the proportions R~ R 2 s' + m s + l s ' + m [ s - l l I s ' - m l s + l I s ' - m l [ s - I [ . Carrying out the associated algebra, we find that the conditions necessary for the existence of an eight-cusp curve become d '=s ' +_m , d=s+_n, n= m (s'-T-m), S l' 2=(s'-T-m)2-s2 + l 2"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000542_pesc.2008.4592377-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000542_pesc.2008.4592377-Figure2-1.png",
        "caption": "Figure 2. The stationary \u03b1-\u03b2, the synchronous rotating d-q and the estimated \u03b3-\u03b4 reference frames.",
        "texts": [
            " Flux/Current Observer of PMSM In sensorless position/speed control, the rotor position cannot be detected, and therefore d-q axis mathematical model cannot be applied directly. Most approaches are based on the estimation of the back electromotive force (EMF) in the stationary reference frame \u03b1-\u03b2. The proposed PMSM mathematical model of sliding mode observer is reflected in an estimated reference frame \u03b3-\u03b4 rotating at an estimated angular velocity \u03c9\u0302 and lagging behind the d-q reference frame by electrical angle error \u03b8 . Fig. 2 shows the relations between the synchronous reference model (dq-axis) and the estimated reference model (\u03b3\u03b4-axis) used in this study. The mathematical model of PMSM in dq-axis synchronous rotating reference frame is presented by the following flux/current state equations. d s d d qr i u\u03bb \u03c9 \u03bb\u22c5= \u2212 + + \u22c5 (1) q s q q dr i u\u03bb \u03c9 \u03bb= \u2212 + \u2212 \u22c5i (2) d d d mL i\u03bb \u03bb\u22c5= + (3) q q qL i\u03bb \u22c5= (4) The conventional d-q axis model can be transformed to \u03b3-\u03b4 axis as follows (see Appendix A). ( )sr i u\u03b3 \u03b3 \u03b3 \u03b4\u03bb \u03c9 \u03b8 \u03bb= \u2212 + + \u2212 (5) ( )sr i u\u03b4 \u03b4 \u03b4 \u03b3\u03bb \u03c9 \u03b8 \u03bb= \u2212 + \u2212 \u2212 (6) d mL i\u03b3 \u03b3 \u03b3\u03bb \u03bb\u22c5= + (7) q mL i\u03b4 \u03b4 \u03b4\u03bb \u03bb\u22c5= + (8) The partial fluxes m\u03b3\u03bb and m\u03b4\u03bb are functions of m\u03bb , \u03b8 and 2L , where ( )2 / 2q dL L L= \u2212 "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001906_10402000802687890-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001906_10402000802687890-Figure1-1.png",
        "caption": "Fig. 1\u2014Schematic of cylindrical roller bearing.",
        "texts": [
            " In view of the aforementioned, the objectives of this work were to (a) independently determine the lives of the inner races, outer races, and roller sets for several classes of radially loaded, cylindrical roller bearings subject to inner-ring interference fit; (b) calculate the reduction in cylindrical roller bearing fatigue life due to the interference fit of the inner ring; and (c) develop life factors applied to the bearing life calculation for the interference fits according to the ANSI/ABMA standards for shaft-fitting practice. A representative cylindrical roller bearing is shown in Fig. 1. The bearing comprises an inner and outer ring and plurality of rollers interspersed between the two rings and positioned by a cage or separator. Figure 2(a) is a schematic of the contact of a cylindrical roller on a race. Figure 2(b) shows the principal stresses at and beneath the surface. From these principal stresses the shearing stresses can be calculated. Four shearing stresses can be applied to bear- ing life analysis: the orthogonal, the octahedral, the von Mises, and the maximum. For the analysis reported herein, only the maximum shearing stresses are considered"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003713_978-1-4614-2419-2-Figure53.3-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003713_978-1-4614-2419-2-Figure53.3-1.png",
        "caption": "Fig. 53.3 BTE Super Hauler wing payload pod showing the mounting hardware and aerodynamic shell",
        "texts": [
            " But, although it is easily constructed and repaired, the resulting complex framework is very difficult to model for a structural analysis. Thus, the relatively complex framework made of materials with varying mechanical properties is best analyzed using an experimental approach. The Super Hauler was custom designed to provide a large unobstructed payload bay in which to mount multiple payloads for flight testing. While this capability has met the needs of the previous payloads tested, the addition of wing payload pods is necessary to conduct new missions and payload evaluation. The payload pod designed is shown in Fig. 53.3. The pod mounts at the intersection of the wing segments with a structural mounting hardware that follows the contour of the wing at this location. The outside of the pod is made of polycarbonate formed into an aerodynamic shape to reduce drag and lessen the impact on flight performance. The overall dimensions of the pod payload are 7.75 in. wide by 11 in. long by 2.5 in. deep and it is designed to carry up to 5 lbs. The wing payload pods will most often be used in pairs to equalize the wing loading and provide a symmetric load on the airframe"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000837_tcst.2006.886438-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000837_tcst.2006.886438-Figure1-1.png",
        "caption": "Fig. 1. Schematic of M3.",
        "texts": [
            " To extend the idea of fuzzy Lyapunov synthesis to a multiinput multi-output (MIMO) system and its application to an M3 system, this brief presents a methodology that uses only the output relative degree and the structural properties of the system model where the latter can be obtained directly from the physical laws. No other a priori knowledge about the system is assumed. The rest of this brief is organized as follows. First, the dynamic model of an M3 is given in Section II. In Section III, the proposed controller is presented and experimental results are given in Section IV. Finally, Section V concludes this brief. The schematic of the M3 test-bed with one flexible macro link and two rigid micro links is shown in Fig. 1. The dynamic equations of a planar M3 when the macro manipulator joints are locked, are given in (1) [13], [14]. Hereafter, this configuration is simply called \u201clocked-joint\u201d configuration (1) 1063-6536/$25.00 \u00a9 2007 IEEE where . and are the vectors of micro joint variables and (truncated number of) flexible modes, respectively, and represent Coriolis and centripetal forces and and are gravity forces, is the joint friction, is the structural damping term, and and are the positive-definite mass and stiffness matrices, respectively"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000894_iros.2008.4651029-Figure5-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000894_iros.2008.4651029-Figure5-1.png",
        "caption": "Fig. 5. From the top, left: the basic configuration, on its right the positioning of rotative actuators, then the positioning of linear actuators between the two passive joints and between two controlled joints. All the configuration are referable to the first form presented through using simple geometric transformations.",
        "texts": [
            " These remarks ideally allow to substitute in a functional analysis the two series of connected four-bar mechanisms with a single kinematic chain and to overlook the presence of the isostatic triangle. The assumption of a parallel architecture results therefore completely justified. The following table 1 presents the basic nomenclature adopted in this paper. B. Functional simplifications Through considering the simplified structure in figure 4, the meso-manipulator architecture presents 132 dof and 129 doc [26-30] The three required actuators can be indifferently imposed following the four possible interchangeable configurations evaluated in the manipulator analysis and presented in figure 5. The conventions adopted for the static and dynamic analyses of the manipulator are shown in figure 6. III. KINEMATIC ANALYSIS The position of the platform centre results univocally defined by a tern of variable distances, i.e. once the position of the feet has been established. By referring to figure 6, the geometry of the structure fixes PiBi and Miptt, with i = 1, 2, 3 representing every leg. These values are time-invariant, so that no additional contributes are introduced into the kinematic analysis"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001256_ie0708881-Figure9-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001256_ie0708881-Figure9-1.png",
        "caption": "Figure 9. Time evolution of TOC/TOC0 during the ozonation of an aqueous mixture of polyphenols in the presence and absence of activated carbon. Conditions as in Figure 2. Black bars correspond to ozonation in the presence of activated carbon.",
        "texts": [
            " As is deduced, the order of reactivity of these polyphenols with ozone is similar to that observed in Figure 2 for the case of individual ozonations: gallic acid > syringic acid > tyrosol. Again, the presence of activated carbon reduces the content of polyphenols compared to the activated carbon free ozonation. After 20 min of reaction, both gallic and syringic acids are nearly completely removed (about 95% removal) and only tyrosol remains in water (70 and 83% removals in the absence and presence of activated carbon, respectively). If TOC is considered, as shown in Figure 9, the presence of activated carbon results in significant decreases of this parameter. Thus, after 180 min, 52 and 88% removals are achieved in the absence and presence of activated carbon, respectively. 3.3. Ozone Consumption. Ozone consumption is a fundamental parameter to establish the suitability of the ozonation process. In this work, ozone consumption has been calculated in both individual and mixture ozonations of polyphenols as two forms: as the mass ratio between ozone consumed and TOC removed and as the mass ratio between ozone fed and TOC removed",
            " The first one is the removal of initial polyphenols which, in accordance with their high ozone reactivity and low influence of activated carbon presence, is likely due to fast direct ozone reactions (in fact, none or very little dissolved ozone was noticed during this initial period of reaction, thus confirming fast ozone reactions). Ozone reactions break the aromatic rings and double bonds of first unsaturated carboxylic acids formed and yield hydrogen peroxide6 and saturated carboxylic acids with reductions of pH in cases of weak buffer ionic strength. Thus, oxidation of saturated car- boxylic acids mainly due to free radical reactions is the second mechanism of the ozone process, which is particularly important when activated carbon is present (see Figure 9). In the absence of activated carbon, the only way for free radical formation is the reaction between ozone and hydrogen peroxide12 but this reaction is favored with pH increase. Since pH diminishes because of carboxylic acids formation, free radical oxidation is rather inhibited without activated carbons. In the presence of activated carbon, however, free radical formation is also likely due to decomposition reactions of hydrogen peroxide and ozone on the carbon surface,1 which explains the higher mineralization achieved"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003973_s12206-012-0811-y-Figure8-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003973_s12206-012-0811-y-Figure8-1.png",
        "caption": "Fig. 8. The deformed shape.",
        "texts": [
            " Material properties required for this analysis are modulus of elasticity, Poisson\u2019s ratio and density. The material properties of various elements of the analysis are listed in Table 1. The next step is to define loads and constraints required for the analysis. To account for an inertia effect like gravity, appropriate values for g (9.81 m/s 2 ) are given. All the degrees of freedom of elements in that portion of the crank shaft which is seated on main bearings have been constrained. Fig. 7 shows the loads and constraints applied on the finite element model. Fig. 8 shows the deformed shape of the assembly. The deformation legend diagram is given (Fig. 9). The maximum de- formation occurs at the piston top surface and it is equal to - 0.04812 mm (along downward direction). Fig. 10 shows the von Mises stress legend diagram. The tolerance allocation of the piston - cylinder assembly is carried out as follows. The assembly parts are listed in Table 2. The associated tolerances T1, T2, T3, T4, T5, T6 and T7 must be determined so that the clearance x\u03a3, between the piston top surface and cylinder top surface must fall within the functionality limits, 0"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001370_icsens.2009.5398292-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001370_icsens.2009.5398292-Figure1-1.png",
        "caption": "Fig. 1. Flexible displacement sensor attached on the McKibben actuator.",
        "texts": [
            " In his research, pressure-length characteristic of MPA was approximated by a linear function. Wakimoto developed an intelligent McKibben actuator with a build-in flexible electroconductive rubber sensor, which can measure the length by actuator itself [7]. However, for gaining the displacement from the flexible sensors, axial tension is need in default position. In this research, a flexible electro-conductive rubber sensor has been developed to estimate the length without losing the flexibility and lightweight properties of MPA. Figure 1 shows the developed flexible sensor. This flexible sensor can be used to measure the circumference displacement. A estimation method was also been proposed to get the axial displacement from the estimated circumference displacement. This flexible sensor has 4 major inherent advantages: (1) Since the sensor can be equipped in MPA directly, power-assisting device does not need exoskeleton to install rigid sensors, e.g. sliding and rotary potentiometers. (2) Higher accuracy and S/N ratio can be obtained by measuring the circumference displacement instead of directly measuring the axial displacement",
            " The sensor is made from electro-conductive rubber, which is flexible and lightweight just like the MPA. Therefore, using this flexibile sensor, the flexibility and lightweight properties of MPA will not lost. 978-1-4244-5335-1/09/$26.00 \u00a92009 IEEE 520 IEEE SENSORS 2009 Conference Figure 2 shows a flat rubber sensor. Two conductive wires are set in each end to measure the resistance of electroconductive rubber. The deformation between these two contact points can be calculated by measuring the change of resistance. This sensor can be winded around the MPA (Fig. 1) to measure the circumference displacement. Using a model of the geometric structure of the MPA, the axial displacement can be estimated from the circumference displacement measured from this flexibile sensor. The hardness of flexible electro-conductive rubber is about 15 (ASKER C in SRIS0101 1), which is more flexible than the rubber used in MPA. The weight is about 10[g], which is lighter than MPA 2. Therefore, MPA will not be affected heavily by the flexible sensor. Table I shows the specification of electro-conductive rubber using in the flexible sensor",
            " All 5 MPAs have same default circumference (\u03c615[mm]). These MPAs will also be used in following experiments. Figure 7 shows the relationship between the length and the circumference of each MPA. Table II shows the estimated length b and turns n of the fiber of each MPAs. The red lines in Fig. 7 show the estimated length L using the estimated b and n in (8). The estimated lengths are closed enough with the measured results and the values of b and n will also be used in next experiments. The flexible sensor is installed to MPA as shown in Fig. 1. The validity of the MPA length estimation method is examined by comparing the estimated length and the true length. Figure 8 and 9 show the experiment devices and the configuration of control system. Control system consists of an air compressor, an air regulator, a control computer, an A/D convertor, a D/A convertor, a DC power source, a MPA and a flexible sensor. The estimation length is caluclated from the circumference displacement measured by the flexible sensor using (8). The true value is measured by the potentiometer"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001931_iros.2010.5649323-Figure18-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001931_iros.2010.5649323-Figure18-1.png",
        "caption": "Fig. 18. The Flow Field of Fig. 15 with the new bj(z)",
        "texts": [
            " However, the velocity of the mobile robot accelerates rapidly when it avoids a moving obstacle. Fig. 16 shows the simulation with the new method by using CT. That is, the ellipse flow field is used in this simulation. The initial conditions of Fig. 16 is the same as those of Fig. 15. In comparison of Fig. 15, Fig. 16 and Fig. 17, the difference of their velocities is shown. It can be seen that the maximum velocity of the new method is a half maximum velocity of our previous method. As a result, the effectiveness of our method by using the ellipse flow field can be shown. Fig. 18 shows the simulation with applying the correction function bj(z) in Fig. 15. And Fig. 19 shows the simulation with applying correction function bj(z) in Fig. 16. Similarity, Fig. 20 also shows the difference of their velocities. It can be seen that the case of Fig. 19 is smoother than the case of Fig. 18. The blue line as shown in Fig. 17 has a discontinuous point at the maximum velocity point. However, the blue line as shown in Fig. 20 has no discontinuous point at all time. As a result, the effectiveness of our method by using the new correction function can be shown. Is the robot motion considered? If the robot is just in front of the obstacle, how does the robot move? These answers are very simple. That is, the robot can move while touching the avoidance circle of the obstacle. If the obstacle does not move, the robot can always avoid the obstacle",
            " 20 come from both the ellipse field and the correction function, and they are only influenced by the parameters of m and uj . However, the m and the uj are the same value, respectively in each simulation. The m denotes a sink value and the standard robot velocity comes from the m. That is, the m can not be changed. This can be seen by the vertical constant velocity lines shown in Fig. 17 and Fig. 20. On the other hand, the uj denotes the obstacle velocity. That is, the uj can not be changed too. All distances of many circles of the moving obstacles shown in both Fig. 15 and Fig. 16, and also both Fig. 18 and Fig. 19 are the same, respectively. Therefore, the changing velocities of these simulations come from the other factors. The factors are using both the ellipse field and the correction function. In this paper, we proposed the improved method by applying the Conformal Transformation and the new correction function to our previous Hydrodynamic Potential method for path planning of a mobile robot to avoid the moving obstacle smoothly. A mobile robot can gradually avoid a moving obstacle from further away, and can be safely guided without rapid acceleration"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000506_j.engfracmech.2008.05.004-Figure3-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000506_j.engfracmech.2008.05.004-Figure3-1.png",
        "caption": "Fig. 3. Load distribution of spur gear, (a) meshing points of spur gear [14\u201318], (b) load distribution along meshing line.",
        "texts": [
            " The aim of the modification is to delay the pitting formation at the single tooth meshing region by means of decreasing the effect of the Hertzian surface pressure in the region caused by a decrease in load sharing, which is a result of an increase in tooth width at the single tooth meshing region. To avoid the effect of surface pressure on gear pitting, a gear width modification was made, thus enabling a homogeneous distribution of the surface pressure. Because the modification of gear width is based on the fact that those areas with great loads are widened, thereby decreasing the surface contact pressure, it is important to first determine the meshing zones and specify the load distribution along the meshing line. The distance AE in Fig. 3 is the meshing distance. The point, P, whereby the centerlines and the mutual tangent of the turning cycles meet, is the pitch-line. Point E is the start point of the meshing, A is the end of the meshing, and B and D are the beginning and ending points for the single meshing, respectively. Although the meshing beginning and ending points have high sliding velocity, in the ED and BA region the theoretical tooth load is half of the total tooth load. Sliding in the DB region is less than the sliding in the tooth tip and tooth root"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003082_0022-2569(70)90069-8-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003082_0022-2569(70)90069-8-Figure2-1.png",
        "caption": "Figure 2.",
        "texts": [
            " Note, however, that for a specific input-output (0, 8), six parameters only may be altered, as we need only express three link-lengths in terms of the fourth. Thus, the number of link-length parameters is really three. We shall not define any dimensionless link parameters at this stage, since the base link that we choose depends upon the nature of the exercise. It will also be necessary later to relate the position of angle O as a function of angle 8. For this, we need to express the location of axis A relative to D; the dimensions shown in Fig. 2 are related as follows: cos a ' = [sin (X + tx)/cos/3]2+ [sinZX + sin2(X + tx) tan\"/3] - sinZo~ 2[sin (h + a)/cos/3] [sin~-~ + sin2(X + a) tan2/3] ,/2 tan/3 tan/.t = sin X sinB' = sin tz cos (h +o~) COS ,u, tan h' = tan (X + a)\" (2.1) (2.2) (2.3) (2.4) J.M. Vol. $, No. 3--H 396 The angle 8~, which the intersection of the plane P and the plane defined by line 1 and C D makes with L is given by tan 8~ = tan/3 cot A. (2.5) Our first task is to express the variable angle ~5 as a function of the independent variable 0",
            " This means that the R - S - S-R linkage can only be assembled in one particular configuration and is a structure. (d) The ellipse intersects the circle in two points. Then the portion of the ellipse within the circle defines a range of possible input angles. (e) The ellipse intersects the circle in four points. Then there exist in general two distinct ranges for the input angle 0. A limiting case could exist where the ellipse intersects the circle in two points and touches the circumference at one point. Two ellipse diagrams for the cases AB as input (Fig. 1), and CD as input (Fig. 2), completely define the input and output angle ranges of motion. The parameter changes for CD as input have been described in Section 2. 399 Figure 4 shows an ellipse diagram for the input angle 0 of an R - S - S - R mechanism (Fig. 1). For any point on the ellipse, we know that X - - L = M cos 0 or x~ = M cos 0 (4.1) (4.1) therefore gives cos 0 at any instant. Also, f rom (3.17) and (3.13), Yl = N cos 0 - P sin 0. (4 .2 ) 400 From (4.1) and (4.2), we may eliminate cos 0 to give sin 0 in terms ofx l and Yr"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001643_09544062jmes1340-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001643_09544062jmes1340-Figure1-1.png",
        "caption": "Fig. 1 (a) Line of action of a skew conical gear drive. (b), (c) Principal directions of the conical gear tooth surface",
        "texts": [
            " Conical gears are a type of involute cylindrical gear with variable profile shifting along the face width [1\u20134, 6]. The tooth profiles in all of its transverse sections are involutes originating from the same base cylinder. The meshing of the conical gears can be also analysed by applying the basic geometrical characteristics of involute cylindrical gearing. Similar to conventional involute cylindrical gearing, the locus of the points of contact between two engaged teeth of a spatial conical gear pair is also a straight line, the \u2018line of action\u2019. Figure 1(a) shows the line of action T1T2 of a skew conical gear pair that is tangential to the base helixes \u03b2Cb1,2 and the base cylinders with radius rCb1,2. The \u2018dual-number\u2019 method [12] is applied to formulate the line of action. Two coordinate systems Sf and Sg are set up for analysis. As shown in Fig. 1(a), the xf -axis and xg-axis are arranged as the common perpendicular of the two skew gear axes; the zf -axis and zg-axis represent the axes of gear 1 and gear 2, respectively. Based on the assembly conditions, and with a shaft angle and an offset d, the axes A\u03022 of gear 2 can be expressed in the form of a dual vector by A\u03022 = \u23a1 \u23a3 0 sin \u2212 \u03b5d cos cos + \u03b5d sin \u23a4 \u23a6 (1) According to the condition of tangency to the base cylinder and the base helices of gear 1 (Fig. 1(a)), the line of action n\u0302 can be also represented as n\u0302 = n1 + \u03b5n1 \u00d7 rG (2) with the following direction vector of the line of action n1 = \u23a1 \u23a3\u00b1cos \u03b2Cb1 sin \u03c81 \u2213cos \u03b2Cb1 cos \u03c81 \u00b1sin \u03b2Cb1 \u23a4 \u23a6 (3) and the position vector for the point of tangency T1 rG = \u23a1 \u23a3rCb1 cos \u03c81 rCb1 sin \u03c81 bL \u23a4 \u23a6 (4) The upper sign in the equations above denotes the left-hand flank in engagement and the lower sign, the right-hand flank. The line of action n\u0302 is also tangential to the base cylinder and the base helix of gear 2",
            "comDownloaded from directions iI1, iI2 of the tooth surface of gear 1 and gear 2, that is iI1 \u00b7 iI2 |iI1||iI2| = cos \u03b7 (12) The first principal direction iI1 at any point y for gear 1, expressed in Sf , is determined from the perpendicular condition to the line of action n1, namely iI1 = \u23a1 \u23a3\u2212sin \u03b2Cb1 sin \u03c81 sin \u03b2Cb1 cos \u03c81 cos \u03b2Cb1 \u23a4 \u23a6 (13) Similarly, iI2 of conical gear 2, represented in Sf , is obtained by coordinate transformation with the assembly condition iI2 = \u23a1 \u23a3 \u2212sin \u03b2Cb2 sin \u03c82 sin \u03b2Cb2 cos \u03c82 cos + cos \u03b2Cb2 sin \u2212sin \u03b2Cb2 cos \u03c82 sin + cos \u03b2Cb2 cos \u23a4 \u23a6 (14) where \u03c82 is the positioning angle for locating the point of tangency T2 (Fig. 1(a)). By utilizing the relation for equality of the normals of the tooth surfaces, that is n1 = \u2212n2 (15) the angle \u03c82 can be solved as follows cos \u03c82 = \u2212cos \u03b2Cb1 cos \u03c81 cos \u03b2Cb2 cos + tan \u03b2Cb2 tan (16) JMES1340 \u00a9 IMechE 2009 Proc. IMechE Vol. 223 Part C: J. Mechanical Engineering Science at UNIV OF MICHIGAN on June 20, 2015pic.sagepub.comDownloaded from Substituting equations (7), (13), (14), and (16) into equation (12) yields cos \u03b7 = sin \u03b2Cb1 sin \u03b2Cb2 + cos cos \u03b2Cb1 cos \u03b2Cb2 (17) Equation (17) indicates that for a specified gear pair the angle \u03b7 is always constant, independent of the position of contact of the teeth",
            " The general assembly relations can be derived along with the working gearing parameters (subscript w) from the working common rack as follows [5, 6, 11] cos = cos \u03b8w1 cos \u03b8w2 cos(\u03b2w1 + \u03b2w2) \u2212 sin \u03b8w1 sin \u03b8w2 (26) d = (rCw1 cos \u03b8w2 + rCw2 cos \u03b8w1) sin(\u03b2w1 + \u03b2w2) sin (27) If the working common rack differs from the rackcutter used for gear generation, this is a case of profile-shifted transmission [16, 17]; otherwise it is a case of standard transmission. The assembly relations for a skew conical\u2013helical gear drive in a standard transmission, where \u03b82 = 0, are expressed as cos = cos \u03b81 cos(\u03b21 + \u03b22) (28) d = (rC1 + rC2 cos \u03b81) sin(\u03b21 + \u03b22) sin (29) In general, the line of action between the base cylinders of a skew cylindrical gear pair is unique, as shown in Fig. 1(a). The tooth surfaces therefore engage via point contact. However, if sin \u03c8 = 0, as in equation (8), the parameter bL cannot be determined definitely. The number of lines of action in such a case is infinite. The line of action is extended to a plane and the gear teeth can be engaged in line contact. The condition of line contact can be obtained from equations (7) and (8) by = \u03b2Cb2 + \u03b2Cb1 (30) d = rCb1 + rCb2 (31) The derived equations can be also explored via the geometrical relation. For the case of line contact, the Proc. IMechE Vol. 223 Part C: J. Mechanical Engineering Science JMES1340 \u00a9 IMechE 2009 at UNIV OF MICHIGAN on June 20, 2015pic.sagepub.comDownloaded from relation between the base cylinder of gear 1 and gear 2 (see Fig. 1(a)) must satisfy the conditions below (as shown in Fig. 3): (a) the shaft angle is the sum of the base helix angle of gear 1 and of gear 2; (b) the offset d is equal to the sum of the base radius of gear 1 and of gear 2. The line contact condition is also valid when the angle \u03b7 is equal to zero. By substituting \u03b7 = 0 into equation (17), the same result as for equation (30) can be obtained. Edge contact occurs when the line of action shifts from its theoretical position outside the face-width.This can occur due to assembly errors or manufacturing errors. The shift of the line of action bL is a suitable criterion for sensitivity analysis of the edge contact. As can be seen in Fig. 1, a shift in the line of action bL is defined as displacement along the axis of gear 1 from the theoretical position bL0 to the actual position b\u2217 L due to the error, that is bL = b\u2217 L \u2212 bL0 (32) When bL is negative in sign, it means that the line of action has shifted towards the toe of the conical gear, and vice versa. The maximum contact stress \u03c3H for infinite face-width is taken into account in equation (18). The specific contact stress \u03c3H/(p)1/3 serves as the criterion for evaluating the surface durability of a gearing design"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002738_we.1545-Figure4-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002738_we.1545-Figure4-1.png",
        "caption": "Figure 4. Typical wind turbine gearbox.",
        "texts": [
            " Moreover, different planet orientations and tooth counts cause these modulation sideband patterns to vary as sidebands are canceled or enhanced depending on the planetary gear design specifications. Wind Energ. (2012) \u00a9 2012 John Wiley & Sons, Ltd. DOI: 10.1002/we In many wind turbines, a specially arranged planetary gear set is used in the first stage of the gear train. In our study, we will be using a planetary gear set where the ring gear is fixed, the carrier is the input driven by the main rotor and the sun gear is the output as shown in Figure 4. For simplicity, in the following derivations, we assume a three-planet stage where the planets are equally spaced. It should be noted that the same derivations for systems with unequally spaced planets will be much more involved. Here, we assume that the gear meshing frequency is a pure tone with frequency of fm, while the modulation frequency is also a pure tone with frequency of fc. No harmonics from gear mesh or from carrier rotation are considered at this stage. Moreover, we assume the ring gear tooth number, Nr, is not divisible by Np = 3, i"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001221_13506501jet677-Figure4-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001221_13506501jet677-Figure4-1.png",
        "caption": "Fig. 4 Whole crankshaft finite-element model: (a) beam element and (b) solid element",
        "texts": [
            " Compared with the free beam method and the continuous beam method, the whole crankshaft beam-element finite-element method uses the whole crankshaft model; hence, it is closer to the actual situation and can calculate the loads of all the main bearings of a multi-cylinder engine directly and simultaneously. Therefore, the whole crankshaft beamelement finite-element method is an accurate and time-saving method for calculating the load of a crankshaft bearing. The whole crankshaft beam-element model is shown in Fig. 4(a), which consists of 164 elements and 165 nodes. The whole crankshaft solid-element model is shown in Fig. 4(b), which is divided by a hexahedron element and is composed of 4540 elements and 23 172 nodes. The loads on all connecting rod journals (including magnitude and acting direction) are obtained by analysing the acting forces on the crank-rod mechanism of an engine, based on the condition that the working phase differences between different cylinders are considered. The centrifugal forces caused by the crankshaft mass are applied automatically by finite-element analysing software in calculation when the material density of the calculating model and crankshaft rotational speed are given"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000556_icma.2007.4303844-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000556_icma.2007.4303844-Figure2-1.png",
        "caption": "Fig. 2 Simplified musculoskeletal system for jumping motion",
        "texts": [
            " The wire is equivalent to the bi-articular muscle which transmits the joint torque between two joints, and the wire constitutes a pantograph mechanism. An ideal jumping motion is defined as ground reaction force turning to the center of gravity. The mechanical function of bi-articular muscle is clarified, and the jumping robot which has the function of biarticular muscle is realized. A simplified musculoskeletal system which picked out the element required for jumping motion from Fig. 1 is shown in Fig. 2. This musculoskeletal system consists of three links: the femur, curs and foot, and of two joints: the knee joint (K) and ankle joint (A). Torque Tk and Ta are the joint torque on the joint K and A, respectively. Muscles for the jumping motioin are the knee extensor (Ke) by the rectus femoris and vastus mediaris, lateraris and intermediaris, the ankle planar extensor (Ae) by the soleus, and bi-articular muscle (Bi) by the gastrocnemius. Force Fk, Fa and Fb are the ground reaction force at the toe (T) generated by the muscle Ke, Ae and Bi, respectively"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003702_cjme.2013.04.801-Figure6-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003702_cjme.2013.04.801-Figure6-1.png",
        "caption": "Fig. 6. Equivalent scheme of heavy-duty six-legged robot under crab-type tripod gait",
        "texts": [
            " mc, mt, and ms are respectively defined as the mass of one coxa, the mass of one thigh, and the mass of one shin. Meanwhile, the mass mc contains the mass of the driver device and actuating device for one abductor joint. The mass mt includes the mass of the driver devices and actuating devices for one hip joint and one knee joint. The equivalent leg is obtained, when the equivalences are respectively executed through legs 4 and 6 to the straight line l1. Because of the symmetry of legs 4 and 6, the moments of couples caused by the equivalences are mutually offset. The equivalent scheme is shown in Fig. 6. In Fig. 6, \u0413 is defined as the included angle between the equivalent coxa and the equivalent shin. \u0413 is viewed as the included angle between the equivalent coxa and the equivalent thigh. Based on Fig. 6, the mathematical expression of the torque is written for foothold o. The normal contact force Fz2 can be solved. Then G G 2 2sin cos 0i x i z z zG l G l F l\u03b1 \u03b1       , (1) where G t 2 s 2sin sinxl l l\u03b2 \u03b2  , CHINESE JOURNAL OF MECHANICAL ENGINEERING \u00b7805\u00b7 G bpzl r l   , 2 t 2 s 2 c bpcos cos 2zl l l l r l\u03b2 \u03b2      , c t scos cosl l l l\u0393 \u0393       , bp c t 2 s 2( cos cos ) cos . 2 r l l l l\u03b2 \u03b2 \u03b8       And G is the weight for the robot and rated load, rbp is the effective radius of the bearing platform, cl  is the effective length of the equivalent coxa, tl  is the effective length of the equivalent thigh, sl  is the effective length of the equivalent shin, l is the projection length of the equivalent leg in the direction of x, lGx is the distance from the vector of the G component along the direction of x to foothold o, lGz is the distance from the vector of the G component in the vertical direction from the plane of the slope to foothold o, and lz2 is the distance from the vector of Fz2 to foothold o"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003211_amr.505.154-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003211_amr.505.154-Figure2-1.png",
        "caption": "Fig. 2 Boundary conditions for friction clutch",
        "texts": [],
        "surrounding_texts": [
            "In this paper, an educational software called Heatflux for computing the amount of heat generated on the flywheel, clutch (flywheel side and pressure plate side) and pressure plate is used. In this software, any function of rotational sliding speed and torque of friction clutch can be inserted to obtain a function for heat generated with radius and time. In the second part of this work the finite element method has been applied to study the influence of non-dimensional radius (R) on the maximum temperature, average temperature and temperature distribution of a friction material. The conclusions obtain from the present analysis can be summarized as follows: 1. The ratio of inner to outer radius of friction surface (R), which is considered the single most important factor affecting the design parameters and thermal behaviour of friction clutch. 2. In this work the analysis was based on the uniform pressure and uniform wear theories. To obtain results with high accuracy, one must know the proper functions of pressure with radius and rotational sliding speed with time. 3. The friction material of clutch should have perfect thermal properties and higher wear resistance for thermal stabilities. 4. The amount of heat generated on the friction clutch side (flywheel side or pressure plate side) is less than 5% from the total heat generated between surfaces in one side. 5. The maximum effect of thermal load (the highest temperatures) at the friction interface for all cases occur approximately at half sliding time (0.2 sec), and the highest average temperatures occur in 0.3 sec (0.75% from the slipping time)."
        ]
    },
    {
        "image_filename": "designv11_12_0000894_iros.2008.4651029-Figure9-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000894_iros.2008.4651029-Figure9-1.png",
        "caption": "Fig. 9. Convention adopted for the inertial moments estimation.",
        "texts": [
            " DYNAMIC ANALYSIS The dynamic problem has been solved by adopting the traditional approach: once defined (6), (7) and (8), the solving expression results (9) [24,30,35]. ( ) TQ J J=M M (6) ( ) ( ), TQ Q Q J J Q= =V V M (7) ( ) ( ), T se qQ F J F F= \u2212 +G (8) ( ) ( ) ( ), , 0Q Q Q Q Q F+ + =M V G (9) To determine the mass matrix M, and the column vector Fse of the external forces applied to the platform, other elements have to be taken into consideration. After having approximated the links to a two concentrated masses model and estimated the inertial moments for the platform along the x and y axis according to the figure 9, the mass matrix elements can be easily defined. Concerning with the Fse vector, the equation (10) must be verified s se siF F F= + (10) with Fsi the column vector of the inertial forces applied to the mobile platform. In order to completely determine the external forces exerted on the platform the last considered contribute is introduced by the material elasticity. If the elastic phenomena are idealized as concentrated into the flexure hinges at the joints, two kinds of elasticity models has been developed as figure 10 shows"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003415_j.mechmachtheory.2011.03.006-Figure4-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003415_j.mechmachtheory.2011.03.006-Figure4-1.png",
        "caption": "Fig. 4. The construction of the subtractive Goldberg 5R linkage. (a) and (b) Two Bennett linkages with one common link a/\u03b1; (c) the subtractive Goldberg 5R linkage; (d) the detailed relationship between the kinematic variables at joints 4 and 5.",
        "texts": [
            " A more generalised 5R linkage was also proposed by Goldberg briefly [5] and later derived in detail by Wohlhart [8]. In the general case, two links which form the rigidified link are not collinearly posed. A variable \u201ckink angle\u201d was introduced, see Fig. 3. Therefore, Goldberg 5R linkage is a special case of the generalised 5R linkage when the kink angle is zero. When the kink angle in the generalised Goldberg 5R linkage equals to \u03c0, the two links adjacent to the common link are overlapped and the resultant linkage is in fact formed by subtracting Bennett linkage B from Bennett linkage A as shown in Fig. 4, which is the first variation of the Goldberg linkages given in [5]. Therefore, we call such linkage as the subtractive Goldberg 5R linkage. Consider two Bennett linkages shown in Figs. 4(a) and (b). In order to construct the subtractive Goldberg 5R linkage, these Bennett linkages must have a common link with the identical geometric parameters, a/\u03b1. So the Bennett linkages A and B should have the following geometric parameters. and aA12 = aA34 = aB12 = aB34 = a;aA23 = aA41 = b;aB23 = aB41 = c; \u03b1A 12 = \u03b1A 34 = \u03b1B 12 = \u03b1B 34 = \u03b1;\u03b1A 23 = \u03b1A 41 = \u03b2;\u03b1B 23 = \u03b1B 41 = \u03b3; sin\u03b1 a = sin\u03b2 b = sin\u03b3 c ; RA i = RB i = 0 i = 1; 2; 3 and 4\u00f0 \u00de: \u00f03\u00de Then the closure equations of Bennett linkages A and B are \u03b8A1 + \u03b8A3 = 2\u03c0;\u03b8A2 + \u03b8A4 = 2\u03c0; tan \u03b8A1 2 tan \u03b8A2 2 = sin \u03b2 + \u03b1 2 sin \u03b2\u2212\u03b1 2 ; \u00f04\u00de \u03b8B1 + \u03b8B3 = 2\u03c0;\u03b8B2 + \u03b8B4 = 2\u03c0; tan \u03b8B1 2 tan \u03b8B2 2 = sin \u03b3 + \u03b1 2 sin \u03b3\u2212\u03b1 2 ; \u00f05\u00de tively. respec The subtractive Goldberg 5R linkage can be formed by removing the common links and joint as shown in Fig. 4(c). The geometric parameters of the subtractive Goldberg 5R linkage are a12 = a34 = a; a23 = b\u2212c; a45 = c;a51 = b; \u03b112 = \u03b134 = \u03b1;\u03b123 = \u03b2\u2212\u03b3;\u03b145 = \u03b3;\u03b151 = \u03b2; sin\u03b1 a = sin\u03b2 b = sin\u03b3 c ; Ri = 0 i = 1; 2;\u2026; 5\u00f0 \u00de: \u00f06\u00de In Fig. 4, the relationship among the revolute variables of the subtractive Goldberg 5R linkage and the Bennett linkages A and B can be set as follows. \u03b81 = \u03b8A1 ;\u03b82 = \u03b8A2 ;\u03b83 = \u03c0\u2212\u03b8B2; \u03b84 = 2\u03c0\u2212\u03b8B1; \u03b85 = \u03c0\u2212 \u03b8B4\u2212\u03b8A4 : \u00f07\u00de And in order to construct the subtractive Goldberg 5R linkage, the compatibility between Bennett linkages A and B, \u03b8A3 = \u03b8B3; \u00f08\u00de must be preserved. Substituting Eqs. (7) and (8) into Eqs. (4) and (5), the closure equations of the subtractive Goldberg 5R linkage can be derived as follows. \u03b81 + \u03b84 = 2\u03c0;\u03b82 + \u03b83 + \u03b85 = 2\u03c0; tan \u03b81 2 tan \u03b82 2 = sin \u03b2 + \u03b1 2 sin \u03b2\u2212\u03b1 2 ; and tan \u03b81 2 tan \u03b83 2 = sin \u03b3 + \u03b1 2 sin \u03b3\u2212\u03b1 2 : \u00f09\u00de Fig. 4 shows that in the subtractive Goldberg 5R linkage, the link-pairs 51\u201312 and 34\u201345 are from Bennett linkages A and B, respectively. Therefore, we can build 6R linkages from two subtractive Goldberg 5R linkages that share a common Bennett linkage A or B, which is similar as the construction of Goldberg 6R linkages [5] and their variations [12]. Alternatively, due to the geometric condition that sin \u03b1/a= sin \u03b3/c in Eq. (3), we can use link-pair 45\u201351 to form another common Bennett linkage, which is different from Bennett linkage A or B"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000533_tia.2007.895763-Figure4-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000533_tia.2007.895763-Figure4-1.png",
        "caption": "Fig. 4. Induction motor scheme with thermocouples location.",
        "texts": [
            "2 A, mechanically coupled to a separately excited dc generator. The induction machine is totally enclosed fan cooled (TEFC), with a cast-aluminum squirrel cage and thermal class F (hot spot 155 \u25e6C) according to IEC 34-1. Other three-phase induction motor parameters are summarized in Table I (see the Appendix). ias = IM cos(\u03c9st + \u03c6) (11) ibs = IM cos ( \u03c9st + \u03c6 \u2212 2\u03c0 3 ) (12) ics = IM cos ( \u03c9st + \u03c6 + 2\u03c0 3 ) . (13) The stator and rotor temperatures are measured by several thermocouples positioned inside the machine according to the schematic presented in Fig. 4. Thermocouples T1, T2, T3, and T4 are for stator temperature measurements, one placed in a stator tooth on the load side and the other three placed in the slots of phases U, V, and W, respectively. Thermocouples T5, T6, T7, T8, and T9 are for rotor temperature measurements. Thermocouples T5, T7, and T9 are placed at the rotor surface, whereas T6 and T8 are placed 5 cm deep inside the rotor. Thermocouple T10 is used to measure the motor surrounding temperature. The temperature evaluation is obtained by a data acquisition system through an infrared transmission device (for the rotor) [14]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000597_978-3-540-73129-0-Figure6.13-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000597_978-3-540-73129-0-Figure6.13-1.png",
        "caption": "Fig. 6.13. Diagram P (E) and S(E) for a typical piezoceramic for T = 0. The actuators operation cycle starts at point E = 0, Sr (derived from [5])",
        "texts": [
            ", the regions consisting of crystallites of uniform dipole orientation) will show a statistically distributed orientation, i. e., the macroscopic body is isotropic and has no piezoelectric properties. Only when a strong electrical dc field is applied, the dipole regions become almost completely arranged (polarization). After switching off the polarization field, this arrangement remains to a large extent, that is, the ceramic body features a remanent polarization Pr, combined with a permanent elongation Sr of the body (see Fig. 6.13). PZT ceramics are chemically inactive and can cope with high mechanical loading, but are also brittle and therefore difficult to process. The permissible compressive stress is considerably higher than the tensile stress. This is why the elements need to be pre-stressed when extensive tensile stress is applied. PZT ceramics belong to the group of ferroelectric materials which feature a hysteretic behaviour shown in the diagram P (E) in Fig. 6.13. Due to the relation P = D \u2212 \u03b50E (P : electric polarization) and D = \u03b5E, the two diagrams P (E) and D(E) differ m erely by the term \u03b50E. For actuator operation the diagram S(E) of the polarized ceramic, the so-called butterfly trajectory shown in Fig. 6.13 (right hand side) is crucial. The maximum achievable strain is limited by the saturation and the repolarization. Precautions must be taken in order to avoid depolarization during actuator operation due to electrical, thermal and mechanical overload. Piezoceramics, for instance, gradually loose their piezoelectric properties even at operating temperatures far below the Curie temperature (depending on the material 120 . . . 500 \u25e6C, for multilayer ceramics (see below) 80 . . . 220 \u25e6C). Under certain applications when the inverse operating voltage is applied, it may not exceed 20% of the rated voltage, or depolarization may occur"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000028_s0022112081000645-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000028_s0022112081000645-Figure1-1.png",
        "caption": "FIGURE 1 . Co-ordinate system used in $ 3 2, 4.",
        "texts": [
            " To evaluate $ I when the vortex is given by its perturbed position (2.10), we write $1 = $ O A 1 + 4 l l + O(a2) ) , (2.18) where $oI is the non-dimensional velocity potential due to the image of the undisturbed vortex in the sphere and the O ( C ~ $ ~ ~ ) terms allow for the perturbation from the straight vortex. The potential f i O I can be determined by as spherical harmonic analysis as follows. 134 M . R. Dhanak In a co-ordinate fmme &@?, fixed with respect to the centre of the sphere, the undisturbed position of the vortex is given by Y = y (see figure 1) where 5? = X([, - 00) - X, ( t ) = (f, 0, [), - 00 < [ < 00. (2.19) In the absence of the sphere, the velocity at a field point Y = (Z;,g,Z) is due to the straight vortex and the velocity potential is given by $,*(Y) = - tan-1 - r y\u201d 27l 5 - f \u2019 so that (4, = $$/Ua) - $,(Y) = 2B tan-1& x - f = 2B tan-l ( sin $ ) r sin0 cos $- f \u2019 (2.20) in terms of spherical polars ( r , 8, $), r2 = E2 + fj2 + .Z2. We seek the disturbance potential $oz when a rigid sphere is introduced. is to satisfy the boundary condition (2",
            " In 3 4 an expression for V, is obtained and the results of the calculations are described in 8 5. The image system of a vortex element in a sphere has been given by Lighthill (1956). This is briefly described here and an expression for the velocity field due to the image system of an infinitely long vortex is obtained. An approaching sphere interacting with a vortex filament 139 Suppose that, with the centre of a sphere of radius a a t the origin, a vortex element of length ds* and circulation r is situated a t YT (see figure 1). The strength of the element J * is defined as Then, writing IY:l = r:, the image system of the vortex element is given by (i) a vortex element of strength a t the inverse point and (ii) a line vortex of circulation - (J* . Y:)/arT stretching from the inverse point to the centre of the sphere. The image system satisfies the boundary condition at the surface of the sphere and the requirement that the vorticity field inside the sphere be solenoidal. The latter condition is necessary if the corresponding Biot-Savart velocity field is to be irrotational"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003924_1.4030612-Figure4-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003924_1.4030612-Figure4-1.png",
        "caption": "Fig. 4 Synchronous motor driven system [19]",
        "texts": [
            " Also, the fixed-point interaction method at each time step is computationally inefficient, especially for the nonlinear problem. The analysis code, including a graphical user interface, was formulated with a MATLAB program. An example of a three single-stage compressors [19] driven by synchronous rotor is applied at first to validate the nonlinear torque calculation and the torsional transient solver. Further, a new full system nonlinear transient system analysis is presented in this paper with some unexpected results. The schematic of the motor\u2013compressor configuration is from Ref. [19] and shown in Fig. 4. A 50 Hz, 4200 kW, four-pole, 1500 rpm synchronous motor with a rated torque of 26,716 N m is the prime driver. The synchronous speed is 1500 rpm and the motor is started with reduced voltage. The driving torque and the load torque versus the motor speed are given in Ref. [19] and are shown in Fig. 5. Since the loads are quite evenly distributed among the three stages, the three compressor load curves are very close to each other. The bull gear connects to three compressors at nodes 8, 11, and 14 with speed ratios of 8"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001368_j.tcs.2009.01.033-Figure12-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001368_j.tcs.2009.01.033-Figure12-1.png",
        "caption": "Fig. 12. The homomorphism h is induced by the map f .",
        "texts": [
            " There is a uniquemap \u03b4 : W nM \u2192 W nq satisfying \u03b4\u03c1 = \u03b9. See the right part of Fig. 10. Let f = \u03b4\u00b5 : W np \u2192 W nq . See Fig. 11. We proceed to show that f \u2217 n = h. For any 1 \u2264 i \u2264 p, f \u2217 n (\u03b1i) = (\u03b4\u00b5)\u2217n(\u03b1i) = (\u03b4\u00b5\u03c1p\u03b8i) \u2217 n(\u03c4 ) = (\u03b9\u03c9\u03b8i) \u2217 n(\u03c4 ) = (\u03b9i\u03c9i) \u2217 n(\u03c4 ) = (\u03b9i\u03b3i\u03c8i) \u2217 n(\u03c4 ) = (\u03b9i) \u2217 n(\u03b3i\u03c8i) \u2217 n(\u03c4 ) = (\u03b9i) \u2217 n( \u2211Mi j=1 \u03b1j) = \u2211Mi j=1(\u03b9i) \u2217 n(\u03b1j) =\u2211Mi j=1 di,j\u03b1ki(j), where di,j = { 1 if \u03bbi,ki(j) > 0 \u22121 if \u03bbi,ki(j) < 0. (5.7) As a result, f \u2217n (\u03b1i) = \u2211q j=1 \u03bbi,j\u03b1j, which means that f \u2217 n = h. Lemma 4 is illustrated in Fig. 12. Lemma 5. Given nice rendezvous tasks T = (K , \u03c3 (\u03a3n), \u03d5) and T \u2032 = (K \u2032, \u03c3 \u2032(\u03a3n), \u03d5\u2032), any homomorphism from sig(T ) to sig(T \u2032) can be induced by a map from (|K |, |\u03d5|) to (|K \u2032|, |\u03d5\u2032|). Proof. Let sig(T ) = (G, e) and sig(T \u2032) = (G\u2032, e\u2032), i.e. e = (|\u03d5|\u03b7)\u2217n(\u03c4 ), e \u2032 = (|\u03d5\u2032|\u03b7)\u2217n(\u03c4 ). Consider a homomorphism h : (G, e)\u2192 (G\u2032, e\u2032). We proceed to construct a map from (K , \u03d5) to (K \u2032, \u03d5\u2032) which induces h. According to Lemma 3, there are integers p and q, and maps \u03bd : |K | \u2192 W np and \u03c9 : W n q \u2192 |K \u2032 |, which induce isomorphisms \u03bd\u2217n : Hn(|K |)\u2192 Hn(W np ) and \u03c9\u2217n : Hn(W n q )\u2192 Hn(|K \u2032|)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000857_j.mechatronics.2008.01.003-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000857_j.mechatronics.2008.01.003-Figure2-1.png",
        "caption": "Fig. 2. The pneumatic components of the supply kit.",
        "texts": [
            " The box has sensing capabilities via two micro-switches, an infrared distance sensor, an encoder, and a flex sensor, all of which are provided as part of the supply kit. To program the controller boxes, the PBASIC programming language is used. Where possible, the kit has been made \u2018\u2018plug-and-play\u201d with many of the difficult operations hard-wired into the board and code examples made available. This simplicity allows the students to focus on the integration of the electronic components into a complete mechatronic device. The pneumatic supplies issued to the students are shown in Fig. 2 and include a one-way pneumatic actuator, a solenoid valve, and a pressure vessel. The one-way actuator has a stroke of approximately 2 in. and can be extended approximately 15 times using the air supplied from the pressure vessel. The inclusion of pneumatics in the supply kit is intended to fulfill two main objectives. First, because students typically have much less experience with fluid systems than mechanical, the kit helps to provide students with important fluid systems experience. Second, the power dense nature of pneumatic systems greatly improves the variety and quality of student designs"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001495_ssp.154.133-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001495_ssp.154.133-Figure1-1.png",
        "caption": "Figure 1. Initial crystallographic orientation related to the rolling direction.",
        "texts": [
            " The specimens in rectangular shape were cut along the [110] and [001] direction of the 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: 152.14.136.77, NCSU North Carolina State University, Raleigh, USA-27/04/15,03:24:50) columnar grains from the flat or round shape ingots of alloys 1 and 2. The relation between initial crystallographic orientation of the samples and the rolling direction is shown in Fig. 1. The alloy 3 was in the parent phase state and the samples were cut from a large flat ingot with random oriented grains. The rolling was carried out on the samples imbedded in a steel channel bar and \u201esealed\u201d in a flat pipe. The samples were heated up to 1000oC and rolled in one direction in several steps. Before each step the sample was reheated. The final reduction of the samples thickness was 28%, 36%, 57% and 69%. The texture was studied by EBSD and X-ray diffraction using the {001}, {110} and {111} pole figures"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003603_0954406212466479-Figure14-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003603_0954406212466479-Figure14-1.png",
        "caption": "Figure 14. Approximate geometry for determination of interference.",
        "texts": [
            " However, its slope and intercept are not constant, as illustrated by equations (59) and (60). Since the tooth number difference is very small compared to the tooth number of the internal gear, an approximation can be obtained cot\u00f0\u20191 \u20192 \u00fe \u00de cot \u00bc k \u00f061\u00de Then, we replace the work profile by an approximate line, as represented by equation (62) y2 \u00bc kx2 \u00fe b0 \u00f062\u00de As shown in Figure 13, the approximate line intersects the addendum circle at point K0, which is used to replace point K in the solution process. Then, we get an approximate profile, as shown in Figure 14. According to the definition of the gear with straightline profile, the approximate line passes through point C and a relation can be obtained xc \u00bc r2 sin c yc \u00bc r2 cos c \u00f063\u00de where (xc, yc) is the coordinate of point C in S2. Here c \u00bc s 2r2 \u00f064\u00de Equations (9) and (62) to (64) yield b0 \u00bc r2 sin sin s 2r2 \u00fe \u00f065\u00de The coordinate of point K0 can be solved by equation (66) yk0 \u00bc kxk0 \u00fe b0 x2 k0 \u00fe y2 k0 \u00bc r2 a2 \u00f066\u00de where (xk0 , yk0 ) is the coordinate of point K0 in S2. Then, the solution of equation (66) is xk0 \u00bc kb0 ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi \u00f01\u00fe k2\u00der2 a2 b02 q 1\u00fe k2 yk0 \u00bc k2b0 k ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi \u00f01\u00fe k2\u00der2 a2 b02 q 1\u00fe k2 \u00fe b0 8>>>>>< >>>>>: \u00f067\u00de at University of Bristol Library on January 6, 2015pic"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003761_9781118818428-Figure20.3-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003761_9781118818428-Figure20.3-1.png",
        "caption": "Figure 20.3 Cyclic variation of local effective surface temperature difference 1\u20442(Twd + Twc) for the following parameter values: temperature ratio NT = 2.77, length/diameter ratio of displacer Ld\u2215D = 2.5, bore/stroke ratio D\u2215Sp = 1.5",
        "texts": [
            " The values are arbitrary, but allows development to proceed pending further insight. With NT to denote temperature parameter TE\u2215TC: TdE\u2215TC = NT + (1 \u2212 NT)\u2215(Ld\u2215Sd + 1) (20.10) TdC\u2215TC = NT + (1 \u2212 NT)\u2215{(Ld\u2215Sd)\u22121 + 1} (20.11) The algebra appears to have introduced an extra parameter Ld\u2215Sd. This, however, is merely the product of other geometric parameters basic to the parallel bore, coaxial configuration, displacer length/diameter ratio Ld\u2215D, bore/stroke ratioD\u2215Sp and Finkelstein\u2019s (1960a) kinematic volume ratio \u03bb: Ld\u2215Sd = \u03bb(Ld\u2215D)(D\u2215Sp) (20.12) Figure 20.3 is a not entirely convincing attempt to portray the cyclic variation of effective gap temperature distribution 1\u20442(Twd + Twc) for the numerical values of the parameters declared The \u2018hot air\u2019 engine 225 in the caption. Relative motion is displayed to scale, but is achieved by holding displacer stationary and depicting the relative motion of selected points equi-spaced axially on the cylinder (the family of lines with strong curvature). Superimposed lines having slight curvature indicate the cyclic fluctuation, referred to the \u2018stationary\u2019 displacer, of local effective surface temperature 1\u20442(Twd + Twc)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000597_978-3-540-73129-0-Figure6.84-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000597_978-3-540-73129-0-Figure6.84-1.png",
        "caption": "Fig. 6.84. Detail of MR fluid shock absorbers on Corvette sports car",
        "texts": [
            " In 1998, a small, real-time controlled MR fluid damper system (the RD-1005-3 described above) was introduced commercially into the heavy-duty truck and off-highway vehicle market for suspended seat applications [152]. That same year, a controllable MR fluid based primary suspension shock absorber for NASCAR race-vehicles was introduced by Carrera [166]. Today, the greatest driving force behind MR fluid technology is automotive, particularly real-time controlled primary suspensions systems. In January 2002, the Cadillac Seville automobile, shown in Fig. 6.84, was introduced by General Motors with a MagneRide\u2122 suspension system having real-time controllable MR fluid shock absorbers and struts as standard equipment [167,168]. The Magneride\u2122 shock absorbers are made by Delphi Corporation with the MR fluid being made by Lord Corporation. Similar, controllable MR fluid-based suspension systems have since become available on numerous other vehicle models including: Corvette sports car [169], Cadillac SRX roadster, Cadillac XLR sport utility vehicle [170,171], Cadillac STS sedan, Cadillac DTS [172] and Buick Lucerne [173]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003737_b978-1-4557-7631-3.00004-1-Figure4.41-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003737_b978-1-4557-7631-3.00004-1-Figure4.41-1.png",
        "caption": "Figure 4.41: Carbon Nanotube-Based FET. (a) Schematic of a CNT-Based FET. (b) Scanning Tunneling Microscope Photograph of a SWNT FET. (c) Image of a MWNT Where Electrical Pulses Are Used to Remove Unwanted Carbon Layers.",
        "texts": [
            " The properties of SWNTs change significantly with the lattice directions. All armchair SWNTs are metals (i.e., conductive). Other SWNTs are semiconductors with the band gap varying from zero to about 2 eV, depending on their lattice directions and the diameters. Since a MWNT comprises an array of different SWNTs, most of MWNTs can be considered to be metals. SWNTs are likely candidates for key sensing element in molecular sensors. Bachtold et al. reported on a SWNT-based FET [58], which suggests basic designs for SWNT-based molecular sensors (Fig. 4.41). From [57]. Various methods have been demonstrated [59] to produce carbon nanotubes, and all of them require high temperatures in their processes. SWCNTs and MWCNTs can be made by laser ablation or pulsed-laser vaporization (PLV) of graphite [60], a method in which the material is made through etching of a solid graphite surface. Other methods include carbon arch (CA) discharge [48] and decomposition in an oxygen free environment [61]. The most commonly used has been chemical vapor deposition (CVD) which produces a more pure sample of the nanotubes [50]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000645_s00170-007-1183-9-Figure4-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000645_s00170-007-1183-9-Figure4-1.png",
        "caption": "Fig. 4 Discretized model of the hexapod with the relevant boundary conditions",
        "texts": [
            " These positions are given in Table 1. Six natural frequencies of the moving platform are obtained from Eq. (16) with the aid of a code written in MATLAB environment. The results are given in Table 2. In order to verify the aforementioned results, the natural frequencies of the upper platform are also obtained by FEM. For this purpose, the solid model developed in CATIA is exported to the Ansys software for modal analysis. The discretized model in the Ansys environment with the relevant boundary conditions is shown in Fig. 4. The results for the same ten positions of the platform are also given in Table 2. As an example, the rotational mode of vibration about X axis is shown in Fig. 5. It is noteworthy that the computational time for each solution has amounted to 3 h. As is evident from Table 2, the results obtained from the analytical model and FEM are in agreement. This can be better visualized from the comparative diagrams of Fig. 6. Some important inferences can be drawn from the foregoing results. The lowest natural frequencies in all modes of vibration occur when the moving platform takes the higher positions"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002226_978-3-642-00644-9_38-Figure6-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002226_978-3-642-00644-9_38-Figure6-1.png",
        "caption": "Fig. 6 SuperBot rolling track of size 6, 8 and 10 in simulation",
        "texts": [
            " Each module selects its joint angle based on its accelerometer values from its own unique orientation in a polygon, therefore, identifier is not required. The control also requires no message exchange avoiding hop delay issue. Rolling track of size 6-module, 8-module and 10-module are demonstrated in simulation. They are implemented in SuperBot simulation using Open Dynamic Engine. Control programs have been loaded into simulated modules without any modifications to the control program. As shown in Figure 6, the rolling tracks of different size are able to detect its configuration and turn while rolling. The experiment suggests the algorithm can support a higher number of modules if the joints are strong enough to support its load for rolling motion. Video of 6-module, 8-module, 10-module rolling track simulation can be viewed at: http://www.isi.edu/robots/superbot/movies/rtSimRolling.avi (or rtSimRolling.swf for faster download) As SuperBot is designed to be self-reconfigurable, we would like to test the future adaptability of the control to loop formation"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000224_1-4020-2933-0_13-Figure4-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000224_1-4020-2933-0_13-Figure4-1.png",
        "caption": "Figure 4. A revolute joint",
        "texts": [
            " This selection issue can be avoided if we define two vectors on body j perpendicular to j s , and then enforcing is to be perpendicular to these two vectors. The most common kinematic joint between two bodies is a spherical (ball) joint, as shown schematically in Figure 3. This joint requires point Pi on body i, and point Pj on body j to remain coincident. This condition, containing three algebraic equations, is expressed as: 3s P P i j \u2261 \u2212 =r r 0( , ) . (7) By combining the constraints in Eqs. (5)-(7), we can represent other types of kinematic joints. As an example, consider a revolute (pin) joint shown schematically in Figure 4. Point Pi on body i and point Pj on body j are defined along the joint axis, where they must remain together. Vector is is defined along the joint axis on body i and vectors ja and jb are defined on body j perpendicular to the joint axis. The following constraints can be written: 3 5 1 1 T 1 1 T 0 0 s P P i j r n i j n i j \u23a7 \u2261 \u2212 = \u23aa\u23aa\u2261 \u2261 =\u23a8 \u23aa \u2261 =\u23aa\u23a9 r r 0 s a s b ( , ) ( , ) ( , ) ( , ) . Similarly, other types of kinematic constraints, such as cylindrical, prismatic, universal can be constructed"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003359_s0025654413010020-Figure7-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003359_s0025654413010020-Figure7-1.png",
        "caption": "Fig. 7.",
        "texts": [
            "4) leaves the strip \u2212\u03c0 \u2264 \u03b2 \u2264 \u03c0. Figure 6 presents the optimal control synthesis picture constructed in the domain D for d = 0.8, e = 0.5, and u0 = 0.5. The switching curve K is, of course, described by Eqs. (3.13), (3.14). As m0 \u2192 \u221e, the optimal control synthesis picture given in Fig. 6 becomes the synthesis picture for the usual pendulum with a fixed suspension point [19\u201321]. MECHANICS OF SOLIDS Vol. 48 No. 1 2013 Now consider a system consisting of a carriage and an n-link pendulum hinged to it (Fig. 7). The carriage can move (roll) without drag along the horizontal line X. Its displacement along the X-axis is denoted by x. The carriage with n-link pendulum hanging downward can be a model of a crane plant with a load suspended from it on a flexible rope. In contrast to the problem considered above, the control input in the crane plant is applied to the crane trolley so as to bring the load suspended from it to the desired place. It is still assumed that the torque M is applied to the first link of the pendulum"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003713_978-1-4614-2419-2-Figure14.3-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003713_978-1-4614-2419-2-Figure14.3-1.png",
        "caption": "Fig. 14.3 (left) Photograph of torsional vibration test stand, (right) CAD drawing of the flywheel and supporting shaft",
        "texts": [
            " The frame supporting the flywheel and drive motor was designed such that its first mode of vibration is at least ten times greater than the highest driving speed (200 Hz). In order to facilitate future works where the test stand will be mounted on a shaker table to evaluate the cross-axis sensitivity of the sensors, the frame and bearings were designed to withstand up to 10 g\u2019s of vertical acceleration with minimal effect on the angular acceleration of the flywheel shaft. A photograph of the test stand can be seen in Fig. 14.3, as well as a schematic showing the flywheel and supporting shaft. The system was designed so that large diameter encoders could be attached on either end of the flywheel shaft, as well as any sensor (such as the angular accelerometers) that can be bolted to the end of the shaft. The structure supporting the flywheel has 1.75-in, 2.25-in, 6.0-in, 12.0-in, and 16.7-in sections so the torsional laser vibrometer can be applied at the various diameters. The following section describes the hardware in detail and the methodology that was used to minimize torsional vibration of the flywheel. A ring gear with 138 teeth is also installed on the flywheel so a gear tooth sensor (see far left in Fig. 14.3) can be used to measure the torsional motion of the flywheel. The flywheel weighs approximately 41.7 kg and has inertia of 1.07 kg-m. The flywheel is connected to a three-phase AC motor through a soft belt, and is dynamically balanced in order to reduce torsional motions of the shaft. The flywheel 14 Comparison of Noise Floors of Various Torsional Vibration Sensors 155 is mounted on a steel shaft, which is supported by two pillow block roller bearings. The frictional torque due to the bearings was expected to be about 0",
            "63 Hz and a belt was found whose stiffness was approximately 1,500 N/m. This should assure that the flywheel is isolated from fluctuations in the motor torque so that the noise floor of each sensor will be visible. After the system was manufactured, a modal test was performed to estimate the modes of the frame/flywheel assembly. The frame was tested in free-free conditions and the results correlated to an approximate finite element model. The first bending mode of the free-free frame was found to be 196 Hz. The bracket that holds the magnetic pickup sensor (see Fig. 14.3) was also tested and its first elastic mode was found to be 484 Hz). The finite element model was then used to estimate the first mode of the assembly when mounted on a rigid base, which was found to be at 259 Hz corresponding to torsional motion of the frame. During all of the noise floor measurement tests described in this work, the frame was bolted to a 36,000 lb seismic mass. This section lists all of the torsional vibration sensors that were considered in this work, briefly discusses their theory of operation, and provides some details regarding how they were attached to the flywheel system during the tests"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003940_iconac.2015.7313942-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003940_iconac.2015.7313942-Figure1-1.png",
        "caption": "Figure 1. Schemetic for a planetary gearbox",
        "texts": [
            " This MSB based approach has been shown to yield outstanding performance in characterizing the small modulating components of motor current signals for diagnosing different electrical and mechanical faults under different load conditions [13][14][15]. Therefore, it is also evaluated in this study to extract the residual sidebands of vibration signal for the purpose of gear and bearing fault diagnosis. III. EXPERIMENTAL SETUPS To verify the effectiveness of MSB-SE based diagnosis, vibration signals were acquired from an inhouse planetary gearbox test system. The maximum torque of planetary gearbox is 670 Nm, the maximum input speed is 2800 rpm and maximum output speed is 388 rpm. The schematic in Fig. 1 shows the position of the accelerometer that mounted on the outer surface of the ring gear and the position of experiment studied bearing. In the experiment, the planetary gearbox operates at 80% of its full speed under 5 load conditions (0%, 25%, 50%, 75% and 90% of the full load). The load setting allows fault diagnoses to be examined with variable load operations which are the cases for many applications such as wind turbine, helicopters etc. The vibration is measured by a general purpose accelerometer with a sensitivity of 31"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003702_cjme.2013.04.801-Figure3-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003702_cjme.2013.04.801-Figure3-1.png",
        "caption": "Fig. 3. Structure of one leg of heavy-duty six-legged robot",
        "texts": [
            " Based on the characteristics of the sine function, it is deduced that the numerical value of F1 gradually increases when the angle of the slope changes from 0\u00b0 to 45\u00b0. Therefore, the rearmost leg experiences the maximum normal contact force when the heavy-duty six-legged robot passes along a slope. torques The electrically driven heavy-duty six-legged robot is called the heavy-duty six-legged robot for short in this paper. The structure of one leg of the heavy-duty six-legged robot is shown in Fig. 3. Every leg of the heavy-duty six-legged robot contains three joints: an abductor joint, hip joint, and knee joint[12]. The axis of the abductor joint runs parallel to the z axis. The axes of the hip joint and knee joint follow the direction of y. The lengths, which refer to the coxa, thigh, and shin, are respectively set as lc, lt, and ls. \u03b2i (i1, 2, 3, 4, 5, 6) is defined as the included angle between the coxa of leg i and the shin of leg i. The range of the included angle is limited to 0\u00b0\u201390\u00b0 for \u03b2i"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001835_s12239-010-0044-y-Figure4-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001835_s12239-010-0044-y-Figure4-1.png",
        "caption": "Figure 4. (a) Photograph of spiders and (b) Schematic diagram of the spider locations.",
        "texts": [
            " Location of Spiders (Rotational Angles) Although the offset and PF are important factors in idle vibration, their values are very difficult to measure and therefore it is almost impossible to find the relationship between the idle vibration and the offset values and the PF directly. Fortunately, however, the variation of the offset is indirectly related to the rotation of the spiders; thus, various spider-positions lead one to useful information regarding the vibrational characteristics during idling. It is not clear where the spiders are located when the vehicle is idling. However, the static motion of spiders in a CV joint can be controlled. For example, the relative locations of spiders, as shown in Figure 4, can be manipulated along with the wheel position. In general, when a vehicle repeats stop-and-start motions, it is hard to figure out the spider locations in the CV joint. To generate meaningful rotations of the spiders, the wheels are rotated by 45o, as shown in Figure 4. Because the spiders are relatively rotated, the starting spider position is termed, 0o when the wheel has the shape of an inverted Y, as shown in Figure 4(a). Due to symmetry, in Figure 4(b), the total number of spider positions with additional rotations of 45o becomes 24. 3.1. Assembly Module Test of the Drive Shaft Before a CV joint is installed in a vehicle, the forces on the assembly module of the CV joint need to be measured because of the possibility of vibrational problems with the CV joint. Moreover, the information regarding vibration in the vehicle before and after installation is useful for revealing the relationship between vibrations and axial forces on the shaft. In this study, a mid-sized vehicle with four cylinders and an automatic transmission was used for the tests",
            " First of all, the vibrational characteristics at the differential gear housing, which is considered to be an exciter, were examined. No variation was observed with regard to the vibrational characteristics. Moreover, as shown in Figure 8, the vibrations on the y axis, which are most closely related to the axial forces in the drive shaft, are very similar to each other for various spider locations. On the other hand, there are noticeable differences in the vibrational characteristics at the knuckle along the axes. Figure 9 shows the vibrations in the vehicle with an Atype CV joint at the spider locations, as shown in Figure 4(b). Along the x and z axes, there are noticeable reductions in the vibration when the excitation from the engine is transferred to the knuckle. However, in the y axis, the reduction of vibration at the knuckle in the frequency range of 20~120 Hz is not as significant as in the other cases. This is probably because of the vibration in the y axis of the drive shaft, which is related to the idle vibrations (Sa et al., 2008). The aforementioned tendency was true in the B- and Ctypes of CV joints, as shown in Figure 10",
            " Therefore, unless direct measurements of the variations of the offset are achieved, it will not be easy to find a clear relationship between the offset and idle vibrations. It is believed that another independent testing machine might be necessary with a different testing protocol. It is meaningful, however, to review data that can lead to additional information on the vibrational characteristics at the knuckle for various spider positions before delicate testing is performed on the offsets of CV joints. Figure 12 shows the variation of the vibrational level measured at the knuckle with the spider locations under idling, as shown in Figure 4. The 24 data-points were collected at intervals of 15o with RMS values in the range of 0~300 Hz. They are connected with straight lines. It is clear that the vibrations are dominant in the z direction, which makes sense because the engine excitations are the strongest in that direction. However, considering the variation of the amplitudes of the RMS values during one complete cycle of the spiders, the vibration is substantive in the y direction as well. Recalling the poor vibrational characteristics in the y direction, as shown in Figure 11(b), it can be inferred that the high variation of the amplitudes in the y direction might be related to the complicated, periodic behavior of the offset, although these are indirectly related"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002400_1.4006324-Figure3-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002400_1.4006324-Figure3-1.png",
        "caption": "Fig. 3 FEM model and boundary conditions of the gears used for loaded tooth contact analysis",
        "texts": [
            " The web inclination angle (simply called web angle) is denoted by h as shown in Figs. 2(a)\u20132(c). When h\u00bc 0 deg, this indicates that the web is a straight web. Figure 2(d) is a solid gear used as the mating gear of these gears when they are engaged. The tooth numbers, the modules, the pressure angles, and the shifting coefficients of all the gears in Fig. 2 are denoted by Z1\u00bc Z2\u00bcZ3\u00bc Z4\u00bc 50, m\u00bc 4, a\u00bc 20 deg, and X1\u00bcX2\u00bcX3 \u00bcX4\u00bc 0, respectively. The structural dimensions of these gears are also shown in Fig. 2. 3.2 FEM Models and Boundary Conditions. Figure 3(a) is used to define the angle A that will be used in Sec. 6. Figure 3(b) is the FEM model used for the LTCA, deformation, and stress calculations of the thin-rimmed inclined web gears when they are engaged with the mating gear as shown in Fig. 2(d). Because the mathematical programming method [11\u201315] is used for the LTCA in this paper, it is only necessary to calculate the deformation influence coefficients and the gaps of the assumed contact point pairs on the contact tooth surfaces when conducting LTCA with the FEM. The simple procedure is as follows. First, a mathematical model used for LTCA is developed based on the principle of the mathematical programming method. Then, the deformation influence coefficients and the gaps of the assumed contact point pairs on the contact tooth surfaces are calculated with the FEM, and the models shown in Fig. 3(b) under the boundary conditions are given in Figs. 3(c) and 3(d). Finally, the tooth loads can be analyzed when the total load of the pair of gears, the deformation influence coefficients, and the gaps are substituted into the mathematical model by solving the equations of this model with the modified simplex method. More details can be found in Refs. [11\u201315]. When calculating the deformation influence coefficients of the assumed contact point pairs by the FEM using the models shown in Figs. 3(b)\u20133(d), the boundary conditions are obtained in the following ways. For the thin-rimmed gear, because the hub of the thin-rimmed gear is very thick relative to the web and the rim, any deformation of the hub is neglected and the entire gear structure except for the hub is used in the FEM model, as shown in Fig. 3(b). The joining part of the hub and the web of the gear are fixed in rotation and displacement in all degrees of freedom for the X, Y, and Z directions, as shown in Fig. 3(c). This boundary condition is also used for the deformation and stress analyses of the thinrimmed gear with the FEM. For the solid gear, only some of the 051001-2 / Vol. 134, MAY 2012 Transactions of the ASMECopyright VC 2012 by ASME Downloaded From: http://mechanicaldesign.asmedigitalcollection.asme.org/ on 01/28/2016 Terms of Use: http://www.asme.org/about-asme/terms-of-use teeth are used for the FEM model, as shown in Fig. 3(b). The three faces shown in Fig. 3(d) with the word \u201cFixed\u201d are fixed in the X, Y, and Z directions as FEM boundary conditions to calculate the deformation influence coefficients of the contact points on the tooth surfaces of the solid gear. It has been previously reported that this method of developing the FEM models and boundary conditions can produce suitable calculation results by experiments through comparing the calculated tooth root strain of the thinrimmed gear with a straight web with the measured strain [14]. The FEM models shown in Fig. 3(b) can be produced automatically for all of the gears and calculations in this paper with the software developed through the efforts of many years. The gearing parameters, the structural parameters, as well as the web angle can be changed freely in FEM modeling. LTCA is conducted for the three types of thin-rimmed gears shown in Fig. 2, when these gears have both straight webs and inclined webs, and they are engaged with the solid mating gear at the highest point of the single pair tooth contact",
            " Figures 14\u201316 show the calculation results of the left, center, and right web gears, respectively, with different web angles. The Journal of Mechanical Design MAY 2012, Vol. 134 / 051001-7 Downloaded From: http://mechanicaldesign.asmedigitalcollection.asme.org/ on 01/28/2016 Terms of Use: http://www.asme.org/about-asme/terms-of-use 051001-8 / Vol. 134, MAY 2012 Transactions of the ASME Downloaded From: http://mechanicaldesign.asmedigitalcollection.asme.org/ on 01/28/2016 Terms of Use: http://www.asme.org/about-asme/terms-of-use abscissas of Figs. 14\u201316 represent the circumferential angle A as shown in Fig. 3(a), and the ordinates of these figures are the equivalent stresses distributed along the joint circle of the rim and the web. From Figs. 14\u201316, it can be seen that the positions of the maximum joint stresses are changed from approximately 92 deg to approximately 100 deg when the web angle is increased. The maximum stresses are determined, and the relationship between the web angle and the maximum stress is shown in Fig. 17. From Fig. 17, it is apparent that the maximum joint stress is increased with an increasing increment in the web angle"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000395_1.2736451-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000395_1.2736451-Figure2-1.png",
        "caption": "Fig. 2 Bearing geometry and surface profile",
        "texts": [
            "org/ on 01/29/2016 Terms y h\u0304, = x h\u0304, 1 2a where C and r are the constants that can be found in Patir and Cheng 3 and is the surface pattern parameter. Assuming both journal and bearing surfaces having the same surface pattern i.e., J= b , the shear flow factor s is expressed as 4 s = 2V\u0304rj \u2212 1 s 2b where s is expressed as 4 s = A1 h\u0304 1e\u2212 2 h\u0304 + 3 h\u0304 2 for h\u0304 5 s = A2e\u22120.25 h\u0304 for h\u0304 5 2c A1 ,A2 , 1 , 2, and 3 are constants and can be obtained from Patir and Cheng 4 . For the journal bearing system shown in Fig. 2, assuming Gaussian distribution of surface heights, the expression for aver- age fluid-film thickness h\u0304T in fully lubricated i.e., for h\u0304 3 and partially lubricated i.e., for h\u0304 3 regions is expressed as 13 h\u0304T = h\u0304 for h\u0304 3 h\u0304 2 1 + erf h\u0304 2 + 1 2 e\u2212 h\u0304 2/2 for h\u0304 3 3 where h\u0304 is the nominal fluid-film thickness, i.e., the fluid-film thickness of a smooth journal bearing system and is expressed as h\u0304 = 1 \u2212 X\u0304J cos \u2212 Z\u0304J sin 4a where X\u0304J = \u0304 sin and Z\u0304J = \u2212 \u0304 cos 4b In the present work, the bearing performance characteristics have been analyzed under fully lubricated condition of the bearing. Restrictor Flow Equation. For an orifice restrictor, the equation of flow in nondimensional form is expressed 14 as \u00af \u00af \u00af 1/2 QR = CS2 1 \u2212 pc 5 Transactions of the ASME of Use: http://www.asme.org/about-asme/terms-of-use t b e c o l m l e t s l w i t s s s n a r c n l e p J Downloaded Fr Fluid-Film Velocity Components. For the thermal analysis, he flow of lubricant between two rough surfaces can be modeled y an equivalent flow model as shown in Fig. 2 c 12 . The quivalent model is defined as two smooth surfaces separated by a learance equal to the average gap h\u0304T . Based on the equivalence f flows through the average fluid-film thickness and through the ocal fluid-film thickness, a group of new pressure flow factors x , y and a shear flow factor s can be derived 12 as x = h\u03043 h\u0304T 3 x; y = h\u03043 h\u0304T 3 y and s = s 6 The mean or expected velocity components can be obtained by odifying the Poiseuille and Couette terms in the expression of ocal velocity components using the new flow factors and are xpressed in nondimensional form as u\u0304 = h\u0304T 2 x p\u0304 0 z\u0304 z\u0304 \u0304 dz\u0304 \u2212 F\u03041 F\u03040 0 z\u0304 dz\u0304 \u0304 + F\u03040 0 z\u0304 dz\u0304 \u0304 + s h\u0304TF\u03040 0 z\u0304 dz\u0304 \u0304 7a v\u0304 = h\u0304T 2 y p\u0304 0 z\u0304 z\u0304 \u0304 dz\u0304 \u2212 F\u03041 F\u03040 0 z\u0304 dz\u0304 \u0304 7b The fluid-film velocity component across the fluid-film is obained from the continuity equation and is expressed in nondimenional form as w\u0303 = w\u0304 \u2212 z\u0304u\u0304 h\u0304T = \u2212 0 z\u0304 h\u0304Tu\u0304dz\u0304 + 0 z\u0304 h\u0304Tv\u0304dz\u0304 7c Non-Newtonian Model"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001531_s11012-009-9232-0-Figure13-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001531_s11012-009-9232-0-Figure13-1.png",
        "caption": "Fig. 13 The prototype of the translator for a twist angle \u03b1 = 90\u25e6",
        "texts": [
            " 2 \u00b7 sin \u03d5 2 \u00b7 sin\u03b1|\u221a sin2 \u03d5 2 \u00b7 sin2 \u03b1 + (sin \u03b8 2 \u00b7 cos \u03d5 2 + cos \u03b8 2 \u00b7 sin \u03d5 2 \u00b7 cos\u03b1)2 (29) The previous relation permits to draw the variation of this distance according to the input angle \u03b8 (Fig. 11a) or to the distance b (Fig. 11b), knowing the variation of this distance (18) according to the input angle. These curves are drawn for three link lengths a = 40, a = 80 and a = 120, and for a twist angle \u03b1 = 150\u25e6. We can also draw the variation of the distance H according to the input distance b for different twist angles. Figure 12 shown this variation for three twist angles \u03b1 = 90\u25e6, \u03b1 = 120\u25e6 and \u03b1 = 150\u25e6. A prototype model (Fig. 13) was developed at the Applied Mechanical Laboratory in Besan\u00e7on (France), for a twist angle \u03b1 = \u03c0/2, and a second prototype with a twist angle \u03b1 = 150\u25e6 is under realization. In this paper we have presented a new and complete method for the calculation of closure equations for Wohlhart symmetric mechanism. This analytic method permits an accurate analysis of the mechanism\u2019s kinematics and gives a viable alternative to geometrical methods used by the other authors. All these calculus have allowed to observe and to demonstrate new geometrical and kinematical properties of these mechanisms"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000800_cnm.1051-Figure7-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000800_cnm.1051-Figure7-1.png",
        "caption": "Figure 7. A thick beam-type structure and the associate graph of the selected part.",
        "texts": [
            "1002/cnm In this section two examples with different topological properties are studied. The models are assumed to be supported in a statically determinate fashion. The effect of the presence of additional supports can separately be included for each special case with no difficulty [26]. The patterns of the null basis matrix B1 and the flexibility matrix G are formed for two examples, and the number of non-zero entries of these matrices is denoted by nz. Example 1 A thick beam-type structure supported in a statically determinate fashion is depicted in Figure 7. This structure is discretized using tetrahedron finite elements. The properties of the model are as follows: number of tetrahedron elements=480; number of nodes=205; elastic modulus E=2e+7kN/m2; Poisson\u2019s ratio =0.2; number of Type I self-stress systems=2032 (89.5%); first Betti number of the associate graph=317 (independent cycles); number of Type II self-stress systems=239; number of internal nodes (Ni )=39; DSIT =2271=(2032+239). Copyright q 2007 John Wiley & Sons, Ltd. Commun. Numer. Meth. Engng 2008; 24:1533\u20131551 DOI: 10"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002273_1.3680609-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002273_1.3680609-Figure2-1.png",
        "caption": "Fig. 2. The three types of ball used in this study: (a) smooth ball, (b) smooth ball modified by gluing a circular loop of string around the ball to simulate a raised seam, and (c) smooth ball modified by gluing a single length of string to the ball as an artificial baseball seam. Each ball was projected in the xdirection with backspin.",
        "texts": [
            " Marks and lines drawn on each ball were used to measure the spin of each ball to within 2%, either by plotting the rotation angle of the ball as a function of time or by counting the number of frames of the video film for the ball to rotate through a fixed number of revolutions. Properties of the four balls selected for this study are shown in Table I. The three polystyrene balls were nominally the same except one (ball 2) was fitted with a circular loop of string glued to the ball to simulate a straight seam, and one (ball 3) was fitted with an artificial baseball seam made from string and glued to the ball, as indicated in Fig. 2. In both cases, the string diameter was 1.5 mm. For ball 2, the string was offset from the center by a distance b\u00bc 30 mm. The baseball seam was scaled directly from measurements of the stitching on an actual baseball. Ball 1 was an unmodified polystyrene ball. The hollow plastic ball was smooth, apart from a small indentation used to inflate the ball. It was manufactured as a child\u2019s basketball and was slightly larger in diameter than an approved soccer ball (218\u2013221 mm). The balls listed in Table I were launched either by hand at relatively low speed and low spin or at higher speed and spin with a homemade lacrosse type ball launcher",
            " The impact point on the target could be measured to within 1 cm, but the horizontal deflection of the ball over the 5-m distance to the target could be measured to an accuracy of only about 9 cm, corresponding to an error of about one degree in the measured accuracy of the horizontal launch angle. In other words, a one-degree change in launch angle (from normal) corresponds to a 9-cm horizontal displacement in the impact point. Table I. Mass (M) and diameter (D) of the four balls used in the experiments. The ball type is shown in Fig. 2. No. Type Material M (g) D (mm) 1 A Polystyrene 8.98 101 2 B Polystyrene 12.15 98 3 C Polystyrene 11.55 100 4 A Hollow plastic 92 228 291 Am. J. Phys., Vol. 80, No. 4, April 2012 Rod Cross 291 Downloaded 11 Nov 2012 to 132.210.244.226. Redistribution subject to AAPT license or copyright; see http://ajp.aapt.org/authors/copyright_permission The horizontal deflection of each ball also depends on the orientation of the spin axis and on the orientation of the seam with respect to the spin axis. If the spin is not pure backspin then the ball can be projected sideways as a result of a sideways component of the Magnus force",
            " 4, April 2012 Rod Cross 294 Downloaded 11 Nov 2012 to 132.210.244.226. Redistribution subject to AAPT license or copyright; see http://ajp.aapt.org/authors/copyright_permission average acceleration in the negative vertical direction during that time was az \u00bc 12:5 m/s2, larger than g despite the fact that the drag force had a component acting vertically upward during this time. Results obtained with ball 2 are shown in Fig. 10. This ball was fitted with an artificial seam of string offset 30 mm from the center of the ball as indicated in Fig. 2(b). It was projected with backspin at speeds from 5 m/s to 17 m/s in an approximately horizontal direction and with the seam oriented as shown in Fig. 2(b). The results in Fig. 10 were obtained with the string on the left of center as viewed by the thrower. When the ball was projected at low speed with the string on the left, the ball deflected to the left, and vice-versa when the string was on the right. The ball also curved in a vertical direction as a result of the Magnus force and the force due to gravity, but the results in Fig. 10 show only the horizontal y-deflection (as defined in Fig. 4) or \u201cbreak\u201d after the ball travelled a horizontal distance of 5 m in the x-direction to the vertical target"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000746_1.2799524-Figure5-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000746_1.2799524-Figure5-1.png",
        "caption": "Fig. 5 \u201cActive\u201d and \u201cpassive\u201d nodes",
        "texts": [
            " The method is based on an exact relationship between uned and mistuned systems, which allows use of large finite ele- ent models, since only one sector is needed to represent the uned and mistuned systems, while the computational cost is inependent of the size of the original single blade segment. The rogram permits computation of the forced response at any seected so-called \u201cactive\u201d DOFs. In the current analysis, the ampliudes of forced response were obtained for all bladed disk sectors t four chosen nodes where the maximum displacements were nticipated. These nodes are shown as black circles in Fig. 5, hile the gray circles indicate the \u201cpassive\u201d nodes where the uniormly distributed loads are applied. A conventional engine order excitation by 3, 6, and 13 engine rders EOs was considered in the analysis over an excitation requency range corresponding to the predominantly first flapwise ibration mode 1F , as shown in Fig. 3 b by a rectangular area ig. 1 Forced response amplification factor as a function of lade frequency mistuning range onstrained between the two dashed horizontal lines. Forced re- 22501-2 / Vol"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001368_j.tcs.2009.01.033-Figure6-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001368_j.tcs.2009.01.033-Figure6-1.png",
        "caption": "Fig. 6. An illustration of Xi,j .",
        "texts": [
            " Consider an arbitrary homomorphism h : Hn(W np )\u2192 Hn(W nq ), and suppose, for 1 \u2264 i \u2264 p, h(\u03b1i) = \u2211q j=1 \u03bbi,j\u03b1j. Now we construct a map which induces h. Fix some i. LetMi = \u2211q j=1 |\u03bbi,j|, and ki : NMi \u2192 Nq be such that \u2211ki(j)\u22121 l=1 |\u03bbi,l| + 1 \u2264 j \u2264 \u2211ki(j) l=1 |\u03bbi,l| for all j \u2208 NMi . It always holds that \u03bbi,ki(j) 6= 0. Recall that Sn = {(t0, . . . , tn\u22122, tn\u22121 sin\u03b2, tn\u22121 cos\u03b2) \u2208 Rn+1|tn\u22121 \u2265 0, 2\u03c0 \u2265 \u03b2 \u2265 0, \u2211n\u22121 j=0 t 2 j = 1}. For 1 \u2264 j \u2264 Mi, let Xi,j be the subspace of Sn with 2(j \u2212 1)\u03c0/Mi \u2264 \u03b2 \u2264 2j\u03c0/Mi, as illustrated in Fig. 6. Let Xi = \u22c3Mi j=1 \u2202Xi,j, where \u2202Xi,j is the boundary of Xi,j. Consider the quotient space Yi = Sn/\u03bei, where \u03bei : Xi \u2192 \u2217. Let \u03c8i : Sn \u2192 Yi be the quotient map. See Fig. 7. Note the facts: \u2022 \u03c8i(Xi,j) is homeomorphic to Sn. \u2022 \u03c8i|X\u0307i,j : X\u0307i,j \u2192 \u03c8i(Xi,j)\u2212 \u03c8i(v0) is a homeomorphism, where X\u0307i,j is the interior of Xi,j. \u2022 Yi is the wedge sum of all \u03c8i(Xi,j) at \u03c8i(v0), 1 \u2264 j \u2264 Mi. Assume \u03b3i,j : \u03c8i(Xi,j) \u2192 Sn to be a homeomorphism. Without loss of generality, assume \u03b3i,j\u03c8i(v0) = v0. Define \u03b6i,j : S n \u2192 \u03c8i(Xi,j), v 7\u2192 { \u03c8i(v) if v \u2208 Xi,j \u03c8i(v0) otherwise"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003903_tdei.2015.005053-Figure16-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003903_tdei.2015.005053-Figure16-1.png",
        "caption": "Figure 16. Double-layer sleeve.",
        "texts": [
            " Figure 15 illustrates that, the installation of silicone rubber sleeve can decrease the surface electric field intensity near iron cap, but the maximum field intensity can reach 7.8 kV/cm at the pin side. When water drop existed on this area, surface discharge will occur. Semi-conductive rubber can be used to limit the surface electric field intensity. Considering the price of semi-conductor rubber is higher than that of silicone rubber, the double-layer rubber sleeve is designed. Its inner layer is semi-conductor rubber, and outer layer is silicone rubber, as shown in Figure 16. Calculation is carried out upon the electric field of insulators provided with three kinds of organic material sleeves such as silicone rubber sleeve, double-layer sleeve and semi-conductive rubber sleeve by means of ANSOFT software, respectively. The results are shown in Figure 17. It can be seen from Figure 17 that the electric field intensity near pin is much higher than that of other parts. More attention should be paid to this area. The local electric field intensity near pin is shown in Figure 18"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003520_j.euromechsol.2011.11.003-Figure4-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003520_j.euromechsol.2011.11.003-Figure4-1.png",
        "caption": "Fig. 4. T2R1-type parallel manipulator with decoupled and bifurcated planar-spatial motion of the moving platform: constraint singularity (a), branch with planar motion (b) and branch with spatial motion (c); limb topology PkRtPtkRtRPtPttRtR-PkRtPS.",
        "texts": [],
        "surrounding_texts": [
            "The term of constraint singularity (CS) has been recently coined (Zlatanov et al., 2002) to characterize the configuration of lower mobility parallel manipulators in which both the connectivity of the moving platform and the mobility of the parallel mechanism increase their instantaneous values. From a constraint singularity, the mechanism can get out with or without branching (Gogu, 2008b). When branching occurs in a constraint singularity, the mechanism can reach different configurations, called branches, and have different independent motions of the moving platform. In this case, the constraint singularity is also called branching or bifurcation singularity. A branch refers to the free-of-singularity configurations of the mechanism inwhich each structural parameter keeps its value. For this reason, this value is called global of full-cycle value for a branch. Two types of branching in constraint singularity (BCS) have been defined in Gogu (2008b). Branching of type BCS1 occurs when a parallel mechanism F ) G1 . Gj . Gk gets out from a constraint singularity in different branches characterized by the same mobility and the same connectivity of the moving platform but with different bases of the vector space of relative velocities between the moving and fixed platforms. Branching of type BCS2 occurs when a parallel mechanism F ) G1 . Gj . Gk gets out from a constraint singularity in different branches characterized by distinct values of mobility and connectivity of the moving platform. The parallel manipulators presented in this paper have constraint singularities with branching of type BCS1. Branching of type BCS2 occurs in kinematotropic mechanisms (Gogu, 2009c). The finite displacements and the velocities in the actuated joints are denoted by qi and _qi, the linear velocities of the characteristic point H of the moving platform, by v1 \u00bc _x; v2 \u00bc _y and the angular velocity of the moving platform by u \u00bc ua \u00bc _a or u \u00bc ud \u00bc _d: In both branches, the moving platform undergoes two planar translations and one rotation but the rotation axis is different in the two branches. In the first case, the rotation velocity u \u00bc ua \u00bc _a is parallel to the x-axis and the moving platform undergoes a spatial motion. In the second case, the rotation velocity u \u00bc ud \u00bc _d is parallel to the z-axis and the moving platform undergoes a planar motion. With these notations, the linear mapping between the actuated joint velocity space and the end-effector velocity space for a branch the T2R1-type PMs is defined by: 2 4 v1 v2 u 3 5 \u00bc \u00bdJ 2 4 _q1 _q2 _q3 3 5 (8) where J is the Jacobian matrix. In the constraint singularity, the T2R1-type parallel manipulators with bifurcated motions has instantaneously iM \u00bc iSF \u00bc 4 and (iRF)\u00bc(v1, v2, ua, ud). The bifurcation in this constraint singularity can be used to change motion type of the moving platform. In the two distinct branches, the parallel mechanism is characterized by M\u00bc SF \u00bc 3 and (RF)\u00bc(v1, v2,ua) or (RF)\u00bc(v1, v2,ud). In the constraint singularity, both connectivity of moving platform and mobility of parallel manipulator increase their instantaneous values. More details about the behaviour of parallel manipulators in a constraint singularity and the associated structural parameters can be found in Gogu (2009c)."
        ]
    },
    {
        "image_filename": "designv11_12_0003463_1.4004588-Figure9-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003463_1.4004588-Figure9-1.png",
        "caption": "Fig. 9 Contact line per lobe: (a) contact line attached to the driven rotor, (b) contact line attached to the driving rotor, and (c) contact line projected to the xoy plane",
        "texts": [
            " On the other hand, area of the blow hole from high pressure chamber to the low pressure chamber is also determined by the rotor profile. To decrease leakage to improve the performance of the pump, a long contact line should be avoided and area of the blow hole should be minimized. Thus, we calculate the length of contact line per lobe and area of the blow hole with different rotor profiles. For a given wrap angle and screw pitch, area of the blow hole and length of the contact line per lobe are totally determined by the transverse profiles of the rotors. The contact line can be figured out by coordinate transformation. Figure 9 demonstrates the contact line within one lobe. 3.3 Area of the Blow Hole. The blow hole is a small triangular-shaped area formed by the housing cusp and the lobe tips of the rotors. The blow hole could be figured out by locating two intersect points, P (P1 and P2 shown in Fig. 10), as the No. Generating segment (Driven rotor) Generated segment (Driving rotor) No. Generating segment (Driven rotor) Generated segment (Driving rotor) 1 Circle AB Circle GF 3 Elliptic arc CD Curve GI 2 Epicycloid BC Point G 4 Circle DE Circle IJ Journal of Mechanical Design SEPTEMBER 2011, Vol. 133 / 094501-3 Downloaded From: http://mechanicaldesign.asmedigitalcollection.asme.org/ on 09/04/2013 Terms of Use: http://asme.org/terms intersection points between the two rotor tips and the housing cusp and the third point Q locates on the contact line as shown in Fig. 9(c). These three points form a plane, which intersects the driving rotor and the driven rotor tips to form a triangular blow hole with two curved sides, as shown in Fig. 10. Lines P1P2 and P01P02 are the housing cusp lines, lines P1Q and P01Q are the intersecting curves between the helical surface of the driving rotor lobe and the blow hole plane, and lines P2Q and P02Q are the intersecting curves between the helical surface of the driven rotor lobe and the blow hole plane. Now, if point C (in Figs"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002752_j.proeng.2010.04.139-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002752_j.proeng.2010.04.139-Figure1-1.png",
        "caption": "Figure 1: The grip triad was constructed using a pipe clamp and 3 slender 3.5\u201d bolts and 3 reflective markers.",
        "texts": [
            " The 6 dof motion consists of the translational (X, Y and Z components) and rotational (3-1-3 Euler angles) position of the grip and its time derivatives (velocity and acceleration). The club kinematics were recorded using an 8-camera Vicon MX-F20 motion capture studio at the UW Kinesiology department. The infrared cameras were used to track the positions of 10 mm diameter reflective spheres attached to the golf club grip, shaft and head at 500 frames per second (fps). The grip\u2019s 6 dof motion, which serves as the main input to the simulation model, was measured using an array of 3 markers attached to the grip of the golf club (Fig. 1). Markers were also attached to the shaft and clubhead to record their motion for validation purposes. All of the calculations needed to extract the position and orientation of the club from the grip array assume that the relative positions of the array markers are fixed. Extracting the 6 dof grip motion to drive the simulation and the club face kinematics for validation required significant effort (Fig. 2). A summary of each step can be found below. 1. Correct Vicon Mislabelling Errors: Vicon tracks a series of unlabeled passive markers, and labels them by fitting a labelled template to the observed data"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000587_2007-01-2232-Figure16-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000587_2007-01-2232-Figure16-1.png",
        "caption": "Fig. 16 Mount Structure Comparison between 6-speed A/T and THS II for RWD",
        "texts": [
            " Also, the mounting position was shifted 80 millimeters farther forward than in the original plan in order to set it at a nodal point of the vibration mode (Fig. 15). Changing the mounting position raised new issues of installation space and the separation of principal elastic axis and the center of gravity, which strongly influences the engine start vibration performance. In order to ensure the adequate installation space, the mounting is embedded to the cross member, a major change from the structure used for a 6-speed A/T (Fig. 16). The distance of the principal elastic axis and the center of gravity was reduced by optimizing the lateral-to-vertical ratio of the mount spring constants. These countermeasures successfully addressed the issues of booming noise and engine start vibration (Fig. 17). The overall length of the transmission of the hybrid luxury sedan is greater than that of the THS II for FWD vehicles (Fig. 18), so the resonance that deforms the entire transmission is generated at a comparatively low frequency. Also, since the MG2 reduction gear ratio is low (Table 1), the resonance is generated by the 24thorder component of MG2 speed in the low vehicle speed range where there is less background noise (Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002093_j.triboint.2008.12.016-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002093_j.triboint.2008.12.016-Figure1-1.png",
        "caption": "Fig. 1. Schematic of the FZG test rig [6].",
        "texts": [
            " This section of the report gives a description of the FZG (Forschungstelle fu\u0308r Zahnra\u0308der und Getriebebau, English translation: research test rig for gears and transmissions) test rig. This was the main apparatus used while conducting the tests. An account is also given of the additional equipment that was used for obtaining the results. The FZG test rig is used for a variety of standardised and nonstandardised tests. It uses a recirculating power loop principle, allowing a fixed torque to be applied to a pair of precision test gears (refer to Fig. 1). Two torsional shafts connect the drive gearbox to the test gearbox. The system is subjected to torsional loading via the load clutch. This is done by loosening the clutch bolts and connecting the loading arm. Known weights are then hung from the loading arm and the clutch bolts are tightened. Hence a known torque is applied to the system. A set of test gears run by the FZG machine, comprises of a pinion and gear wheel (ratio of 1.5). They are commonly referred to as having type A profile. The gear teeth are finished with a special cross hatch grind pattern"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003713_978-1-4614-2419-2-Figure36.1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003713_978-1-4614-2419-2-Figure36.1-1.png",
        "caption": "Fig. 36.1 Geometry of a 2-D plate",
        "texts": [
            " Since the sensitivity is directly related to the wave structure, the preservations of wave structures can certainly maintain the defect detection sensitivity of the guided wave modes. For frequencies higher than the first cutoff, however, mode conversions may occur. The goal of the vibration mode decomposition work is to assess the defect sensitivity of plate vibration modes by the same transverse wave structure analysis approach that has proved to be successful for transient guided waves. a 2-D plate. The geometry of the 2-D Plate is shown in Fig. 36.1. It is assumed that the steady state vibration of a plate is a superposition of guided wave modes. Based on the orthogonality of guided wave modes, the stress field for the vibration can then be expressed as: s x; z\u00f0 \u00de \u00bc X m am sm\u00f0z\u00deei kmx ot\u00f0 \u00de \u00fe X m a0m s0m\u00f0z\u00deei k0m x L\u00f0 \u00de ot\u00f0 \u00de (36.1) Where s is the stress wave structure, the subscript m denotes the mth guided wave mode, am is the amplitude vector of the mth mode, o is the radian frequency, and k is the wave number, the prime sign differentiate guided wave modes propagating in the \u2013x direction with the ones propagating in the x direction"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003867_ascc.2015.7244739-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003867_ascc.2015.7244739-Figure1-1.png",
        "caption": "Fig. 1. Quadrotor reference frame system [23]",
        "texts": [
            " Due to the fact that quadrotors deal with a lot of disturbances and noise (vibrations from rotors), fuzzy logic is one of the intelligent control techniques that are capable of dealing with noisy inputs. The rest of the paper is structured as follows. System overview and modelling are presented in section II. Design of the controller is presented in section III. Section IV contains the results and discussion, and the concluding remarks are given in section V. The quadrotor is modelled in a cross configuration, with the centre of gravity coinciding with with the centre of mass as seen in Fig. 1 [23]. The four rotors can be controlled independently in order for the quadrotor to move within a three-dimensional space. Certain modelling assumptions need to be taken into consideration in order to mathematically model the quadrotor for simulation. The Earth is approximated as the inertial reference frame since it is stationary. Two reference frames are required to model the quadrotor UAV: the body-fixed frame and the inertial Earth-fixed frame. The inertial reference frame is required in order for Newton\u2019s equation of motion to be applicable"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002717_0954406212454390-Figure5-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002717_0954406212454390-Figure5-1.png",
        "caption": "Figure 5. Control volumes for the internal gear pump.",
        "texts": [
            " Here, displacement equals the fluid volume delivered by the pump each revolution and is usually used to evaluate the volumetric capacity of the pump. And, pulsation coefficient is used to describe the fluctuation level of the pump flowrate. Besides these two indices, volume of trapped fluid is also an important index to evaluate the performances of gear pumps.1 Derivation of the instantaneous flowrate formula The control volume approach is often used in the derivation of flowrate formula of gear pumps.3\u201312 Two control volumes, V1 and V2, are built for the pinion and the internal gear, respectively (Figure 5). With the assumptions of incompressible fluid, no fluid leakage and rigid parts of the pump, the discharge volume of fluid, dV, equals the difference between the input volumes, dVi1 and dVi2, and the output volumes, dVo1 and dVo2, as represented by equation (21). dV \u00bc dVi dVo \u00bc \u00f0dVi1 \u00fe dVi2\u00de \u00f0dVo1 \u00fe dVo2\u00de \u00f021\u00de where dVi1 \u00bc 1 2 r2a1d 1, dVo1 \u00bc 1 2 r2f1d 1, dVi2 \u00bc 1 2 r2a2d 2, dVo2 \u00bc 1 2 r2f2d 2: \u00f022\u00de Here, the axial thickness of the gear is a unit thickness. Given the fundamental law of gearing, it is known d 2 \u00bc r1 r2 d 1, "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003713_978-1-4614-2419-2-Figure1.2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003713_978-1-4614-2419-2-Figure1.2-1.png",
        "caption": "Fig. 1.2 Joints modeling",
        "texts": [
            " By using the developed model, dynamic simulations were performed to obtain the impact forces on the lander, penetration depth into the lunar soil, and tip-over characteristics. Figure 1.1 shows the design feature of the lunar lander considered in this paper. It mainly consists of lander body and landing gear assembly, and the latter includes struts, footpads and shock absorber (honeycomb type). The connection parts between center struts and main body are modeled as revolute joints as shown in Fig. 1.2. The honeycomb and strut are connected by a sliding joint. The force-displacement relationship of the shock absorber as shown in Fig. 1.3 was obtained by experiments and modeled by a nonlinear spring and damper element. The soil properties of lunar surface should be properly considered in the contact element between footpads and the lunar surface. The stiffness and damping properties of the lunar surface was determined by parametric simulation. Figure 1.4 shows some results from case studies for damping determination"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003509_j.proeng.2013.08.203-Figure6-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003509_j.proeng.2013.08.203-Figure6-1.png",
        "caption": "Fig. 6. Frobenius norm to evaluate the conditioning of the Jacobian matrix (a). Stiffness analysis (b)",
        "texts": [
            " As a first approach, to simplify the problem, the elements are considered as non-shrinking elements and the study focuses in a small region of the workspace where the Jacobian is well conditioned. Simplification can be done and the simplified relation between the Cartesian stiffness Kx, the joint stiffness K and the Jacobian J can be written: (3) A stiffness analysis has been realized on the IRB6660\u00ae with the method presented in Dumas et al. (2011). Firstly, the Frobenius norm allows to understand where the Jacobian matrix of the IRB6660\u00ae has a good conditioning (Fig. 6a). Secondly, different weights and configurations allow defining the stiffness of the 2nd to 6th axis by measuring the displacement produced and the forces felt inside the force sensor (Fig. 6b). Once, the stiffness is known for the 5 last joints, tests are then realized to analyze the stiffness of the first axis. A criterion rp is introduced to focus the displacement on the stiffest joints: (4) 1-- JKJK T x )max( 6 1 6 1 i i j rp q k k j i The kinematic performance criteria of manipulators allow defining the ability to move and to generate a given speed from its current position. Two matrices Wx and Wq, positive and diagonal, which allow to split the relative influence of the end effector speed and the articular speed (Dubey and Luh (1988))",
            " The general relations between the speed of the equivalent structure and the speed of the actuator are: 2/1-2/1 qxv JWWJ 1),( iiW x 2 max 1),( i q q iiW d T vv T dv uJJu 1)( ][][ 3221 1 32 A T Jqq ESA q I Jq 3 1 00 00 001 11 011 AJ and (9) The mechanical advantage performance allows defining the ability to transmit a force f along a d\u2019 direction. Two matrices Wf and W , positive and diagonal, which allow to split the relative influence of the end effector force and the articular torque are introduced (Dubey and Luh (1988)). The application allows defining the value of Wf and W . If a preferred direction is expected from the end effector, Wf can be weighted. However, the d\u2019 advance direction evolves during the path: Wf is so chosen as an identity matrix. Though, to take into account the maximal torque of the articulation imax (Fig. 6a), Jv is defined from the Jacobian matrix J as: and and and (10) A ms criterion concerning the speed capacity for the hybrid robot with parallelogram architecture is defined by: (11) with ud\u2019 a unit speed vector in the d\u2019 advance direction Rms ratio is the ratio between the operational force norm |f| and the articular torque norm | |: (12) Motor torque can be easily integrated into the model as explained above. Concerning the Tricept\u00ae, more investigation has to be taken into account to improve the posture towards the loading"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003887_j.procir.2015.06.103-Figure6-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003887_j.procir.2015.06.103-Figure6-1.png",
        "caption": "Fig. 6. Basic tooth profile of the double circular-arc profile",
        "texts": [
            " The transformation matrix from cS to 0S is 1000 0100 00cossin 00sincos 22 22 0 cM (17) The transformation matrix from 0S to 0S is 1000 0sincos0 0cossin0 0001 11 110 0M (18) The transformation matrix from 0S to 1S is 1000 0100 00cossin 00sincos 11 11 1 0M (19) The transformation matrix from cS to 1S is 1000 0 0 0 333231 232221 131211 00 0 1 0 1 aaa aaa aaa cc MMMM (20) Where 2112111 sinsinsincoscosa 2112112 cossinsinsincosa 1113 cossina 2112121 sinsincoscossina 2112122 cossincossinsina 1123 coscosa 2131 sincosa 2132 coscosa 133 sina (21) The parameter 1 represents the pitch cone angle of external spiral bevel gear. Similarly, the transformation matrix from cS to 2S in the Fig. 5 can be obtained as 1000 0 0 0 333231 232221 131211 00 0 2 0 2 bbb bbb bbb cc MMMM (22) Where 2232311 sinsinsincoscosb 2232312 cossinsinsincosb 2313 cossinb 2232321 sinsincoscossinb 2232322 cossincossinsinb 2323 coscosb 2231 sincosb 2232 coscosb 233 sinb (23) 2.2. Tooth Equation of the Double Circular-Arc Spiral Bevel Gear As shown in Fig. 6, the basic tooth profile of the spiral bevel gears in the normal section is a double circular-arc profile, and adopts the profile of the model GB 12759-91 as the basic tooth profile, which consists of eight sections. For the purpose of establishing the equation of the crown gear\u2019s tooth profile, the coordinate system ),,( nnnnS zyx is attached to the tooth profile in Fig. 6, and the coordinates of a point on the basic tooth profile can be illustrated in the coordinate system nS as 0 sin cos ,, iini iini T nininini Fr Er zyxr . (24) Where )8,,1(ii represents the number of arcs, ii FE , represents the circle centre of arc i , nir presents the arc radius and i presents the angle of arc i . As shown in Fig. 7, the coordinate system ),,( ccccS kji is attached to the crown gear and rotates about the axis ck of the centre of the gear. And the curve presents the actual boundary curve of the tooth profile; the curve presents the actual centre curve of boundary curve"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000399_s10999-008-9077-z-Figure21-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000399_s10999-008-9077-z-Figure21-1.png",
        "caption": "Fig. 21 Schematic of (a) disc and (b) fir-tree geometry",
        "texts": [
            " It is generally accepted that fretting initiated fatigue cracks are due to intermittent high frequency engine resonance and load fluctuations resulting from a change in engine speed or power requirements. It was therefore the objective of this study to evaluate the contact behaviour of turbine disc assemblies using the finite element method. Specifically, our attention was devoted to examining the effect of the critical geometric features upon the contact stress distribution at the different teeth of a fir-tree joint. These features, shown in Fig. 21, include the number of fir-tree teeth n, flank length l, contact angle a, flank angles b and c which define the tooth pitch. The fir-tree models were meshed with four-noded quadrilateral elements (Fig. 22). Plane stress conditions were assumed. No attempt was made to model the blade in details, except insofar as providing the necessary centrifugal loading at the interface. All models were subjected to centrifugal loadings. The material properties used for modelling the blade and the disc were that of a typical Nickel alloy used in disc design; namely, INCONEL 720"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000410_0094-114x(80)90020-8-Figure4-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000410_0094-114x(80)90020-8-Figure4-1.png",
        "caption": "Figure 4.",
        "texts": [
            "3) would not be satisfied for the same joint variable values. 3. A New Application We gave reasons for selecting the R-S-C-R- linkage as our prime example for application of the screw motor approach, but it could be objected that this loop may be readily and successfully analysed for uncertainty configurations by means of alternative algebraic methods. Such is not the case, however, for the R-C-C-R-R- chain, for which a stationary position analysis was first carried out in [2, 3]. The R-C-C-R-R- loop is illustrated in Fig. 4, and again the location of the fundamental frame of reference is carefully chosen. In [2, 3] it was shown that the stationary configurations for joint 5 could be determined by solving five simultaneous equations in the joint angles 0~- 0s. It was also indicated that the variables r2 and r3 could be subsequently found from the closure equations, the latter being given by 1 , r3s02 = -- sa-~23aM(c02c03 - s02s03c0123) + a23c02 + a12 q- a51c01 + R5sO1scl51 + a45(c05c0~ - sOssO~cas~) + R4(s05cOl so45 + cOssOi sa45ca51 + s01 ca45sasO}"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000645_s00170-007-1183-9-Figure5-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000645_s00170-007-1183-9-Figure5-1.png",
        "caption": "Fig. 5 An example of FEM simulation for rotational mode of vibration about X axis",
        "texts": [
            " The results are given in Table 2. In order to verify the aforementioned results, the natural frequencies of the upper platform are also obtained by FEM. For this purpose, the solid model developed in CATIA is exported to the Ansys software for modal analysis. The discretized model in the Ansys environment with the relevant boundary conditions is shown in Fig. 4. The results for the same ten positions of the platform are also given in Table 2. As an example, the rotational mode of vibration about X axis is shown in Fig. 5. It is noteworthy that the computational time for each solution has amounted to 3 h. As is evident from Table 2, the results obtained from the analytical model and FEM are in agreement. This can be better visualized from the comparative diagrams of Fig. 6. Some important inferences can be drawn from the foregoing results. The lowest natural frequencies in all modes of vibration occur when the moving platform takes the higher positions. This implies that high cutting speeds should be used in order to avoid dynamic instability, and the hexapod table is most suitable for high-speed machining and finishing operations"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003990_wcica.2012.6359128-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003990_wcica.2012.6359128-Figure1-1.png",
        "caption": "Fig. 1. Solidworks drawing of the helicopter",
        "texts": [
            " The goal is achieved by taking the advantage of linear quadratic regulator (LQR) control approaches [7], which are favorable for solving the control problems of multi-variable system in the manner of optimal profile. Extensive knowledge and comprehensive understanding in helicopter modeling are essential to accurately capture the helicopter\u2019s aerodynamics so as to design the desirable control system. First-principles modeling approach as the fundamental and physics-based one is employed to obtain the nonlinear model with modest level of complexity. The effortless extraction of physical parameters is accessible by employing the faithful Solidworks model of the helicopter shown in Fig. 1. The derived high-fidelity nonlinear model is sufficient for simulation and the linearized model is adequate for the controller design. A nonlinear model of the UAV helicopter is generated from first-principles modeling approach. A fairly accurate model is desired, whereas complexity in the dynamic equations increases with the model accuracy. The helicopter model is 3927 978-1-4673-1398-8/12/$31.00 \u00a92012 IEEE developed with medium-complexity based on only basic flight dynamics. Satisfactory accuracy of the model can be achieved and the obtained model is adequate for flight simulations over a large portion of the flight envelope"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003353_iros.2012.6385499-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003353_iros.2012.6385499-Figure2-1.png",
        "caption": "Fig. 2. Simulation model of combined rimless wheel with wobbling mass",
        "texts": [
            " We discuss the effect mainly from the viewpoint of CoM trajectory, and investigate the potentiality of speeding-up of PDW through numerical simulations and experiments. Fig. 1 shows the overview of our prototype CRW with a wobbling mass that moves up-and-down passively along the guide rail in the body frame. The two RWs are connected via a rigid rod so that they mutually synchronize or the phase difference is kept zero during motion. LEDs are attached on the wobbling mass and body frame for clearly observing the oscillation. In the following, we describe the mathematical model. Fig. 2 shows the ideal model. For simplicity, the rear RW is called \u201cRW1\u201d, and the fore one is called \u201cRW2\u201d. The CRW model is identical to that in our previous work (see [4] for further detail). We then add a passive wobbling mass to the body frame. Let qi = [ xi zi \u03b8i ]T be the generalized coordinate vectors of the RWs (i = 1, 2) and body frame (i = 3). We then define the generalized coordinate vector of the whole system, q \u2208 R 10, as q = [ qT 1 qT 2 qT 3 Lc ]T where Lc is the length between the wobbling mass and body frame"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002619_app.1965.070090425-Figure4-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002619_app.1965.070090425-Figure4-1.png",
        "caption": "Fig. 4. Chart record from photoelectric turbidmeter X-Y recorder. M , / M n of polymer = 8.0.",
        "texts": [
            " Two variable potentiometers are placed in series with the motordriven potentiometer so that the cooling rate may be changed from O.Ei\u00b0C./min. to 2.89'C./min. The temperature and turbidity of the solution are recorded by an X-I.' recorder, the temperature signal coming from the thermocouple and the turbidity signal coming from the photoelectric bridge circuit as shown in Figure 3. Electronic span adjustments make it possible to select any level of sensitivity desired. The shape of the turbidity plot is that of an S. It will be seen from Figure 4 that a typical plot actually does resemble the integral molecular weight distribution expected from a cumulative distribution. It will be noted that in the middle of the S-shaped plot the slope is quite constant and can be represented by a straight line. Various samples having different molecular weight distributions show a different slope at this middle portion of the turbidity curve as well as a different shape on the two ends. A parameter designated as S was chosen to be the difference in temperature between points representing 20% of the maximum turbidity and 50% of the maximum turbidity as suggested by Taylor"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003863_icuas.2015.7152374-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003863_icuas.2015.7152374-Figure1-1.png",
        "caption": "Figure 1. An undirected graph for the topology of multi-UAS with five agents.",
        "texts": [
            " Next, an adjacency matrix , , , 1,2N N i jA a i j      ,..., N with element ,i ja is defined as 1ija  if  ,i j E . If there exists a path connecting each pair of distinct vertices, the graph is connected. Each agent in multi-UAS is assumed to have the equal omni-directional sensing capability which indicates that there is a mutual sensing among the connected multi-UAS. Moreover, the adjacency matrix is symmetric, i.e. TA A , and the graph is undirected. An undirected graph topology example is shown in Figure 1. Further, the degree matrix of graph G can be defined as { }D diag A where the diagonal element ijd is derived as 1, N ij ij j j i d a     . The Laplacian matrix   N N ijL lp   , ,i j  1,2..., N can be represented as L D A  . In Figure 1, a multi- UAS including five UAS is demonstrated as an undirected graph. Next, the optimal flocking control problem for multi-UAS is formulated. III. PROBLEM FORMULATION For the sake of simplification, several recent literatures [5],[6],[7] supposed that multi-UAS have the homogeneous simple system dynamics such as a double integrator. However, due to the uncertainty and nonlinearity, this assumption is not suitable for most practical UAS. Therefore, in this paper, the more realistic heterogeneous nonlinear dynamics are considered as ( , ) ( , ) i i i i i i i i i i p v v f p v g p v u    , 1,2,"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000597_978-3-540-73129-0-Figure6.109-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000597_978-3-540-73129-0-Figure6.109-1.png",
        "caption": "Fig. 6.109. Linear contractile dielectric elastomer actuators: a helical and b folded configuration",
        "texts": [
            " The simplest one, from a conceptual point of view, consists of a stack of elementary actuating units, made of planar actuators connected in electrical parallel and mechanical series [280, 290]. The thickness contraction of each element causes the axial contraction of the entire structure. This configuration can enable very interesting per- formances [290]. Nevertheless, its discontinuous structure can complicate its fabrication. Therefore, new solutions for contractile actuators may be of help. As an example, two types of new configurations have been recently presented, as shown in Fig. 6.109. The first is termed an helical dielectric elastomer actuator (Fig. 6.109a) [291]. It consists of a hollow cylinder of dielectric elastomer, having two helical compliant electrodes integrated within its wall. The second is termed a folded dielectric elastomer actuator (Fig. 6.109b) [292]. It is made of a monolithic strip of electroded elastomer which is folded up. For both these configurations, a high voltage difference applied between the electrodes induces attractions among opposite charges of the two electrodes, as well as repulsions among the same type of charges of each electrode: accordingly, these effects determine the compression of the dielectric included between the electrodes, causing an axial contraction and a radial expansion of the structure. Such devices might be useful for applications requiring spring-like contractions of an elastomeric device activated and modulated by an electrical signal"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002319_j.jsv.2012.05.009-Figure8-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002319_j.jsv.2012.05.009-Figure8-1.png",
        "caption": "Fig. 8. Snapshots of the flapping motion at w\u00bc0.8 with the time interval of 0.25: (a) t\u00bcT0; (b) t\u00bcT0\u00fe0.25; (c) t\u00bcT0\u00fe0.5; (d) t\u00bcT0\u00fe0.75; (e) t\u00bcT0\u00fe1.0; (f) t\u00bcT0\u00fe1.25. (Re, G , rr, l)\u00bc(500, 50, 16, 30), Lz\u00bc2 (BBC).",
        "texts": [
            " 6, indicating a very narrow plate should not exhibit apparent asymmetrical flapping motion. Figs. 7\u20139 show the snapshots of the flapping motion with the time interval of 0.25 for w\u00bc0.6, 0.8 and 1.4, respectively. Fig. 7 shows the asymmetrical flapping in the spanwise direction at w\u00bc0.6, and the plate often exhibits the spanwise warping deformation. For w\u00bc0.8 and 1.4, the flapping is symmetrical, but the stronger local deformation can be observed. For w\u00bc0.8, the strong deformation appears at the center region in the spanwise direction, as shown in Fig. 8, whereas for w\u00bc1.4, the deformation mode becomes complex: two spanwise waves can be observed at the free end, but only one spanwise wave is observed in the middle region (Fig. 9). The vortex structures in the wake of the flapping plate for w\u00bc0.6, 1.0 and 1.6 are shown in Fig. 10. The lci criteria (i.e., the imaginary part of the complex eigenvalue of the velocity gradient tensor) proposed by Zhou et al. [42] is employed for the vortex identification. It is not surprising that the hairpin-shaped vortices predominate in the wake"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000388_14644193jmbd97-Figure12-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000388_14644193jmbd97-Figure12-1.png",
        "caption": "Fig. 12 Forces on the spindle system",
        "texts": [
            " Rolling bearing analysis, 2001 (John Wiley & Sons, USA). 35 Aini, R., Rahnejat, H., and Gohar, R. An experimental investigation into bearing induced spindle vibration. Proc. Instn Mech. Engrs, Part C: J. Mechanical Engineering Science, 1995, 209, 107\u2013114. 36 Vafaei, S., Rahnejat,H., andAini, R.Vibrationmonitoring of high speed spindles using spectral analysis techniques. Int. J. Mach. Tool Manuf., 2002, 42, 1223\u20131234. APPENDIX Notation a a distance between the external force and the first bearing (Fig. 12) (m) a1 a distance between the gravitational centre of shaft and the first bearing (Fig. 12) (m) A distance between centers of curvature of inner and outer races grooves (m) b1 a distance between the gravitational centre of shaft and the second bearing (Fig. 12) (m) B total curvature d diameter (m) E modulus of elasticity (N/m2) g acceleration due to gravity (m/s2) K stiffness factor (N/m3/2) m number of balls M mass of spindle (kg) n rotational speed of the spindle (r/min) r radius (m) t time (s) W restoring force (N) x, y, z fixed coordinates defining spindle centre movement and deflections along these coordinate axes (m) a contact angle (rad) g angle between two adjacent ball (rad) d contact deflection (m) z damping ratio u angular position of a ball (rad) r curvature (1/m) v angular velocity (rad/s) q the angle between the fixed and moving reference axes (Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000470_j.ijfatigue.2007.12.003-Figure9-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000470_j.ijfatigue.2007.12.003-Figure9-1.png",
        "caption": "Fig. 9. A finite element model for residual stress analysis.",
        "texts": [
            " In this paper, therefore, the criterion for surface removal to obtain the maximum contact fatigue life considering residual stress change is developed by conducting finite element analyses. As a result, the optimal depth for surface removal considering the change of residual stress is proposed. A wheel manufactured for a freight train (SSW1) is chosen for the analysis, and a corresponding axisymmetric model is used to adopt the symmetric characteristics of wheel geometry and loading conditions. Fig. 9 shows a finite element model used for stress analysis. The loading condition and analysis procedure are adopted from the previous study [12]. The compressive residual stress in the circumferential direction is observed as expected. Since the wheel is designed to maintain a constant thickness up to a certain limit for the reprofilling process which requires contact surface removal, the wheel surface thickness is continuously reduced due to the surface removal and the wear over the lifetime. Since the residual stress on contact surface is redistributed due to the surface removal and the thickness reduction, it should be considered to decide the optimal surface removal depth"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002185_bit.260090407-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002185_bit.260090407-Figure1-1.png",
        "caption": "Fig. 1. Culture vessel dimensions (in cm.). Dotted line indicates extent of travel of sparge pipe. At its topmost position (0% \u201cimmersion\u201d) the end of the sparge pipe is 10.5 cm. above the bottom of the vessel. A t the lowest position ( 100~o immersion) the distance is 3 cm.",
        "texts": [
            " In order to obtain the relatively slow control movement meritioncd abovc, suitable gearing is inserted between the sprtrge pipe motor and the sparge pipe itself. As an on/off controller for the system we have used both single and multichannel recorders with either mercury switches or cam-operated microswitches. These systems provide high and lorn level control with variable deadband. Culture Vessel A 5-liter fermenter of the type described by Elsworth et aL8 was used for both batch and continuous culture. The dimensions of the vessel are shown in Figure 1. The culture (3 liters) was stirred a t 880 rpm and the temperature maintained at 30 f 0.5\"C. The pH of the medium was controlled at 7.15. The air and nitrogen werc separately metered into the vessel arid sterilized by filtration. The oxygen probe was inserted through the base plate of the vessel. The filtered and dried effluent gas was supplied to an infrared type carbon dioxide analyzer (model SBK, Hilger & Watts Ltd., London NWl) and to a paramagnetic oxygen analyzer (model OA 137, Servomex Controls Ltd"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001307_tie.2009.2021588-Figure7-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001307_tie.2009.2021588-Figure7-1.png",
        "caption": "Fig. 7. Experimental setup of impaired cooling conditions. (a) Impaired cooling condition with cooling fan removed. (b) Impaired cooling condition with additional thermal insulation.",
        "texts": [
            " 6. To test the feasibility of the proposed cooling-condition monitoring system, three different cooling conditions are tested: 1) the motor is first operated under normal cooling condition; 2) the motor fan is removed to emulate the broken cooling fan condition; 3) a fiberglass thermal-insulation foil is used to cover part of the induction-motor frame to emulate the impaired cooling caused by motor-frame dust build-up. The inductionmotor setups under the impaired cooling conditions are shown in Fig. 7. The stator-winding temperature-estimation results under different cooling conditions and different load conditions are shown in Fig. 8. The estimated stator-winding temperature is then filtered to reduce the estimation error, using a statisticalmodel-based Kalman filtering technique, as proposed in [29]. Fig. 8(a)\u2013(c), respectively, shows the Ts estimation results under three different cooling conditions: the normal cooling condition, the impaired cooling condition with a broken cooling fan, and the impaired cooling condition with additional thermal insulation"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003702_cjme.2013.04.801-Figure14-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003702_cjme.2013.04.801-Figure14-1.png",
        "caption": "Fig. 14. Analysis model for maximum static torque of abductor joint",
        "texts": [
            " Based on the machine characteristics of the robot and the need to prevent the bearing platform from rolling, yawing, and pitching, a range of \u201330\u00b0 to 30\u00b0 is confirmed for \u03b82, \u03b84, and \u03b86. The foothold of leg 2 is defined as point 2o . Point 4o represents the foothold of leg 4. Point 6o is the foothold of leg 6. f2, f4, and f6 are respectively considered to be the frictions on footholds 2o , 4o , and 6o . Based on Fig. 13, some relations can be obtained: f2  2xF , f4  4xF , and f6  6xF . The analysis model for the maximum static torque of the ZHUANG Hongchao, et al: Method for Analyzing Articulated Torques of Heavy-duty Six-legged Robot \u00b7808\u00b7 abductor joint is shown in Fig. 14, where d2 is the length between foothold 2o and the vector of force 6xF . The distance from foothold 2o to the vector of force 4xF is expressed as d3. do represents the distance from foothold 2o to the vector of the G component in the direction of x. The arm of the force, whose length is from the vector of force 2xF to the axis of the abductor joint of leg 2, is regarded as d1. Thanks to the equal poses between legs 4 and 6, length d2 is equal to length d3. Based on the above analysis, the matrix of the distance can be obtained and written as r bpr d Rl c , (3) where d is a matrix of the distance, R is a matrix of the rotation angles for the joints, l is a matrix of the length in one leg, cr is a matrix of the coefficient for the radius of the bearing platform, and they can be expressed as follows: 2 2 2 2 2 2 6 2 6 6 2 2 4 2 4 4 2 2 2 2 2 2 6 6 2 2 4 4 2 2 cos cos cos cos cos cos cos cos cos cos cos cos cos cos cos cos cos cos cos cos cos cos , cos cos cos cos cos cos cos cos \u03b8 \u03b2 \u03b8 \u03b8 \u03b2 \u03b8 \u03b8 \u03b8 \u03b2 \u03b8 \u03b2 \u03b8 \u03b8 \u03b8 \u03b2 \u03b8 \u03b2 \u03b8 \u03b2 \u03b8 \u03b2 \u03b8 \u03b2 \u03b8 \u03b2 \u03b8 \u03b2 \u03b8 \u03b2 \u03b8              R T c t s( ) ,l l ll T r 3 31 0 , 2 2      c T o 1 2 3( ) "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000155_pime_proc_1976_190_050_02-Figure4-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000155_pime_proc_1976_190_050_02-Figure4-1.png",
        "caption": "Fig. 4 The surface tractions relationship",
        "texts": [],
        "surrounding_texts": [
            "J . HALLING 486\n6 .\n7.\n8.\n9 .\nHALLING, J. A C o n t r i b u t i o n t o t h e 1 2 . TAVERNELLI I \" A C o m p i l a t i o n a n d T h e o r y o f F r i c t i o n . J .F . & COFFIN, I n t e r p r e t a t i o n o f\nU.S.M.E. /T/40/75. T e s t s \" L.F. C y c l i c / S t r a i n F a t i q u e\nTABOR, D . T h e H a r d n e s s o f M e t a l s O x f o r d U n i v e r s i t y P r e s s ,\nT r a n s . A m SOC. Me ta l s , 5 1 , 4 3 8 , 1 9 5 9 .\nR e l a t i o n s h i p b e t w e e n 1 9 5 1 . 1 3 . SUH,N.P. &\nSRIDHARAN,P. t h e c o e f f i c i e n t o f JEFFERIS,M.A. & JOHNSON,K.L. T r a c t i o n s i n E l a s t o -\nF r i c t i o n a n d t h e Wear R a t e o f Meta ls . h y d r o d y n a m i c s L u b r i c a t -\ni o n .\nP r o c . I . Mech.E., 1 8 2 , B u t t e r w o r t h P r e s s , P t . l , N o . 1 4 , 2 8 1 , 1 9 7 0 . L o n d o n , 1 9 6 5 .\n1 4 . K R A G E L S K I I , I . V . F r i c t i o n a n d Wear\nTIMOTHENK0,S T h e o r y of E l a s t i c i t y 1 5 . HALLIDAY,J.S. S u r f a c e E x a m i n a t i o n b y & G O O D I E R , J . N . M c G r a w - H i l l , 1 9 5 1 . R e f l e c t i o n E l e c t r o n\nMicros C O D Y - 10. HALLING,J. & A S t a t i s t i c a l M o d e l f o r\nEL-REFAIE,M E n g i n e e r i n g S u r f a c e s I n s t n . Mech. E n g r s . T r i b o l o g y C o n v e n t i o n , 1 9 7 1 .\nProc .1 .Mech.E. 1 9 6 9 , 7 7 7 , 1 9 5 5 .\n11. G R E E N W O O D , J . A . C o n t a c t s Of N o m i n a l l y This paper is presented for written discussion. The MS was received on & WILLIAMSON, F l a t S u r f a c e s . J .B .P . 1st November 1975 and accepted forpublication on 10th June 1976. 33\nP r o c . Roy. SOC. , A , 2 9 5 Communications are invited for publication in the hceedings. 3 0 0 , 1 9 6 6 . Contributors should read the instructions on page ii of cover.\nRoc Instn Mech Engrs Vol190 43/76 IMechE 1976\nat WEST VIRGINA UNIV on June 5, 2016pme.sagepub.comDownloaded from",
            "Roc Imtn Mech Engra Voll90 43/76\nat WEST VIRGINA UNIV on June 5, 2016pme.sagepub.comDownloaded from",
            "J. HALLING\nFig. A.l: The relation between u and q\nRoc lnstn Mech Engrs Vol190 43/76\nat WEST VIRGINA UNIV on June 5, 2016pme.sagepub.comDownloaded from"
        ]
    },
    {
        "image_filename": "designv11_12_0000058_j.triboint.2005.11.022-Figure5-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000058_j.triboint.2005.11.022-Figure5-1.png",
        "caption": "Fig. 5. Variation of non-dimensional bush temperature (Tb) along circumferential direction for e \u00bc 0:1, L=D \u00bc 1, a \u00bc 70, c\u0304g \u00bc 0:1, ng \u00bc 8, d \u00bc 1.",
        "texts": [
            " The pressure distribution is wave like structure due to the grooves, whose average value is similar to plain journal bearing due to the hydrodynamic action. The temperature distribution at the middle layer of the fluid film is shown in the Fig. 4. Due to the grooves the fluid temperature forms wave like pattern, similar to that of pressure. The temperature of the fluid film is maximum corresponding to the maximum pressure zone. Variation of bush temperature at middle section of the bush is shown in Fig. 5. It is also observed that the pattern of bush temperature is similar to the temperature profile of the fluid film. Due to the rotation of the shaft, the shaft temperature along y direction remains constant as shown in Fig. 6 and it gradually increases from the end towards ARTICLE IN PRESS M. Sahu et al. / Tribology International 39 (2006) 1395\u20131404 1401 the middle section of the bearing. Fig. 7 shows the variation of maximum non-dimensional temperature of fluid, bush and shaft with eccentricity ratio"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001884_s11431-010-3100-y-Figure4-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001884_s11431-010-3100-y-Figure4-1.png",
        "caption": "Figure 4 3-RPS mechanism.",
        "texts": [
            " This paper will introduce the application of method of virtual loop through several mechanisms. Like the examples in this paper, in most cases, dA,B j1 can be obtained by observation as mentioned above. However, due to the complicated structures of mechanisms and dA,B j1 dj, the determination of dA,B j1 is easier than dj in some cases, and in few cases, determination of dA,B j1 is more complex than dj. The more general method for determining dA,B j1 still waits for more systematic investigation, which is beyond the scope of this paper. Example 1. Figure 4 is a 3-RPS mechanism, where the three link groups are uniformly distributed. In each link group, the R-axis is perpendicular to the prismatic pair P. In loop I, there is no common constraint, here dX I = 6, i.e., FI =PI i=1fi dX I =106 = 4. The moving platform has three rotations, three translations and a local freedom around CD-axis, therefore, d3,6 I (x y z)=6. In loop II, the link group GHK cannot translate along x-, y-axes, the rank is dgz II (0 y z)=5, so dgz II = d3,6 I (x y z)+dX II (0 y z) = dX II(x y z)=6, FII =Pj i=1fi dX II=56=1, F=FI +FII = 4+(1)=3"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003761_9781118818428-Figure1.2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003761_9781118818428-Figure1.2-1.png",
        "caption": "Figure 1.2 How the quarry engine might have appeared after inversion of the cylinder unit to allow heating to a viable operating temperature",
        "texts": [
            " Stirling myth \u2013 and Stirling reality 3 Stirling\u2019s own hand-written description cites temperature difference (between upper and lower extremes of the cylinder) of 480 \u25e6F \u2013 or 297 \u25e6C. If ambient temperature were 30 \u25e6C this would require the hottest part of the cylinder to be at 337 \u25e6C. The quarry engine doubtless functioned \u2013 but not in the elegant configuration of the patent drawing. A drawing of an enginewhich, by contrast, is readily reconciledwith brother James Stirling\u2019s retrospective (1852) account is shown at Figure 1.2: flywheel, linkmechanism, cylinder, piston, and displacer are re-used. However, the entire assembly now operates upside down relative to the patent illustrations, with cylinder head immediately above the furnace and the flame in direct contact. Achieving 1\u20442 hp no longer requires Stirling to have optimized the thermal and flow design of the regenerator. Brother James\u2019 account of the eventual failure of the engine now makes sense: \u2018\u2026 the bottom of the air vessel became over-heated.\u2019 Nothing could be further from the truth"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003520_j.euromechsol.2011.11.003-Figure8-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003520_j.euromechsol.2011.11.003-Figure8-1.png",
        "caption": "Fig. 8. Maximally regular T2R1-type parallel manipulator with bifurcated planarspatial motion of the moving platform: constraint singularity (a), branch with planar motion (b) and branch with spatial motion (c); limb topology PkRtPtkRtRPaPattPtkPtRtRSPtR.",
        "texts": [],
        "surrounding_texts": [
            "The term of constraint singularity (CS) has been recently coined (Zlatanov et al., 2002) to characterize the configuration of lower mobility parallel manipulators in which both the connectivity of the moving platform and the mobility of the parallel mechanism increase their instantaneous values. From a constraint singularity, the mechanism can get out with or without branching (Gogu, 2008b). When branching occurs in a constraint singularity, the mechanism can reach different configurations, called branches, and have different independent motions of the moving platform. In this case, the constraint singularity is also called branching or bifurcation singularity. A branch refers to the free-of-singularity configurations of the mechanism inwhich each structural parameter keeps its value. For this reason, this value is called global of full-cycle value for a branch. Two types of branching in constraint singularity (BCS) have been defined in Gogu (2008b). Branching of type BCS1 occurs when a parallel mechanism F ) G1 . Gj . Gk gets out from a constraint singularity in different branches characterized by the same mobility and the same connectivity of the moving platform but with different bases of the vector space of relative velocities between the moving and fixed platforms. Branching of type BCS2 occurs when a parallel mechanism F ) G1 . Gj . Gk gets out from a constraint singularity in different branches characterized by distinct values of mobility and connectivity of the moving platform. The parallel manipulators presented in this paper have constraint singularities with branching of type BCS1. Branching of type BCS2 occurs in kinematotropic mechanisms (Gogu, 2009c). The finite displacements and the velocities in the actuated joints are denoted by qi and _qi, the linear velocities of the characteristic point H of the moving platform, by v1 \u00bc _x; v2 \u00bc _y and the angular velocity of the moving platform by u \u00bc ua \u00bc _a or u \u00bc ud \u00bc _d: In both branches, the moving platform undergoes two planar translations and one rotation but the rotation axis is different in the two branches. In the first case, the rotation velocity u \u00bc ua \u00bc _a is parallel to the x-axis and the moving platform undergoes a spatial motion. In the second case, the rotation velocity u \u00bc ud \u00bc _d is parallel to the z-axis and the moving platform undergoes a planar motion. With these notations, the linear mapping between the actuated joint velocity space and the end-effector velocity space for a branch the T2R1-type PMs is defined by: 2 4 v1 v2 u 3 5 \u00bc \u00bdJ 2 4 _q1 _q2 _q3 3 5 (8) where J is the Jacobian matrix. In the constraint singularity, the T2R1-type parallel manipulators with bifurcated motions has instantaneously iM \u00bc iSF \u00bc 4 and (iRF)\u00bc(v1, v2, ua, ud). The bifurcation in this constraint singularity can be used to change motion type of the moving platform. In the two distinct branches, the parallel mechanism is characterized by M\u00bc SF \u00bc 3 and (RF)\u00bc(v1, v2,ua) or (RF)\u00bc(v1, v2,ud). In the constraint singularity, both connectivity of moving platform and mobility of parallel manipulator increase their instantaneous values. More details about the behaviour of parallel manipulators in a constraint singularity and the associated structural parameters can be found in Gogu (2009c)."
        ]
    },
    {
        "image_filename": "designv11_12_0003761_9781118818428-Figure17.3-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003761_9781118818428-Figure17.3-1.png",
        "caption": "Figure 17.3 Measures for minimizing effects of \u2018jetting\u2019: plenum volume and provision for tangential entry",
        "texts": [
            " The result will be credible if the flow itself can reasonably be considered one-dimensional \u2013 or at least axi-symmetric. Free-flow areaAffxe of the expansion exchanger tube bundle can be an order of magnitude less than Affr of the regenerator. For the P-40 Affxe\u2215Affr = 0.0328, or 3.2%! There is incontrovertible evidence that flow from an individual tube of the expansion2 exchanger emerges as a jet penetrating several layers of the gauze stack before dissipating. The phenomenon is evidently mitigated somewhat by provision of a plenum volume between tube exit and the first gauze of the stack (Figure 17.3). The matter is far from incidental: Feulner (2013) recounts experiments 1Aiming for \u00b6v > 0.75 minimizes likelihood of need for remedial action later when pumping penalty is estimated. 2The problem cannot be verified by the same means at the the compression end, but is likely to be less severe, as total inflow is distributed over a far larger number smaller tubes. 174 Stirling Cycle Engines Wire-mesh regenerator \u2013 \u2018back of envelope\u2019 sums 175 176 Stirling Cycle Engines with engines which, with the exchanger end-plate in contact with the first gauze layer, failed to run but which, after introduction of a gap, operated satisfactorily. Stirling engine science would advance more rapidly if Peter Feulner could find time from his management responsibilities to publish an account of his impressive design and experimental work. Even when a plenum volume is provided, a degree of \u2018jetting\u2019 evidently persists, but can probably be reduced by introducing the flow with a tangential component, as suggested for a specimen tube in Figure 17.3. With slab flow assumed, the algebraic start point is that for steady flow in the uniform- temperature matrix: \u0394p = \u22121\u20442\u03c1u2CfLr\u2215rh (17.5) Substituting \u03c1 = p\u2215RT for the ideal gas introduces local Mach number Ma = |u|\u2215 \u221a (\u03b3RT) and allows re-writing: \u0394p\u2215p = \u22121\u20442\u03b3Ma2CfLr\u2215rh (17.5a) Equation 17.5a introduces a further dimension to \u2018jetting\u2019, to which machines using heavy gases \u2013 air and N2 \u2013 are vulnerable: Flow through the individual gauze aperture is subject to the effects of compressibility (choking) at values of approach Mach number, Ma, an order of magnitude lower than the text-book criterion of Ma \u2265 0"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000880_tmag.2007.891399-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000880_tmag.2007.891399-Figure2-1.png",
        "caption": "Fig. 2. Procedure of rotation motion. (a) Initial position; (b) first step; (c) after one period.",
        "texts": [
            " The re-generation method that the nodes between the rotary area and the stationary area are automatically connected has been applied to a motor [5]. However, it is difficult to connect the nodes between both areas because this model has many composition surfaces. In 0018-9464/$25.00 \u00a9 2007 IEEE our method, the simple computation procedure for the rotation motion analysis is proposed without solving the motion equation to save the CPU time. Instead of re-generation of mesh, the material number is changed according to the rotation movement. The outline of rotation motion analysis is shown in Fig. 2. 1) First, the rotation magnet region is equally divided in the initial position as shown in Fig. 2(a). The number of divisions of this region is decided by the rotation speed and time interval. It is assumed that the magnet is rotated in a counterclockwise direction. 2) Next, the material data of each element of rotation magnet is moved to the adjoining element and the magnetization direction is simultaneously rotated as the rotation angle as shown in Fig. 2(b). Here, the 3-D finite element mesh is not changed. 3) The material data and magnetization directions of the rotation magnet are changed in every step for the rotation motion analysis as shown in Fig. 2(c). The rotation speed of magnet is constant in this analysis. The Lorentz force acting on the conductor can be calculated using the magnetic flux density vectors and the eddy current density vectors [6]. The magnetic attractive force between the magnet and the core should be considered in this analysis. Therefore, the total transient torque acting on the stator is calculated by the Maxwell stress tensor. The integral surface is set around the rotation magnet. The equation of forced sinusoidal oscillation with the viscous damping is described as follows [7]: (4) where is the rotation angle, is the moment of inertia, is the coefficient of viscous damping, is the torsion spring constant, is the torque, and is the angular frequency"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002558_1.50800-Figure4-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002558_1.50800-Figure4-1.png",
        "caption": "Fig. 4 Target interception scheme: a) no interception and b) interception.",
        "texts": [
            " These ellipsoids are centered upon the UAV center of mass and mimic the aircraft motion. 1. Object Detection Scheme Sensors onboard the UAV allow the aircraft to characterize its surrounding environment. An object (obstacle or target) is detected by the UAV when the ellipsoidal UAV sensor range intersects the ellipsoidal object [27] (Fig. 3, Table C1 in Appendix C). 2. Target Interception Scheme The UAVintercepts the target when its center of mass is contained within the ellipsoidal target [27] (Fig. 4, Table C1 in Appendix C). Themission is deemed completewhen all UAVs have intercepted the target. 3. Unmanned Aerial Vehicle Communication Scheme Information on objects (obstacles and target) detected by fellow UAVs that are in communication range is also relayed to the UAV. Two UAVs can communicate when the ellipsoids modeling their respective communication range encompass the other UAV center of mass [28] (Fig. 5, Table C1 in Appendix C). A multihopping communication scheme [29] is used when more than two UAVs cooperate"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001369_j.jsv.2008.09.045-Figure5-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001369_j.jsv.2008.09.045-Figure5-1.png",
        "caption": "Fig. 5. Link\u2013pulley contact description.",
        "texts": [
            " 2, the dotted arrows represent the forces (fx and fy), which only arise when a chain link comes in contact with the pulley sheave. Using Newton\u2013Euler formulation [8\u201312] and the Theory of Unilateral Contacts [2,21], the equation of motion for all links under unilateral contact conditions with the pulleys can be written as M\u20acq h \u00f0WN \u00feWSl\u0302SjWT\u00de kN kT ! \u00bc 0, where l\u0302S \u00bc f mi\u00f0 _gi\u00de sign\u00f0 _gi\u00deg. (1) kN, kT are the normal and the sticking constraint forces of the links that are in contact with the pulley and _gi is the relative velocity between a link, i, and the pulley in the sliding plane (refer to Fig. 5). WN represents a matrix with coefficients of relative acceleration (in the normal direction) between the links and the pulleys in the configuration space. WS represents a matrix with coefficients of relative acceleration (in the slip direction) between the links and the pulleys in the configuration space when the links are slipping on the pulley sheave, whereas WT represents a matrix with coefficients of relative acceleration (in the slip direction) between the links and the pulleys in the configuration space when the links are sticking to the pulley sheaves",
            " Haque / Journal of Sound and Vibration 321 (2009) 319\u2013341 325 A chain link contacts a pulley at the ends of a rocker pin. As the plates move, the rocker pins of adjacent links interact with each other. However, assuming negligible dynamic interaction between a pair of rocker pins, the rocker pins are modeled as a single bolt. So, every link is associated with one bolt through which it contacts the pulley sheaves. The bolt dynamics is characterized by a linear massless spring. The surfaces of the bolt are loaded with the normal and frictional contact forces. Fig. 5 illustrates the free body diagram for the interactions between the bolt and a pulley. In this figure, Fr and Ft represent the components of the resultant friction force vector Ff between a chain link and the pulley, which act in the plane of the pulley sheave, and N is the normal force between the link and the pulley. It is necessary to quantify the bolt spring force, Fb, in order to derive the contact forces. The bolt force depends on the bolt length lb and stiffness Kb as well as on the local distance z between the pulley\u2019s surfaces",
            " (6) The coefficient of friction mentioned in case 1 describes the classical Coulomb\u2013Amonton friction law which aptly captures the dynamics associated with kinetic friction and is most commonly referenced in literature. However, the coefficient of friction mentioned in case 2 is more detailed as it not only captures the dynamics associated with kinetic friction, but also captures the dynamics associated with stiction- and Stribeck-effects (which are prominent under dry and lubricated contact conditions, respectively). ARTICLE IN PRESS N. Srivastava, I. Haque / Journal of Sound and Vibration 321 (2009) 319\u2013341 327 Using continuous (or smooth) friction models and the slip angle, g, (refer to Fig. 5) the normal force (N or lN) on the bolt (or pulley) can be obtained as N\u00f0or lN \u00de \u00bc Fr tan b\u00fe Fb cos b , F r \u00bc mN sign\u00f0 _gr\u00de sin g. (7) Substituting the normal force into Eqs. (4) and (5) yields a solvable system of equations for the generalized coordinates, q, of the chain CVT system. Since certain friction characteristics vary continuously with respect to velocity (as in the case of Stribeck friction), friction-driven systems are capable of exhibiting complex and rich nonlinear dynamic behavior such as bifurcations, self-excited vibrations, chaos, etc"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002400_1.4006324-Figure22-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002400_1.4006324-Figure22-1.png",
        "caption": "Fig. 22 Deformation of the right web gears",
        "texts": [],
        "surrounding_texts": [
            "7.1 Gear Deformation and Discussion. To obtain a good understanding of the tooth root stresses of thin-rimmed gears, the deformation of the gear is increased by 2000 times, and the images of the deformed gears are given in Figs. 20\u201322. Figures 20(a)\u201320(d) show images of deformed left web gears with web angles of 0, 15, 30, and 45 deg, respectively. Figures 21(a)\u201321(d) show images of the deformed center web gears with web angles of 0, 15, 30, and 45 deg, respectively. Figures 22(a)\u201322(d) show images of the deformed right web gears with web angles of 0, 15, 30, and 45 deg, respectively. To understand the deformation characteristics of the loaded tooth of the thin-rimmed gear well, an image of the deformed right web gear with a web angle of 0 deg is shown in Fig. 23(a), and an enlarged view of the loaded tooth is also shown in Fig. 23(b). Figure 23 aids in explaining the deformation characteristics of the loaded tooth of the thin-rimmed right web gear with a web angle of 0 deg in the following discussion. Based on Fig. 23, the tooth positions before deformation and after deformation are sketched in Fig. 24(a). From Fig. 24(a), the deformation of the loaded tooth can be roughly divided into two types of deformation: one type is an upward and downward deformation of the loaded tooth as shown in Fig. 24(b), and the other type is a rotation deformation of the loaded tooth as shown in Fig. 24(c). In Fig. 24(b), end A of the loaded tooth has a downward deformation, and end B has an upward deformation. This is because end A is further away from the web than end B, so end A is more flexible than end B. Thus, end A can be deformed more easily than end B. Therefore, when the tooth is very rigid, an inclined deformation (end A is down and end B is up) of the loaded tooth, as shown in Fig. 24(b), occurs. In Fig. 24(c), when the tooth is loaded, end A moves toward the right and end B moves toward the left. It appears that the loaded tooth rotates around the web (the web is the axis). This deformation is called a rotation deformation of the loaded tooth in this paper. The asymmetrical web position (the web center is offset from Journal of Mechanical Design MAY 2012, Vol. 134 / 051001-9 Downloaded From: http://mechanicaldesign.asmedigitalcollection.asme.org/ on 01/28/2016 Terms of Use: http://www.asme.org/about-asme/terms-of-use the tooth center) can be considered to be the main reason for the tooth rotation deformation. As stated above, because end A is further away from the web than end B, the rigidity of end B is greater than that of end A. Therefore, end B shares most of the tooth\u2019s load as shown in Fig. 8(a). When the tooth is very rigid, the greater tooth load on end B will cause the tooth to rotate around the web, as shown in Fig. 24(c). When the rotation deformation of the loaded tooth occurs, end A of the loaded tooth will approach the neighboring tooth, and end B will move far away from the neighboring tooth on the right side, as shown in Fig. 24(a). This allows the tooth root of the loaded tooth to experience a counterintuitive compressive stress on end A and a tensile stress on end B on the side of the loaded tooth surface. When the web is inclined, the supporting rigidity of the web to the teeth will become smaller. In this case, the deformation of the loaded tooth will be affected by the web\u2019s supporting rigidity, and the deformation of the loaded tooth will become more 051001-10 / Vol. 134, MAY 2012 Transactions of the ASME Downloaded From: http://mechanicaldesign.asmedigitalcollection.asme.org/ on 01/28/2016 Terms of Use: http://www.asme.org/about-asme/terms-of-use complicated than the statements mentioned above. It is necessary to consider the web deformation on the loaded tooth to understand the root stress of thin-rimmed gears with an inclined web. 7.2 Deformation-Sharing Ratios of the Tooth, the Rim, and the Web. A tooth\u2019s relative deformation of a pair of contact gears along the line of action can be calculated by LTCA under the application of a torque load. Table 4 shows the results. In Table 4, the tooth\u2019s relative deformation of a pair of solid gears (the mating gears in Fig. 2(d)) is calculated at the worst load position, and the result is shown in the first column (denoted by \u201csolid gears\u201d). The relative tooth deformations of thin-rimmed left web gears with different web angles are also calculated when the left web gears are engaged with the solid mating gear. The calculation results are shown in the columns denoted by \u201cThin-rimmed left web gear\u201d in Table 4. In Table 4, the deformations are divided into the gear deformation (the total deformation of a pair of gears), the tooth deformation (the deformation resulting only from the contact teeth), and the rim and web deformation (the deformation resulting only from the rim and the web). Therefore, the deformation of the gear is equal to the deformation of the teeth plus the deformation of the rim and the web. In Table 4, the tooth\u2019s relative deformation of the pair of solid gears is 3.8 lm. Because it can be considered that the solid gears have no rim or web deformations, 3.8 lm represents the deformation resulting from the contact teeth. Therefore, in Table 4, the deformation of the teeth is also 3.8 lm. This deformation can be regarded as the tooth deformation of all of the thin-rimmed gears used in this paper when they are engaged with the solid mating gear. The tooth relative deformations of the thin-rimmed left inclined web gears with different web angles are also calculated when they are engaged with the solid mating gear at the worst load positions of the tooth contact. The calculation results are given in Table 4. The deformations of the rim and the web of the thin-rimmed left inclined web gears are obtained by calculating the value of the gear deformations minus the tooth deformation. The deformation-sharing ratio of the teeth can be obtained by dividing the tooth deformation by the gear deformation. The deformation-sharing ratio of the rim and the web can be calculated by dividing the rim and the web deformation by the gear deformation. The calculation results are shown in Table 4. From Table 4, it can be seen that the teeth share only 29.6% of the total deformation of the gears while the rim and the web share 70.4% of the total deformation for the thin-rimmed straight web gear. The deformation-sharing ratio of the teeth decreases while the deformation-sharing ratio of the rim and the web increases when the web angle is increased. When the web angle is increased to 60 deg, the rim and the web share 93.4% of the total deformation of the gears while the teeth only share 6.6% of the total deformation. These findings indicate that the tooth deformation can be neglected in the engineering calculations for thin-rimmed left web gears with large web angles. The deformation-sharing ratios of the center web and right web gears are calculated in the same way, and the calculation results are given in Tables 5 and 6, respectively. From Tables 5 and 6, we find that the same conclusions can be obtained."
        ]
    },
    {
        "image_filename": "designv11_12_0001517_j.mechmachtheory.2009.12.004-Figure4-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001517_j.mechmachtheory.2009.12.004-Figure4-1.png",
        "caption": "Fig. 4. A six-bar linkage with three sliders.",
        "texts": [
            " Since arctan(x5) 2 ( p/2, p/2], each real solution of x5 corresponds to a unique h5 if h5 2 ( p, p]. 4R1P Loop: The Stephenson type linkage in Fig. 3 contains a 4R1P loop. The loop equation for the 4R1P chain ABEFGA in Fig. 3 can be expressed as a2eih2 \u00fe a4ei\u00f0h3\u00feb\u00de a1 a7 a5eih5 s2eia1 \u00bc 0: \u00f021\u00de Eliminating s2 and using the tangent half-angle formula of Eq. (14), Eq. (21) can be written in the quadratic form as Eq. (15) and from which equations in the form of Eqs. (17) and (20) can be derived. 3R2P Loop: The Stephenson type linkage in Fig. 4 contains a 3R2P loop. The loop equation for the 3R2P chain ABEFGA in Fig. 4 can be expressed as a2eih2 \u00fe a4ei\u00f0h3\u00feb\u00de a1 a7 s3ei\u00f0a1\u00fec\u00de s2eia1 \u00bc 0: \u00f022\u00de Eliminating s2, Eq. (22) can be written as s3 sin c a2 sin\u00f0h2 a1\u00de a7 sin a1 a4 sin\u00f0h3 \u00fe b a1\u00de \u00bc 0: \u00f023\u00de The discriminant of Eq. (23) with s3 as the unknown parameter is D3 \u00bc sin2 c: \u00f024\u00de Since sin2 c P 0 is always true, Eq. (24) can be automatically satisfied. If sin c \u2013 0, for each pair of h2 and h3 values, there is only one solution for s3 in Eq. (23), i.e. only one linkage configuration. Thus, the branch of the four-bar chain is also the branch of the whole linkage in Fig. 4. If sin c = 0, there are infinite solutions to s3 and such linkage is useless. Thus, sin c cannot be allowed to be zero. Virtual five-bar loop: A Watt six-bar chain consists of two four-bar loops and a six-bar loop ABCEFGA (Fig. 2). The six-bar loop may be also transformed to a five-bar loop through the stretch rotation of a four-bar loop [11]. Thus, a Watt six-bar linkage may be regarded as a degenerate Stephenson six-bar linkage and the mobility analysis method for Stephenson six-bar linkages is fully applicable",
            " If the linkage configurations for a given branch of the four-bar chain satisfy Eq. (17), the given branch of the four-bar chain can be two branches of single-DOF double-loop linkages. Otherwise, the given branch is not valid and the single-DOF double-loop linkage cannot be assembled. Each solution set of Eq. (20) represents a branch of single-DOF double-loop linkages. (4) Repeat the above three steps if the four-bar chain has another branch. (5) It is noted that the single-DOF double-loop linkages with three prismatic joints in Fig. 4 has always Type I branches, which are determined only by the four-bar loop. Each branch of the four-bar loop represents one branch for the linkage. For Type II branch of single-DOF double-loop linkages, branch points exist and the branches of single-DOF double-loop linkages are caused by interaction between these two loops. The mobility of the four-bar loop is blocked by the other loop (Fig. 9). The coupling between the four-bar loop and the other loop will affect the rotatability of the four-bar loop",
            "6 ] in Eqs. (9a) and (20b). Branch 3-4: This branch contains segment 3-4. Sub-branch 3-4: Segment 3-4 with h2 2 [ 105.2 , 15.6 ] in Eqs. (9a) and (20a); Sub-branch 4-3: Segment 4-3 with h2 2 [ 105.2 , 15.6 ] in Eqs. (9a) and (20b). Branch 6-5: This branch contains segment 6-5. Sub-branch 6-5: Segment 6-5 with h2 2 [7.9 , 72.6 ] in Eqs. (9b) and (20a); Sub-branch 5-6: Segment 5-6 with h2 2 [7.9 , 72.6 ] in Eqs. (9b) and (20b). Example 4. Given the same dimensions for the single-DOF double-loop linkage in Fig. 4 as in the above example and sin c \u2013 0. It may be noted that with the same dimensions for the single-DOF double-loop linkage in Fig. 4 as in Fig. 3, the branches of the linkage are determined by the 3R1P chain and each branch of the 3R1P chain represents one branch of the whole linkage since Eq. (24) is always satisfied and there is only one solution to s3 for a given pair of h2 and h3 and sin c \u2013 0. Thus, there is only one branch for the linkage in Fig. 4 in this example. This paper presents the first successful attempt that extends the discriminant method to the mobility identification of any type of single degree-of-freedom (DOF) double-loop planar linkages under any input and output condition. The approach is the algebraic counterpart of the joint rotation space method for the mobility of Stephenson six-bar linkages [10]. The method has the following desirable features. 1. The applicability of the method is only affected by the essential polynomial equation form of the linkage rather than the type of joints or even the physical appearance of the linkage"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002215_1.3204650-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002215_1.3204650-Figure1-1.png",
        "caption": "Fig. 1 S",
        "texts": [
            "3204650 Keywords: uncertainty, uncertain rotor, uncertain bearings, rotordynamics, nonparametric stochastic modeling, random matrices Introduction A series of algorithms were described in the first part of this eries 1 to simulate an ensemble of uncertain stiffness, mass, yroscopic, and centripetal effect matrices. In this second part, hese algorithms are applied to a representative rotor to provide a rst perspective on the effects of uncertainty in rotordynamic sysems. To avoid needless duplication, all equation and figure numbers arked by an asterisk in this part will refer to the equations and gures with the same number of Part I. Examples of Application The symmetric rotor system with asymmetric bearings decribed in Ref. 2 see Fig. 1 and Ref. 2 for properties was first onsidered to exemplify the concepts introduced in Part I. The otor was discretized by finite elements using 25 key points with our degrees of freedom each for a total of 100 degrees of freeom , 24 beam elements, five disk elements, and two speed deendent bearing elements. The Campbell diagram of this rotor is hown in Fig. 2. The next task focused on the development of a reduced order odel of the 100 degrees of freedom finite element model. This educed order model was achieved using the mode shapes of the ree-free rotor at rest, and a convergence study see Fig",
            " 6\u20138 are typical probability density unctions of the normalized deviations of the magnitude of the igenvalue denoted here by , i.e., \u2212 \u0304 / \u0304, where \u0304 denotes he eigenvalue magnitude of the mean model, for 800 rpm, 593 rpm, 2400 rpm, 3200 rpm, and 3614 rpm. These speeds were elected on the following basis. At the speed of 800 rpm, the earing mode frequencies are well separated from those of the orward/backward modes, which themselves have started to split rom one another. The speed of 1593 rpm corresponds, as 00 rpm, to well separated frequencies but is also the critical ig. 2 Campbell diagram of the rotor of Fig. 1. \u201ea\u2026 Imaginary art of eigenvalues. \u201eb\u2026 Real part of eigenvalues. 92502-2 / Vol. 132, SEPTEMBER 2010 om: http://gasturbinespower.asmedigitalcollection.asme.org/ on 01/27/201 speed of the rotor. Further, at the speed of 2400 rpm, the bearing mode and the forward/backward modes exhibit equal imaginary parts of the eigenvalues not of the real parts . Finally, the two higher speeds of 3200 rpm and 3614 rpm were considered in the original investigation 2 . The determination of the probability density functions requires first the assignment of the eigenvalues to particular sets; e",
            "13 CB,xz, CB,zx 0.17\u20130.33 Bearing 2 B,xx 0.16\u20130.17 KB,zz 0.15\u20130.16 KB,zx 0.15\u20130.16 B,xx 0.14 CB,zz 0.16\u20130.19 CB,xz, CB,zx 0.14 ig. 20 Evolution versus rotor speed of the mean model vales of KB,xz and of the mean and mean \u00b11 standard deviation of he simulated values of this coefficient. \u201ea\u2026 Bearing 1. \u201eb\u2026 Bearng 2. ournal of Engineering for Gas Turbines and Power om: http://gasturbinespower.asmedigitalcollection.asme.org/ on 01/27/201 The above discussion was carried out on the symmetric rotor model of Fig. 1, and uncertainty was restricted to maintain the symmetry of the rotor. In practical situations, it would be expected that uncertainties in the rotor geometry, material properties, etc., would induce a small level of asymmetry. In this light, it was desired to make a preliminary assessment of the differences in the effects induced by symmetric and asymmetric uncertainties. This Fig. 22 Probability density function of the normalized eigenvalue magnitude for different rotor speeds, first bearing mode, and uncertainty in bearing properties. SEPTEMBER 2010, Vol. 132 / 092502-7 6 Terms of Use: http://www.asme.org/about-asme/terms-of-use a s d 1 a p n F t c F v m F t c 0 Downloaded Fr nalysis was carried out on the rotor of Fig. 1 but with bearings uch that KB,xx=KB,zz and KB,xz=KB,zx=0 and similarly for the amping coefficients see Fig. 26 . Then, the coefficients of Eq. 53 are constant, thereby simplifying the analysis. Shown in Fig. 27 are the Campbell diagrams of the rotor of Fig. on the bearings of Fig. 26 both in the fixed frame Fig. 27 a nd in the rotating one Fig. 27 b . The derivation of an accurate reduced order model was accomlished next using, as in the previous discussion, the modes of the onrotating rotor free of its bearings",
            " 30 Probability density function of the absolute deviation n eigenvalue imaginary part at 2200 rpm for mode I, asymmetic and symmetric uncertainties in stiffness ig. 31 Probability density function of the absolute deviation n eigenvalue imaginary part at 2590 rpm for mode I, asymmetic and symmetric uncertainties in stiffness 92502-10 / Vol. 132, SEPTEMBER 2010 om: http://gasturbinespower.asmedigitalcollection.asme.org/ on 01/27/201 The methodology described in Part I for the consideration of uncertainty on the response of rotordynamic systems was exemplified here on the simple but representative system of Fig. 1. The effects of uncertainty on its eigenvalues, eigenvectors, unbalanced forced response, and instability threshold were all studied under uncertainty from either the rotor stiffness or mass properties or the bearing properties. Further, the analysis was performed in both conditions of uncertainty maintaining or violating the symmetry of the rotor. These results provide a first perspective on the effects Transactions of the ASME 6 Terms of Use: http://www.asme.org/about-asme/terms-of-use o e m R J Downloaded Fr f uncertainty in rotordynamic systems but also demonstrate the ase with which such analyses can be performed on complex rotor odels using the nonparametric methodology of Part I"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003878_1.4031025-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003878_1.4031025-Figure1-1.png",
        "caption": "Fig. 1 Spherical involute surface definitions",
        "texts": [
            " (14) tan g \u00bc \u00f0tan\u00bd\u00f0h\u00fe u\u00desin b \u00de=\u00f0sin u cos b\u00de (15) Hence, for any given angle u, angle h is directly found from Eq. (13), since b is known. Angles h and u are replaced in Eq. (15) to find g. Another variation of this procedure with forward approach is to find a unique u value for any given h through solution of nonlinear Eq. (13) and replacing u and h again in Eq. (15). The latter approach is the forward approach; however, it is mathematically easier to use u as given in Eq. (13) and solve for roll angle h. Referring to Fig. 1, coordinates of any point P that is generated at roll angle h from base cone K at cone distance r and base cone angle 2b are (where T represents matrix transpose) P \u00bc rsin bsin h rsin b cos h r cos b\u00bd T (16) Note that Fig. 1 is established only for one cross section of r \u00bc q, where r should vary across face width of the tooth to generate entire bevel gear surface. To calculate normal n to the spherical involute surface at point P; coordinate systems x0y0z0 and x00y00z00 with origins at the same point as coordinate system xyz are established. x0y0z0 coordinate system is established by rotation of the coordinate system xyz around its z axis as much as h, such that arc OP \u00f0 coincides with plane y0 z0. Coordinate system x00y00z00 is established by rotating the coordinate system x0y0z0 around x0 axis as much as \u00feg, such that line CP coincides with z00 axis. With this in spherical triangle 4OPB if tangent to arc OP \u00f0 (shown as n0 in Fig. 1(a)) at point P is rotated around z00 axis for angle \u00fe#, it results in normal to the curve C (shown as n in Fig. 1(a)). In coordinate system x00y00z00, unit tangent vector to arc OP \u00f0 at point P is n0 \u00bc 0 1 0\u00bd T (superscript T means transpose); therefore, n in x00y00z00 is nx00y00z00 \u00bc sin# cos# 0\u00bd T (17) Hence, unit normal vector n in xyz coordinate system is nxyz \u00bc cos h sin h 0 sin h cos h 0 0 0 1 2 4 3 5 1 0 0 0 cos g sin g 0 sin g cos g 2 4 3 5 sin# cos# 0 2 4 3 5 (18) or, after simplification nxyz \u00bc cos hsin# sin h cos g cos# sin hsin# cos h cos g cos# sin g cos# 2 64 3 75 (19) From Eqs. (16) and (19), coordinates and unit normal vectors of any point P on spherical involute surface are calculated based on two independent surface parameters of r and u"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000540_amc.2008.4516037-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000540_amc.2008.4516037-Figure1-1.png",
        "caption": "Fig. 1. Two-wheel inverted mobile manipulator",
        "texts": [
            " \u2022 Nullspace which is made up of active joint is constructed by separating passive joint from Kinematics \u2022 Attitude stabilization is performed by internal motion by using Nullspace This paper is composed of 4 parts. Modeling of two-wheel inverted mobile manipulator is shown in section II. Section III describes the control law to realize end-effector force control abd pushing operation of two-wheel inverted mobile manipulator. Then, simulation is performed to confirm the validity of proposed method in SectionIV. Finally, conclusion and future works are summarized in section V. II. MODELING A. Two-Wheel Inverted Mobile Manipulator Model of the mobile manipulator is shown in Fig.1. This robot has the mechanism that two wheels whose axis sustains the upper structure. Upper structure has four links manipulator and three joints. Note that the joint q1 is passive joint. 978-1-4244-1703-2/08/$25.00 \u00a92008 IEEE 33 1) Kinematics: A position of vehicle part xv , an endeffector position of mobile manipulator xe, and the joint angle of manipulator part qm are expressed by these equation. xv = (xv, zv)T \u2208 R2 (1) xe = (xe, ze)T \u2208 R2 (2) qm = (q2, \u00b7 \u00b7 \u00b7 , q4) T \u2208 R4 (3) Angles of the wheels qv , inclination angle of platform q1, angle of manipulator joint qm are dealt as a whole by using a jointspace vector of a mobile manipulator q defined in eq"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003581_aim.2011.6027001-Figure6-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003581_aim.2011.6027001-Figure6-1.png",
        "caption": "Fig. 6. Conceptional drawing of the completed robot",
        "texts": [
            "07 MPa, the length of one unit is about 28 mm and the contraction rate of one unit is 20.3 %, and diameter is about 28 mm and expansion rate is 97.7 %. We confirm that the robot can vary the shape of the segments between short/thick and long/thin. This change is a characteristic of peristaltic crawling. A. Proposed Robot Mechanism For treating the affliction of the small intestine, the robot needs to arrange the electric wire of the camera and the treatment tool that is attached to conventional endoscope within its body. Figure 6 shows a schematic illustration of the completed robot. When this robot is detached from the endoscope, the robot observes the small intestine and spontaneously, the endoscope observes the large intestine. The robot should have a space at the center for arranging the treatment tool and electrical wires. These components must not hinder the robot\u2019s motion. We use a \u201cTube-Slide Mechanism\u201d that was developed by our laboratory [15]. Figure 7 shows a sectional view of this mechanism. Each unit has a space called a chamber between the bellows and the artificial muscle"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002273_1.3680609-Figure5-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002273_1.3680609-Figure5-1.png",
        "caption": "Fig. 5. The forces acting in the xz-plane on a ball with backspin include the gravitational force mg, the drag force FD, the lift force FL, and the buoyant force FB.",
        "texts": [
            " Experiment 3 was performed in essentially the same manner as experiment 2 but using the polystyrene ball with a baseball seam. The ball was launched by hand with backspin, varying the orientation of the seam on a trial and error basis in order to maximize the sideways deflection. Experiment 3 was performed after Professor Alan Nathan sent the author a video clip showing a baseball deflecting sideways in the opposite direction to that expected from the Magnus force.20 Consider a ball of mass m traveling with backspin in the vertical xz-plane at speed v and angle h with the horizontal, as shown in Fig. 5. The main forces on the ball consist of the gravitational force mg, a drag force FD acting in a direction opposite the velocity, and a lift force FL acting in a direction perpendicular to the velocity and the spin axis. For relatively light or large balls, the vertical buoyant force FB \u00bc mAg may also be significant; mA being the mass of air displaced by the ball. Because of buoyancy, m cannot be measured directly on a scale because the scale reading is m mA. The mass m was therefore determined by adding mA to the scale reading"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003713_978-1-4614-2419-2-Figure15.1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003713_978-1-4614-2419-2-Figure15.1-1.png",
        "caption": "Fig. 15.1 The inner surface of the wheel hub is restricted in x, y and z directions, the contact area shifts one segment in each time-step",
        "texts": [
            " Johansson \u2022 F. Larsson Department of Applied Mechanics, Chalmers University of Technology, SE-412 96 G\u20acoteborg, Sweden e-mail: ronasi@chalmers.se R. Allemang et al. (eds.), Topics in Modal Analysis II, Volume 6, Conference Proceedings of the Society for Experimental Mechanics Series 31, DOI 10.1007/978-1-4614-2419-2_15, # The Society for Experimental Mechanics, Inc. 2012 169 A wheel from train type \u2018Regina\u2019 is analyzed under dynamic condition using the Finite Element software Abaqus-standard, cf. Fig. 15.1. The material properties is chosen close to steel properties, E \u00bc 210 GPa, n \u00bc 0. 29, density of 7,800 kg/m3. The internal hub surface of the wheel is restrained in x, y and z directions. The inner-side and outer-side strain gauges are positioned at the radius of 208 and 200 mm from the wheel center respectively, i.e. the radial strains are extracted at these radii. Compensation for centrifugal and gyroscopic forces due to rotation and translation of the wheel will be considered later. Acting surface of the contact pressure rotates around the wheel rim in analysis-steps. The wheel is partitioned in 72 segments, Fig. 15.1, that creates 72 surfaces on the wheel rim on which the contact pressure can be applied. The analysis is performed for a one full cycle in 72 analysis-steps (Dynamic-Implicit) of 0.0015 s, i.e. one full rotation in 0.108 s which corresponds to axel linear speed of about 200 km/h. Eight-node linear brick elements with reduced integration and hourglass control are used for the analysis. The contact force between the wheel and the rail typically amounts to a pressure over a ellipsoid surface of 5\u201315 mm in diameter",
            " In the present investigation, our interest is to determine the force magnitude, assuming that the location of the force is known. Therefore, we allow ourself to let the small contact surface be represented by a uniform contact pressure area of the same size as the element surface. Since the force is traveling over several elements, it is important to have a smooth transition when the force switches elements. Therefore, in each analysis-step, the contact pressure is applied on two neighboring contact surfaces, Fig. 15.1. While the amplitude of the pressure on the first surface-area a1 is linearly decreasing from 1 to 0, the amplitude of the second surfacearea a2 increases linearly from 0 to 1 over the time-steps in the analysis-step. In the next analysis-steps the two contact surfaces shift by one segment which means the surface with increasing amplitude in the previous analysis-step has a decreasing amplitude in the new analysis-step. This way the contact surface switches smoothly around the wheel rim over analysis-steps representing a rotating wheel"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002444_c1an15346c-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002444_c1an15346c-Figure2-1.png",
        "caption": "Fig. 2 Cyclic voltammograms of an ACMGCE in a 0.1 M phosphate buffer solution (pH 7.0) at scan rate 20 mV s 1 in (a) absence, (b) and (c) presence of 0.2 and 2.5 mM AA and GSH respectively.",
        "texts": [
            " In these conditions, the kinetic parameters of the surface electron transfer rate constant (ks), and the charge transfer coefficient (a) for electron transfer between the electrode surface and the immobilized AC film can be estimated based on the plot slopes in Fig. S1D\u2020 and DEp according to the Laviron theory.66 Using the data in Fig. S1D\u2020, the values of ks and a were found to be 6.94 s 1 and 0.52 respectively at pH 7.0. In order to test the electrocatalytic activity of the ACMGCE, as a bifunctional electrocatalyst, for the oxidation of AA and GSH, This journal is \u00aa The Royal Society of Chemistry 2011 cyclic voltammograms of the ACMGCE were obtained in the absence and presence of AA or GSH. Fig. 2 shows the cyclic voltammograms of the ACMGCE in the absence (voltammogram a), and the presence of 0.2 mMAA (voltammogram b), and 2.5 mM GSH (voltammogram c). A comparison of voltammograms of (a) and (b) as well as (a) and (c) indicates that redox couple I of the electrodeposited AC has an electrocatalytic effect on the oxidation of AA. Also, redox couple II of the modified electrode shows a similar behavior for the oxidation of GSH. In other words, the ACMGCE plays the role of a bifunctional electrocatalyst for the oxidation of AA and GSH in two oxidation peak potentials which are well separated from each other"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000017_1.14796-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000017_1.14796-Figure2-1.png",
        "caption": "Fig. 2 Front and side views of spheroid-cone airship \u201cWhite Diamond.\u201d",
        "texts": [
            " With the preceding background in mind, in 2002 the author instigated the development of a new airship for an on-going project concerned with scientific exploration of rainforest canopy.28,29 In essence the airship had to lift two people (including the pilot) and be capable of stable yet maneuverable flight in quiescent (still air) tropical conditions, effectively replacing an older design tested by the author.29,30 Because the required maximum flight speed target was unusually low, just 3 m/s, it followed that an airship with low \u03bb, or even \u03bb = 1, might provide the best design solution. Indeed, it was finally decided to adopt a spheroid-cone hull form with \u03bb = 1.54 (Fig. 2). To reduce risk, this major design decision was preceded by a brief evaluation study using several wind-tunnel models. Two closed-circuit wind tunnels at the author\u2019s present address were used for testing: tunnel 1 with a contraction of 7.2:1 and a quasi-octagonal working section 1.24 by 1.0 m and tunnel 2 with a contraction of 5.6:1 and a rectangular working section 1.0 by 0.77 m. The turbulence levels of both tunnels were estimated (using hot-wire techniques) to be less than 0.4% at the test flow speeds involved, about 35\u201343 m/s",
            " At the end of the evaluation study it was, therefore, sensible to assume that, provided similarly favorable flow effects exist at the much higher Reynolds number expected for the full-scale airship, an overall performance gain would be possible by the attachment of a rear-facing cone to a spherical balloon, reducing CDVtot and, hence, reducing the mass (and cost) of a propulsion system. Following the wind-tunnel evaluation study, a decision was made to commit to the full-size airship design shown in Fig. 2. This spheroid-cone airship was tested using tethers inside shed 1 at Cardington, United Kingdom (in January 2004), and then flown freely over the Kaieteur National Park, Guyana (in July 2004). The main balloon had a diameter dmax = 9.7 \u00b1 0.05 m and, hence, was capable of containing VHe = 480 \u00b1 5 m3 of helium when filled at its maximum allowable internal pressure. The tail cone was made of a light fabric and had a length of about 6 m with a semi-apex angle of 0.52 rad. The cone was held in tension by an internal alloy tube running along the central longitudinal axis, and tightly attached to the balloon in 24 places on a pitch circle diameter of 7 m",
            "71-m-diam propeller. Flight speed was measured using a three-axis Metek USA-1 Basic sonic anemometer with a data acquisition rate of 10 Hz. In tunnel 1, this anemometer produced a fairly uniform white noise error at 10 Hz, with a root-mean-square velocity fluctuation of about urms = 0.05 m/s, at flow speeds between 1 and 3 m/s. The anemometer was mounted on a sting in front of the gondola at a radial distance of about r = 6.6 m from the balloon\u2019s center at about \u03b8 = 1.26 rad from the major axis (Fig. 2). A (fixed) flow speed correction factor D ow nl oa de d by G E O R G E W A SH IN G T O N U N IV o n Ja nu ar y 27 , 2 01 5 | h ttp :// ar c. ai aa .o rg | D O I: 1 0. 25 14 /1 .1 47 96 of 0.88 was calculated to estimate the forward flight speed, U = U\u221e, from the measured (higher) value, U\u03b8 , at the sting position using the potential flow equation for flow around a sphere31 U\u03b8 = \u2212 1 2 U\u221e sin \u03b8 [ 2 + ( 1 2 dmax/r )3] (13) No account was taken for effective variations in anemometer position \u03b8 brought about by yaw and pitch changes",
            " When the estimation methods presented by Hoerner1 are used, these parasitic terms probably only accounted for about 10\u201320% of the total drag. However, because of interference with the hull, their combined effect was probably considerably larger. Some experimental evidence9 suggests that interference effects of any hull protuberances will be more pronounced on hull forms with low \u03bb. In particular, note that at the base of the balloon there was a cylindrical fabric protrusion (about 0.6 m in diameter, 0.4 m depth) caused by one of the helium relief valves (Fig. 2). Based on previous experiments,9 it is likely that this valve unit alone caused a significant drag rise. Another possibility is that the rigging lines tended to produce scalloping in the lower balloon envelope and that may have also caused flow separation upstream of the maximum diameter. Flutter in the fabric tail cone and fin might have also brought about adverse flow effects. (Both surfaces were stretched taut, but flutter was observed while the airship was being held on the ground during sporadic breezes exceeding 2 m/s"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000568_00368790710746110-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000568_00368790710746110-Figure1-1.png",
        "caption": "Figure 1 Geometry and coordinates of the problem",
        "texts": [
            "htm Industrial Lubrication and Tribology 59/3 (2007) 143\u2013147 q Emerald Group Publishing Limited [ISSN 0036-8792] [DOI 10.1108/00368790710746110] with magnetic fluid-based squeeze film between curved circular plates. Recently, Patel and Deheri (2002a, b) analyzed magnetic fluid-based squeeze film between two curved plates lying along the surfaces determined by a secant and hyperbolic functions. Here, it has been proposed to analyze the configuration of Prakash and Vij (1973) in the presence of a magnetic fluid with regards to conical plates. The configuration of the bearing is shown in Figure 1. In the analysis the assumptions of usual hydrodynamical lubrication theory are retained. Also, it is taken into account that the porous matrix is homogeneous and isotropic. Developing the analysis of Prakash and Vij (1973) the concerned Reynolds\u2019 equation governing the film pressure is obtained as: 1 x d dx x d dx \u00f0 p2 0:5m0 mH2\u00de \u00bc 12m_h sinv h3sin3v\u00fe 12fH \u00f01\u00de where: H2 \u00bc a2 12 x2 sin2v a2 ; m is the fluid viscosity, m represents the magnetic susceptibility. m0 stands for the permeability of the free space, while the inclination of the external magnetic field with the lower plate is taken as in Bhat and Deheri (1991)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001866_bf00251592-Figure7-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001866_bf00251592-Figure7-1.png",
        "caption": "Fig. 7. Stephenson-! medm~sm",
        "texts": [
            "ix-Bar Motion II. The Stephenson-1 and Stephenson-2 Mechanisms E. J . F . PRIMROSE, F. FREUDENSTEIN & B. ROTH Communicated by E. LEIMANIS I. Abstract Algebraic, geometric and kinematic properties are derived for the curves generated by points on the floating links of the six-link mechanisms derived from the Stephenson kinematic chain. II. The Stephenson-1 Mechanism 1. The Mechanism (Fig. 7) This mechanism - the type studied by MUELL~R [9]* -- is derived from the Stephenson chain by fixing a binary link, which is connected to two ternary links. The notation in Fig.7 corresponds largely to that of MtrELLER, in order to facilitate comparisons. It is possible also, of course, to have Stephenson-1 mechanisms without crossing links; in such cases angles a and a' are of opposite sign. Point G describes the six-bar curve, which has also been called the \"kneecurve\" [9]. A is the origin and A E is along the positive x-axis of our right-handed Cartesian coordinate system. * References refer to those given at the end of Part I of this work, this Archive, page 41. Six-Bar Motion. II 2. The Stephenson-1 Six-Bar Curve From Fig. 7, we have [ X - s e t C~,-~)] [ Y - s e - ' (~'-')] = l 2, and if we set e ~ = t , this becomes . 4 t 2 - p t + A = O where A = X s e i~, P = X y+s2-12 and the bar denotes complex conjugation. Similarly, we have E X _ m _ s , e~(~'-~')]EY_m_s,e-i(*'-~')]=1,2, and if we set e i~''= u, we obtain Jl' u 2 - p ' u + A ' = O where A ' = ( X - m ) s ' e ~', p ' = ( X - m ) ( Y - m ) + s ' 2 - 1 '2. From the quadrilateral A BDE, we have also [m+ r ' e ' q \" - r e 'q'] [m + r' e - ' q \" - r e - '~] =n 2 , which yields (r r ' - m r' t) U2 + (rrt r - Q t + m r t2) u + ( - m r' t + r r' t2)-----0 where Q = m 2 + r' 2 + r E _ n 2"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000224_1-4020-2933-0_13-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000224_1-4020-2933-0_13-Figure1-1.png",
        "caption": "Figure 1. A rigid body positioned in an inertial frame",
        "texts": [
            " This formulation is simple to learn and easy to implement in either special- or general-purpose computer programs, however, the computational efficiency is not the best. Several well known commercial multibody programs such as ADAMS and DADS employ this type of formulation. This formulation yields a large set of loosely coupled equations of motion. The fundamental assumption in this formulation is that a body is assigned a set of translational and rotational coordinates. Therefore, as shown in Figure 1, a typical rigid body i is positioned in an inertial (non-moving, global) x\u2013y\u2013z frame by vector ir . This vector locates the origin of a bodyattached (moving, local) iii \u03b6\u03b7\u03be \u2212\u2212 frame that is at the mass center of the body. This reference frame may or may not coincide with the body principal axes. The x\u2013y\u2013z components of vector ir represent the translational coordinates of the body and they are denoted by the algebraic vector { }T i i x y z=r . Angular orientation of iii \u03b6\u03b7\u03be \u2212\u2212 frame with respect to x\u2013y\u2013z frame can be described by a set of Euler angles or Euler parameters [1]",
            " The translational and rotational velocity vectors for the body are defined respectively as { }T i i x y z=r and { }( ) ( ) ( ) T i x y z i \u03c9 \u03c9 \u03c9= or { }( ) ( ) ( ) T i i\u03be \u03b7 \u03b6\u03c9 \u03c9 \u03c9= , where \u03c9i = Ai\u03c9i. Similarly, the acceleration vectors are defined as { }T i i x y z=r and { }( ) ( ) ( ) T i x y z i \u03c9 \u03c9 \u03c9= or { }( ) ( ) ( ) T i i\u03be \u03b7 \u03b6\u03c9 \u03c9 \u03c9= In order to simplify the discussion, we adopt the x\u2013y\u2013 z components of angular velocity and acceleration vectors in our formulations. A vector or a point can be defined attached to a rigid body. As shown in Figure 1, vector is is defined on a body representing a particular axis. However, vector P is locates point P from the origin of the body. Local components of these vectors are constants and are respectively denoted as { }( ) ( ) ( ) T i i s s s\u03be \u03b7 \u03b6=s and { }T i PPPP i \u03b6\u03b7\u03be=s . Obviously, the following transformations can be performed in order to compute the global components of these vectors: i i i =s A s (or P P i i i =s A s ). (1) The velocity and acceleration of a body-fixed vector, either i s or P i s , are computed as: i i i i i i i i i i i = = \u2212 = \u2212 + s s s s s s "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001847_1.4000269-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001847_1.4000269-Figure1-1.png",
        "caption": "Fig. 1 Geometry of test bearing",
        "texts": [
            " Editor: Michael D. Bryant. ournal of Tribology Copyright \u00a9 20 om: http://tribology.asmedigitalcollection.asme.org/ on 09/01/2017 Terms control for the stabilization method by using starved lubrication conditions is devised and stabilization verification experiment is performed by applying the above experimental method to a highspeed bearing test rig. The results obtained by the verification of the efficiency of the method are discussed in this paper. 2.1 Experimental Test Rig and Experimental Method. Figure 1 is a schematic diagram of a cylindrical journal bearing used in the experiments, and its main dimensions are given in Table 1. The upper part of a test bearing is provided with an oil supply groove, which allows supplying lubricating oil from the groove into a bearing clearance. The bearing clearance is set relatively large because the research subject of our experiment is concentrated to small-bore high-speed bearing used in small size lightly loaded, relatively low cost machinery such as small-size compressor or turbine"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001392_6.2007-6671-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001392_6.2007-6671-Figure1-1.png",
        "caption": "Figure 1. Key Missile Quantities",
        "texts": [
            " W2 weights the importance of shaping the transient response, or the frequency response traits of control set. Wv weights the importance of matching actual control set e\u00a4ect to the desired control e\u00a4ect, or virtual control. In the next several sections all of the nonlinear and linearized equations will be developed to de ne the problem. The following sections will then outline the speci c problems examined here, along with results. Finally, some conclusions will be drawn, with suggestions for future work. II. The Nonlinear Equations Figure 1 shows a missile with some of the key variables and axes identi ed. The relevant nonlinear equations of motion are given by _x = f (x) + g (v(x; u)) y = h(x) + k(v(x; u)) where v are forces and moments and u are control de ections. More speci cally, these equations are 2 of 32 American Institute of Aeronautics and Astronautics _Vm = 1 V [AX +AY tan +AZ tan ] _ = P cos2 tan +Q R sin cos tan AX V sin cos Vm +AZ V cos2 Vm _ = P cos2 tan +Q sin cos tan R AX V sin cos Vm +AY V cos2 Vm _P = 1 (IxxIzz I2xz) \" [Ixz (Ixx + Izz Iyy)]PQ+ Izz (Iyy Izz) I2xz QR +IzzL+ IxzN + IzzMTX + IxzMTZ # _Q = 1 Iyy Ixz R2 P 2 + (Izz Ixx)PR+M+MTY _R = 1 (IxxIzz I2xz) \" Ixx (Ixx Iyy) + I2xz PQ+ [Ixz (Iyy Ixx Izz)]QR +IxzL+ IxxN + IxzMTX + IxxMTZ # where V q 1 + tan2 + tan2 and AX = FX m + TX m AY = FY m + TY m AZ = FZ m + TZ m where FX , FY , and FZ are the aerodynamic forces; TX , TY , and TZ are the propulsive forces; L,M, and N are the roll, pitch and yaw aerodynamic moments; and MTX , MTy , and MTz are the propulsive moments"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003761_9781118818428-FigureA.4-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003761_9781118818428-FigureA.4-1.png",
        "caption": "Figure A.4.1 Notation (to no scale!) for analysis of a quasi-linear \u2018lever-crank\u2019 drive mechanism offering thermodynamic flexibility combined with minimal side thrust. For viable operation H + RQ > r + c",
        "texts": [
            " The site reviews a title by Evesham (2010), which sounds sufficiently indispensable to be cited here (see Bibliography) before this author has had the chance to acquire a copy. Doerfler is evidently no ordinary mathematician, and generously makes his e-mail address available (see his site) for exchanges on his wide-ranging interests. At a lighter level, readers new to the subject might find the account by Earle (1977) more accessible than the specialist sources. Appendix 4 Kinematics of lever-crank drive In Figure A.4.1 the crankshaft rotates clockwise about point O, driving crank-pin P, offset from the crankshaft axis by radial distance r. Crank angle \u03c6 is measured clock-wise from the upper vertical axis. Lever S-R-Q is pivoted at ground point Q. Bell-crank P-R-T pivots at crank-pin P and point R on the lever. Points O-P-R-Q define the classic \u2018four-bar linkage\u2019. Point Smay drive the piston and Tmay drive the displacer \u2013 or T the piston and S the displacer, reversing the direction in which the cycle work is positive"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003874_icra.2015.7139806-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003874_icra.2015.7139806-Figure2-1.png",
        "caption": "Fig. 2. Schematic diagram of the robot model and its coordinates",
        "texts": [
            " Output of the optimization problem is a vector of the desired joint accelerations (q\u0308d) which goes through the inverse dynamic function and gives us the required joint torques for applying to the robot. In this section, we implement our proposed control algorithm on a four-link planar robot in simulation. The controller is to balance the robot on the compliant ground while external disturbances are applying to the robot. The robot consists of a foot, a shank, a thigh and a torso. A schematic diagram of the robot is shown in Fig. 2. The foot has three degrees of freedom (DoF) which are under-actuated and denoted by q1, q2 and q3. In fact, q1 and q2 are the displacements of the origin of the foot in horizontal and 1 vertical directions, respectively, and q3 is the rotation angle of the foot. Thus, the whole robot has six DoF which only three of them (i.e. q4, q5 and q6) are actuated. The lengths of the links and their inertial parameters are mentioned in Table I. In this table, IG is the moment of inertia about the CoM of each link and lc is the location of the CoM of each link with respect to its predecessor joint"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003777_20110828-6-it-1002.01580-Figure5-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003777_20110828-6-it-1002.01580-Figure5-1.png",
        "caption": "Fig. 5. Relation z \u03b8",
        "texts": [
            "4rad/sec: this is a little slower than the one chosen for the weighting function on the tracking error We\u03b8 (this explain the value of \u03b3 > 1). In Roche et al. (2010), an LFT (Linear Fractional Representation)approach has been used to build a LPV model of the AUV by considering the sampling interval as varying parameter. Here, the polytopic approach will be used to take into account the sampling variation in the model formulation (as in Robert et al. (2010)). The geometrical relation between the altitude and the pitch angle will be used to obtain a simple model for the altitude control design(see Figure 5). z = l sin \u03b8 i.e. z\u0307 = l\u0307 sin \u03b8 + l\u03b8\u0307 cos \u03b8 Moreover, the longitudinal speed u is equal to l\u0307: z\u0307 = u sin \u03b8 + l\u03b8\u0307 cos \u03b8 (2) A 1st order limited development of equation (2) for \u03b8\u0307 = 0 and \u03b8 in the neighborhood of 0 implies : z\u0307 \u2243 u\u03b8 Gz = z \u03b8 = 1 p (3) Therefore, the inner loop composed by the non linear model and the pitch controller Gz can be approximated by an integrator. In this case, the obtained controller will be of low order. As emphasized in the introduction, the objective is to handle asynchronous measurements in the control algorithm"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001465_robot.2010.5509554-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001465_robot.2010.5509554-Figure1-1.png",
        "caption": "Figure 1. WMRA coordinate frames",
        "texts": [
            " Redundancy resolution is to be optimally solved to avoid singularities and joint limits. While the end-effector follows a primary trajectory, we introduce a secondary trajectory to be followed by the wheelchair as part of the redundancy resolution and optimization algorithm. Two of the DoFs are provided by the non-holonomic motion of the wheelchair. This subsystem is controlled using 2 input variables: the linear position of the wheelchair along its x-axis, and the angular position of the wheelchair about its z-axis (see figure 1). The planar motion of the wheelchair includes three variables: the x and y positions, and the zorientation of the wheelchair [12]. Assuming that the manipulator is mounted on the wheelchair with L2 and L3 offset distances from the center of the differential drive across the x and y coordinates respectively (see figure 1), the mapping of the wheels\u2019 velocities to the manipulator\u2019s end effector velocities along its coordinates is defined by: c c W cr J J V= \u22c5 \u22c5 (1) where Jc and Jw are the jacobians that map the end-effector velocities to the arm base velocities (without arm motion) and the arm base velocities to the wheels\u2019 velocities respectively and the end effector velocity and manipulator velocity are: T cr x y z = \u03b1 \u03b2 \u03d5        , l c r V  \u03b8 =  \u03b8    , and [ ] xg yg 2x2 xg yg c 2x2 3x1 2x2 6x3 (P S P C ) I P C P S J [0] [0] [0] 1  \u2212 \u22c5 \u03d5 + \u22c5 \u03d5    \u22c5 \u03d5 \u2212 \u22c5 \u03d5    =          , c 2 c 3 c c 2 c 3 c 1 1 5 W c 2 c 3 c c 2 c 3 c 1 1 1 1 3x2 2 2c (l s l c ) c (l s l c ) l l l 2 2J s (l c l s ) s (l c l s ) 2 l l 2 2 l l   \u03d5 + \u03d5 + \u03d5 \u03d5 \u2212 \u03d5 + \u03d5      = \u03d5 \u2212 \u03d5 \u2212 \u03d5 \u03d5 + \u03d5 \u2212 \u03d5     \u2212     where Pxg and Pyg are the x-y coordinates of the end-effector based on the arm base frame, \u03d5 is the angle of the arm base frame, which is the same as the angle of the wheelchair based on the ground frame, and L5 is the wheels\u2019 radius"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001522_detc2009-86119-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001522_detc2009-86119-Figure2-1.png",
        "caption": "Figure 2. CNC hypoid generator with six degrees-of-freedom",
        "texts": [
            "2. Advanced Manufacture of Spiral Bevel Gears on a CNC Hypoid Generating Machine. The CNC machine for generation of spiral bevel and hypoid gears is provided with six degrees-of-freedom for three rotational motions ( 1R , 2R , 3R ), and three translational Downloaded From: http://proceedings.asmedigitalcollection.asme.org/pdfaccess.ashx?url=/data/conferences/idetc/cie2009/70876/ on 03/25/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use 5 Copyright \u00a9 2009 by ASME motions (X, Y, Z, Fig. 2). The six axes of CNC generator are directly driven by the servo motors and able to implement prescribed functions of motions. Two rotational motions are provided as rotation of the workpiece (pinion/gear) ( 3R ) and the rotation that enables the machine to change the angle between the axes of the workpiece and the tool ( 2R ). The third rotational motion ( 1R ) is provided as rotation of the tool about its axis and generally it is related to the cutting process. The following coordinate systems are applied to describe the relations and motions in the CNC generator (Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000177_tro.2006.882921-Figure5-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000177_tro.2006.882921-Figure5-1.png",
        "caption": "Fig. 5. Guaranteed region p p q q near an acute corner. h is chosen at an arbitrary point on line 1 below h .",
        "texts": [
            " Thus, we can find a segment on side two such that the potential change for this segment is zero when the sink is moved from to by . Solve the following equation for : For side two, the point experiences the worst change of the potential. We can find an interval on side one such that , where is the worst potential change for side two, excluding the segment . Then is the worst possible change of the potential when the sink is moved from to . To extend the idea to a region, we make a qualitative argument rather than a quantitative one, because the equations are similar to the obtuse vertex case. For side one in Fig. 5, the change of the average potential of the line segment for the sink which is moving from to is bigger than that of for the sink moving from to . Notice that we choose the length of to be the same as that of , and to be placed at an arbitrary location below . Obviously, the change of the average potential of for the sink moving from to is bigger than that for the sink moving from to , which is actually equivalent to the change of the average potential of for the sink moving from to . Thus, the average change of the potential for the line segment is the worst case for side one",
            " 6 shows the geometry, although we omit the details for brevity. These bounds provide the means for an algorithm to peel off the boundary, as given in Algorithm 2. Algorithm 2: Guaranteeing no pivot points near the boundary Require Ensure perimeter {if not, reduce } if CORNER then if OBTUSE then % Step: where is found by (Fig. 2). % Region: if Convex then Side 1: where satisfies (Fig. 2) Side 2: (Fig. 2) else {Concave} Side 1: (Fig. 3) Side 2: (Fig. 3) end if else {ACUTE} % Step: (Fig. 4) % Region: if Convex then Side 1: (Fig. 5) Side 2: (Fig. 5) else {Concave} Side 1: (Fig. 6) Side 2: (Fig. 6) end if end if else {SIDE} (Fig. 1) end if For a given object, we have checked each vertex and the boundary, and peeled off the boundary all around the object with a finite thickness everywhere, except points on the boundary where the force is zero. The analysis on the existence of a finite resolution (away from critical points) is guaranteed, and suggests a constructive algorithm. The arguments so far prove the existence of a finite resolution where we can guarantee no pivot point exists, except in regions near actual critical points, i"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001927_nme.1620030405-Figure4-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001927_nme.1620030405-Figure4-1.png",
        "caption": "Figure 4. Global and local co-ordinates. (a) ( r , z) global co-ordinates; (b) (U, W) global co-ordinates",
        "texts": [
            " First, equations (2) and/or (3), together with the expression for meridional rotation, are represented as {u(<)} = [+(<)I (4 (4) { U P = 0% u2 x> 3x1 3x8 8x1 where Then, by substituting the co-ordinates of the element nodal points into equation (4), one gets the nodal displacements (q} in terms of generalized co-ordinates : 3x1 3x1 2x1 where (q}T = <u: U; j ~1 xi j Um ~ n > 1 1 Finally, in physical co-ordinates, the stiffness matrix and the load vector become {Q+ (3 = L*-lIT(QI,J Q2a 6x1 2x 1 2x1 8x1 6x1 A REFINED CURVED ELEMENT 499 where {Q1} are the equivalent generalized loads acting on the external nodes, {Q2} are the loads acting on the two internal nodes, and the subscript OL denotes the generalized co-ordinates. For the assembly of elements two different sets of global co-ordinates were employed. In the one, for the analysis of axisymmetric shells with discontinuous meridional slope, the (r, z) co-ordinates were taken as the global co-ordinates, Figure 4(a). For the other cases,'for the analysis of shells with continuous meridional slope, the (U, W) co-ordinates were taken as the global co-ordinates, Figure 4(b). The required transformation is Ti 0 0 [ \" = [ O Ti 0 ) 1\") rlc (7) UlC 0 0 1 8x1 8x8 3x1 8x1 2x1 where T's are 3 x 3 rotation matrices, I is a 2 x 2 identity matrix, and the global displacements of the nodal points for the two cases are respectively, and For the cap element, Ti in equation (7) is replaced by an identity matrix. By using equation (7) the element stiffness matrices and the load vectors, equations (a), can be transformed into global co-ordinates. However, before proceeding with the actual assembly of the stiffness matrices, the internal degrees of freedom must be condensed out from the equations"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003392_s10529-011-0604-x-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003392_s10529-011-0604-x-Figure1-1.png",
        "caption": "Fig. 1 (a) SEM image of a nano-probe with a silicon nitride cantilever and a tip. (b) SEM image of a Pt-UME (shown as a white line) at the end of a tip. (c) Schematic design of the micro-fluidic system immobilizing a cell. (d) Optical image of a target cell, Chlamydomonas reinhardtii trapped in the open micro-fluidic channels",
        "texts": [
            " Cells Unicellular algal cells, Chlamydomonas reinhardtii, or purified chloroplasts from Peperomia metallica were used for oxygen sensing. The pf18 mutant strain of the algal cells was grown in TAP medium and used for oxygen sensing since it exhibits no motility. Chloroplasts purified from Peperomia metallica leaves were also prepared as described elsewhere (Bulychev et al. 1976). Nano-probes with an UME and characterization A nano-probe with a Pt-UME was fabricated using silicon fabrication technology and was employed to detect oxygen evolution inside a single cell. A nanoprobe with a Pt-UME is shown in Fig. 1a and b. The nano-probe was composed of a silicon nitride cantilever and a high aspect ratio tip. The cantilever exhibited a low spring constant of the silicon nitride layer which enabled the tip to be inserted into the cytosolic cell space without disrupting cell physiology. At the end of the cantilever, a needle-shaped tip containing a localized Pt-UME was embedded in the silicon nitride. The detailed nano-probe fabrication process and its characterization are published elsewhere (Bai et al. 2008)",
            " An open micro-fluidic channel system was fabricated to immobilize single cells while allowing access to the nano-probe. An optically transparent PMMA film (50 lm thick) was micro-molded using micro-fabricated silicon wafers. The resulting micro-fluidic system was a channel structure with a width of 20 lm and a depth of 5\u201310 lm. Micro-traps were placed inside the channels to immobilize cells. After the micro-molding, a 100 nm thick gold layer was sputtered on top of the PMMA film and was used as a pseudo-reference electrode during oxygen sensing experiments. Figure 1c shows the schematic design of the micro-fluidic channel system with a cell and a nano-probe. Additional details about the fabrication of the micro-fluidic channel system are published elsewhere (Ryu et al. 2008). Chlamydomonas reinhardtii or purified chloroplasts from Peperomia metallica were delivered through the micro-fluidic channels by capillary flow, as shown in Fig. 1d. The cell solution was contained in the micro-fluidic channels, guaranteeing that cells were kept in an aqueous environment, while immersion of the nano-probe was minimized to reduce the generation of non-faradaic currents. Once a target cell was immobilized, the nano-probe was positioned onto it through an AFM (PicoSPM, Molecular Imaging, Phoenix, AZ), as shown in Fig. 1d. Electrical Set-up for current sensing from single chloroplasts For amperometric oxygen measurements, an electrical setup was built as described in Fig. 2a. A nano-probe linked on an AFM was inserted into a single cell. For oxygen sensing, the thin layer of gold and the Pt-UME were used as a pseudo-reference electrode and a working electrode, respectively. The Pt-UME on the nano-probe was polarized at -0.7 V with respect to the pseudo-reference electrode. While changing the intensity of light (10\u2013220 lmol photons/m2 s), the corresponding current was monitored"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001523_tpas.1972.293335-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001523_tpas.1972.293335-Figure1-1.png",
        "caption": "Fig. 1. Two-Phase Synchronous Machine.",
        "texts": [
            " The small perturbation equations are developed in a general freely rotating reference frame (81 and the saturated equations for the syn- chronous and d-q reference frames are developed and discussed. The concepts of eigenvalues and eigenvectors are given physical meanings so as to show that there is more to modal analysis than saying: the system is stable provided that all its eigenvalues have negative real parts. The eigenvectors yield information on the coupling between the modes and are used to measure the effect of saturation. Description of the Model in the Stationary Reference Frame The specific model chosen is shown diagrammatically in Figure 1. It is a five-winding model having a balanced two-phase stator winding, a direct axis rotor excitation winding and d and q axes damper windings. Windings nos. 1 and 2 are the two-phase coordinates of the actual phase quantities. The magnetic axis of winding 1 is taken as datum, and the generator convention commonly used in the power industry is adopted [ 1, [6]. Positive rotation is counterclockwise. Basic Assumptions and Definition of Air-Gap Quantities The angle 0 gives the instantaneous location of the d axis and fag is the air-gap flux; it lags the d axis by an angle 6g which must be distinguished from the load angle 6, as shown on the vector diagram of Figure 2"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003702_cjme.2013.04.801-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003702_cjme.2013.04.801-Figure2-1.png",
        "caption": "Fig. 2. Analysis model for determining leg experiencing maximum normal contact force",
        "texts": [
            " In the final section, the conclusions are presented. By observing the poses of hexapods while passing along a slope, including ants and crabs, it can be found that the undersides of their bodies move parallel to the plane of the slope. Moreover, the moving processes of their legs are basically equal to or the opposite of each other in one period. To determine which leg of a robot experiences the maximum normal contact force, a demonstration is implemented as follows. The analysis model is shown in Fig. 2. In Fig. 2, object Q contains two legs. The vertical height from the center of gravity of object Q to the plane of the slope is defined as h. G is viewed as the weight of object Q. Points o1 and o2 represent footholds of object Q on the CHINESE JOURNAL OF MECHANICAL ENGINEERING \u00b7803\u00b7 slope. The distance is regarded as l between foothold o1 and foothold o2. Foothold o1 has friction f1 and normal contact force F1, whereas foothold o2 has friction f2 and normal contact force F2. The angle of the slope, which varies from 0\u00b0 to 90\u00b0, is defined as \u03d5"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002390_jmems.2012.2194777-Figure5-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002390_jmems.2012.2194777-Figure5-1.png",
        "caption": "Fig. 5. Simple setup for observing the image of the liquid lens.",
        "texts": [
            " Because the surface tension of the metal blades is fairly small (\u223c20 mN/m), excessive BK-7 will form a convex shape and have a contact angle on the border of the opening aperture, as shown in Fig. 4(b). Next, we fully filled the top chamber with silicon oil, as shown in Fig. 4(c). After that, the silicon oil was sealed with another glass plate (top glass), as shown in Fig. 4(d). In this ID, the base plate and the rotatable disc are combined together, so they are treated as one periphery. To evaluate the lens performance, one simple way is to observe the imaging property of the liquid lens. A simple setup is used to observe the image of an object through the liquid lens, as shown in Fig. 5. Here, we typed small letters from 1 to 9 on a piece of paper as an object. The object was placed at \u223c8 cm behind the lens cell. The lens cell was placed in vertical direction. When the aperture was adjusted to 2r = 3.2 mm, a clear image was observed in white light circumstance, as shown in Fig. 6(a). The image was taken using a digital camera in front of the lens. When the aperture of the ID is decreased by rotating its actuator handle in clockwise direction, the size of the observed image is magnified"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001531_s11012-009-9232-0-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001531_s11012-009-9232-0-Figure2-1.png",
        "caption": "Fig. 2 Wohlhart symmetric mechanism",
        "texts": [
            " The homogeneous form for the transfer matrix is: i\u22121Qii = \u23a1 \u23a3 cos \u03b8i \u2212 cos\u03b1i \u00b7 sin \u03b8i sin\u03b1i \u00b7 sin \u03b8i ai \u00b7 cos \u03b8i sin \u03b8i cos\u03b1i \u00b7 cos \u03b8i \u2212 sin\u03b1i \u00b7 cos \u03b8i ai \u00b7 sin \u03b8i 0 sin\u03b1i cos\u03b1i di 0 0 0 1 \u23a4 \u23a6 (1) The closure condition for a 6R single loop mechanism expresses that the six transfer matrix product is equal with the unity matrix. This condition can also be written: 3Q4 \u00b7 4Q5 \u00b7 5Q6 = 3Q2 \u00b7 2Q1 \u00b7 1Q6 Developing this equation we obtain a twelve equations system. This system is identical with the system obtained by Waldron and cited by Baker [2]. We consider now Wohlhart symmetric mechanism (Fig. 2) with the next geometrical conditions: \u23a7\u23aa\u23a8 \u23aa\u23a9 d1 = d2 = d3 = d4 = d5 = d6 = 0 \u03b11 = \u03b13 = \u03b15 = \u03b1 \u03b12 = \u03b14 = \u03b16 = 2\u03c0 \u2212 \u03b1 a1 = a2 = a3 = a4 = a5 = a6 = a (2) With these conditions (2), the twelve equations system is apparently simplified (Appendix 1). From (A.11) and (A.12) relations we obtain the two first closure equations: \u03b82 = \u03b86 (3) \u03b83 = \u03b85 (4) These two relations, introduced in (A.9) and (A.10), prove that these relations are dependent, so: 4 \u00b7 sin \u03b82 2 \u00b7 sin \u03b83 2 \u00b7 cos\u03b1 \u00b7 ( cos \u03b82 2 \u00b7 sin \u03b81 + \u03b83 2 \u2212 sin \u03b81 2 \u00b7 sin \u03b82 2 \u00b7 sin \u03b83 2 \u00b7 cos\u03b1 ) = cos \u03b81 2 \u00b7 cos \u03b82 \u00b7 sin \u03b83 + sin \u03b81 2 \u00b7 cos \u03b83 (5) Introducing the relations (3) and (4) in (A"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001313_09544100jaero155-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001313_09544100jaero155-Figure2-1.png",
        "caption": "Fig. 2 Circuit diagram of a DC motor",
        "texts": [
            " The control objective is to make the beam of the TRMS tracks a predetermine trajectory. Figure 1 shows the TRMS considered in this investigation. The dynamic model as supplied by the manufacturer has been improved in this study and the DC motors are simulated with respect to the corresponding equations. The TRMS possesses two permanent magnet DC motors; one for the main and the other for the tail propelling. The motors are identical with different mechanical loads. The mathematical model of the main motor, as shown in Fig. 2, is presented in equations (1) to (5). The mathematical model of the remaining parts of the system in vertical plane is described in equations (6) to (8) (see Fig. 3) Uv = Eav + Raviav + Lav diav dt (1) Eav = kav\u03d5v\u03c9v (2) Tev = TLv + Jmr d\u03c9v dt + Bmr\u03c9v (3) Tev = kav\u03d5viav (4) TLv = ktv|\u03c9v|\u03c9v (5) In equation (6) the first term denotes the torque of the propulsive force due to the main rotor, the second Proc. IMechE Vol. 221 Part G: J. Aerospace Engineering JAERO155 \u00a9 IMechE 2007 at UNIV CALGARY LIBRARY on May 25, 2015pig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003909_acs.2480-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003909_acs.2480-Figure1-1.png",
        "caption": "Figure 1. Airplane position relative to the Earth and segment-fixed reference frames.",
        "texts": [
            " Let us define the absolute motion of the aircraft in an Earth-fixed reference frame FE , and to introduce the body and segment-fixed reference frames, FB and FS , respectively. FE , FB , and FS are two-dimensional frames because only the lateral dynamics are considered in this study. Denoting the origin and the coordinate axes of FE by (OE , xE , yE ), then OExE is chosen northward and OEyE points east. FB is a reference frame in which the origin is the mass center of the vehicle, and the axes direction is according to Figure 1. In terms of the segment-fixed reference frame, one of its axis is aligned with the line segment to be followed while the other is chosen to define a clockwise rotation of FE through angle d , where d is the orientation of the desired segment. In real conditions, an airplane is generally exposed to wind that, depending on its inertial direction and magnitude, could affect both the groundspeed/airspeed and course/heading relations. Because the characteristics of the wind have a notable effect on aircraft performance, it is necessary to include them into the equations of motion",
            " Then, the DEs for the coordinates of the flight path in FE are PxE PyE D BTB NVB or PxE D u Ec c C v E s s c v Ec s PyE D u Ec s C v E s s s C v Ec c with uE D uCWN c c CWEc s vE D v CWN s s c WN c s CWE s s s CWEc c where xE and yE represent the inertial position in the x-axis (north) and in the y-axis (east), respectively. Remember that the pitch and roll angles are small so that sin. ; / 0 and cos. ; / 1. Moreover, considering a symmetrical airplane with a rigid spinning rotor placed in front of its body, it can then be considered that, without loss of generality, V acts only on the x-axis (Figure 1). Hence, the following expression can be stated v << 1 u \u00d0 V and consequently PxE D V cos C ! cos ! (2a) PyE D V sin C ! sin ! (2b) where ! cos ! D WN , ! sin ! D WE , ! is the wind velocity, and ! describes the wind direction. The coordinates of the inertial velocity vector in the rotated frame are given by the rotation matrix R. d / as it follows Ppn_s Ppe_s D R. d / PxE PyE (3) where R. d / D cos d sin d sin d cos d represents the complete transformation from FE to FS . The motion of the airplane with respect to a stationary desired straight-line path of angle d can then be expressed from (2) and (3) as Ppn_s D V cos ",
            "1002/acs P r Pr c where r stands for the yaw rate, represents the yawing moment, and c is a constant related to the aircraft moment of inertia. The use of the segment-fixed reference frame is of great interest when making the assumption of a straight-line reference path. In this particular case, FS is preferred, rather than FE , to define the aircraft kinematics because it allows to formulate the path following problem as a regulation problem. As depicted in the geometrical representation from Figure 1, minimizing the position error of the airplane relative to the segment line AB implies that Pe_s ! 0 and Pn_s to move along the segment, where Pe_s and Pn_s are defined in FS , and they are given by (4). Accordingly, we will further assume that the path to be followed consists of combinations of straight-line segments of different orientations. This assumption is not very restrictive because every smooth path can be divided into a series of successive straight-line segments. In addition, we will consider for control design that the wind velocity and direction change slowly such that they can be considered quasi-constant"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002595_0022-2569(69)90050-0-Figure3-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002595_0022-2569(69)90050-0-Figure3-1.png",
        "caption": "Figure 3. Extension to plane mechanism.",
        "texts": [
            " This was expected since this axis concerning as a Br-axis corresponds to a real root of (2.32). As a consequence we have that: /f ro>2x/6, there are always two values of cranklength (pO for which the mechanism has a Ball-axis with excess I in the plane x =0 . There are still an infinity of such mechanisms by varying to. Moreover, owing to the symmetry of the position on hand this Ball-axis is even of excess 2. 2.9. The plane mechanism as a result of the spherical mechanism The spherical crank-slider mechanism on hand lies on a sphere of radius R. In Fig. 3 the zero-position is drawn in the plane x = 0 . The said Ball-point with excess 2 is indicated by Br 5. The crank-angle Pl is taken negative (thus r I =co tgp l <0). The condition to be fulfilled (for Br 5 to be a Ball-point with excess 2) was (see 2.38): ,-~- 3ror~ + 2r~ + 6 = O. (2.41) Putting MA t =a , M1D=b and B. Br~=c we have R R R r I = - , r o = - and cot 7 o = - (2.42) a b C Substituting this into equation (2.41) one finds bZ- 3ab + 2a2 + ~ = O (2.43) which as tends R to infinity tends to b 2 - 3ab + 2a 2 = 0",
            " The problem of finding the path of each point is now reduced to the numerical integration of three simultaneous differential equations. We are especially interested in the motion of Ball-axis with excess 2 (Br 5, see equation 2.34). Moreover, we want to know \"how long\" (in terms of =) this Ball-axis remains in the neighbourhood of the tangent plane of the path in the position ==0. Define tgT=(Z/Y)~ i.e. 7 is the angle between the plane through the X-axis and the Ball-axis, and the plane XO Y. The initial values are (see also equation 2.34 and Fig. 3): ~ = 0 , Xs=O, tgyo=(z/y)s=ro+rori\" +2 ('3.2) r I - - F o The maximum value of 7 is (c~= 180 \u00b0) 7m.~ =~'o+2pi y 7o- ' / (3.3) 2pz thus, for ct=0, 7 ,=0 and for ~=180 \u00b0 , 7 ,= 1. Thus 7. indicates the position of the Ball-axis with respect to the tangent plane defined by (3.2). We have calculated )'n as a function of 7 for various values of Po (with their corresponding values o fp t , according to equation 2.40). In Fig. 4 the result is shown for po= I 1 \u00b0 with p ~ = - 8 - 2 0 9 \u00b0 : (the second solution of pt ~ -6 "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000597_978-3-540-73129-0-Figure6.120-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000597_978-3-540-73129-0-Figure6.120-1.png",
        "caption": "Fig. 6.120. Microanalysis system [source: IFT]",
        "texts": [
            " The pump is currently undergoing clinical testing on patients receiving painkilling medication. In addition to valves, pumps, nozzles and dispensers, microtechnologies make available other modular fluid-flow components such as flow sensors, micromixers and reaction chambers. Customized fluid systems can now be produced solely on the basis of these modular components. Typical applications are microanalysis systems and microdosing systems, for example for dosing medications, chemical reagents, lubricants and adhesives. A microsystem for analysing water (Fig. 6.120) was developed within the scope of a joint project (VIMAS) funded by the German Ministry of Education and Research (BMBF) under the leadership of the Fraunhofer Institute for Solid-State Technology. Using appropriate sensors, this system determines environmentally relevant parameters (concentrations of nitrates, oxygen and carbolic acid; pH values; opaqueness). The dimensions of the base plate are 31 \u00d7 32mm2. Miniaturized lubricating systems are under development at the IMIT. The first application will be for improving the \u2018wick\u2019 lubricating process"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000597_978-3-540-73129-0-Figure6.90-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000597_978-3-540-73129-0-Figure6.90-1.png",
        "caption": "Fig. 6.90. Haptic feedback for steer-by-wire systems",
        "texts": [
            " In contrast, researchers in Hong Kong and Changsha, China have demonstrated that very small MR dampers, if properly tuned, can have a profound effect on mitigating cable galloping even when located very close to the cable anchor location as shown in Fig. 6.89. The trend in vehicle industries toward control-by-wire (steer-by-wire, shiftby-wire, throttle-by-wire, brake-by-wire, etc.) has created a need for highly controllable, rugged, cost-effective haptic devices to provide realistic forcefeedback sensations to the operator, whether the manual device is a wheel, a joystick, a pedal, or a lever. British forklift manufacturer Linde uses MR brakes to control over-steer in their R14 industrial forklift [177]. The R14 vehicle, shown in Fig. 6.90, is an all-electric forklift intended for close maneuvering and manipulation in confined, clean-spaces such as food handling warehouses with large drive-in freezers. There is no mechanical connection between the steering wheel and the ground wheels. Steering is accomplished entirely by electrical control. Rotation of the steering wheel turns an optical encoder, which supplies an electrical signal that is transmitted to the drive ground wheel and causes a motor to orient them in the desired direction. The steering wheel and the optical encoder are both mounted to the shaft of a MR brake. The brake provides a variable amount of rotational resistance depending on the instantaneous vehicular motion and orientation of the ground wheels. Such tactile feedback to the operator is necessary to insure stable operation. The MR brake and magnetic rotary encoder are packaged into a common package as shown in Fig. 6.90 and mount directly to the dashboard of the forklift. As a final example of a MR fluid controlled adaptronic system, the smart prosthesis knee developed by Biedermann Motech GmbH [178\u2013181] is presented. This system shown in Fig. 6.91 is a complete artificial knee that automatically adapts and responds in real-time to changing conditions to provide the most natural gait possible for above-knee amputees. The heart of this system is a small magnetorheological fluid damper that is used to semi-actively control the motion of the knee based on inputs from a group of sensors located in the prosthesis"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003003_j.jlumin.2013.08.052-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003003_j.jlumin.2013.08.052-Figure1-1.png",
        "caption": "Fig. 1. Schematic diagram of the FI-CL system.",
        "texts": [
            "0500 g Jacobsen's catalyst in 8 mL of 95% ethanol and diluted with water to100 mL. All working solutions were stored in a refrigerator and protected from light. The CL study was performed on a model IFFM-E flow injection CL analyzer (Remax, Xi'an, China), which consisted of a model IFFM-E flow injection system and a model IFFS-A luminometer. The UV\u2013vis absorption spectra were studied on an UV-1800 spectrophotometer (Shimadzu, Japan). CL spectra were measured on a RF-5301 spectrophotometer (Shimadzu Corporation) combined with a flow-injection system. As depicted in Fig. 1, all solutions were delivered by two peristaltic pumps (P1, P2). P1 was used to deliver Jacobsen's catalyst solution (tube a) and H2O2 solution (tube b) at a flow rate of 1.3 mL/min. luminol in NaOH solution (line c) and sample or blank solution (line d) were delivered by P2 at a flow rate of 1.8 mL/min. Polytetrafluoroethylene (PTFE) tubing (0.8 mm i.d.) was employed to connect all components in the flow system. In order to get the best conditions and good stability, the flow injection system were run for at least 10 min before the first measurement"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003463_1.4004588-Figure3-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003463_1.4004588-Figure3-1.png",
        "caption": "Fig. 3 Applied coordinate systems",
        "texts": [
            " Although many different edge blunting methods have been proposed and many related research works have been carried out, they are all resulted two-dimensionally. To understand the different profile more clearly and design the profile of the rotor upon flow capacity and working duty of the pump, it is necessary to research the different profiles three-dimensionally, because the pump is three-dimensional. So, the paper calculates performance parameters of the new proposed profile, and it is compared with the existing designs three-dimensionally. Obtained results provide important reference to design efficient pumps. 2.1 Coordinate System. As shown in Fig. 3, coordinate systems S1, S2, and Sf are rigidly connected to the driving rotor, the driven rotor, and the frame, respectively. Due to the equal radii of the operating pitch circle, both of the rotors rotate with the same angle velocity. R1 and R2 are the equations of the transverse profiles of the rotors presented in S1 and S2, respectively. The coordinate transformation in transition from S2 to S1 is based on the matrix equation as follows: R1 \u00bcM12R2 \u00bcMf 1Mf 2 R2 (1) Where Mf 2 \u00bc cos/ sin/ r sin/ cos/ 0 0 0 1 2 64 3 75; Mf 1 \u00bc cos/ sin/ r cos/ sin/ cos/ r sin/ 0 0 1 2 64 3 75 M12 \u00bc cos 2/ sin 2/ Ecos/ sin 2/ cos 2/ E sin/ 0 0 1 2 64 3 75 1Corresponding author"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001274_cdc.2009.5400332-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001274_cdc.2009.5400332-Figure1-1.png",
        "caption": "Fig. 1. State space plot of sample executions x(t).",
        "texts": [
            " X0 \u2229 XT = \u2205 and \u2200x0 \u2208 X0, \u2203\u03c4 > 0, t f > \u03c4 : x[t0,t f ](t, x0, \u03c4) \u2208 XT , A2) The switching time constraint (5) is not effective at any locally optimal solution M\u0304\u22c6, and A3) The functions fk, Lk, \u03d51 and \u03c8 are at least C2, timeinvariant and sufficiently regular. A1 and A2 are necessary, but not sufficient for feasibility of Problem 1. Clearly, the mere existence of at least one admissible trajectory x[t0,t f ](t, x0, \u03c4) for each initial state does not ensure that all associated switching points x1 are located on a (n \u2212 1)-dimensional surface in the state space. A3, on the other hand, ensures a certain regularity of the solution M\u0304\u22c6, if the solution exists. Fig. 1 illustrates the considered problem setting in state space. Both depicted trajectories originate from X0 and intersect the surface M\u22c610 at switch points x\u22c61 , x\u2032\u22c61 , for which the subsequent evolutions eventually enter the terminal region XT . While switching at any point other than x\u22c61 , x\u2032\u22c61 must not necessarily lead to a violation of the terminal constraints (6), the corresponding trajectory will accrue higher costs. Remark 1: In contrast to most literature on optimal control of switched autonomous systems, the control problem 1 is cast over a variable time horizon, which is crucial for the existence of stationary optimal switching surfaces (2)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002148_j.otsr.2009.11.005-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002148_j.otsr.2009.11.005-Figure1-1.png",
        "caption": "Figure 1 Photograph of the special custom-made jig for torsional testing including the specially constructed rotation sensor.",
        "texts": [
            " Before biomechanical testing, the frozen tibiae were hawed at room temperature and kept moist using gauze oaked in 0.9% NaCl solution during the entire test period. he embedded specimens were mounted in specially contructed jigs for compressive, 4-point-bending, or torsional esting in a Material Testing System (MTS [Model 858, MTS orp., Minneapolis, USA]). The order of stiffness testing as randomised and measurements were performed on each pecimen by using compressive, bending (anteroposterior nd mediolateral) and torsional load. The resulting deforation was detected by custom-made compression, torsion Fig. 1) and deflection (Fig. 2) sensors (LVTD and precision otentiometer). For each stiffness testing-procedure, a preonditioning of 10 cycles was conducted before the actual esting in order to assure repeatability. The callus tissue ithin the specimens was loaded during the different types f testing up to 15 Nm for torsional, to 750 N for compresive and to 6.5 Nm for bending load. During testing, load nd deformation were simultaneously recorded in order to etermine stiffness, which is defined as the slope of the Stiffness of callus tissue during distraction osteogenesis 157 F d i l s t b S S t ( t o c s a s t w r l a S d t c o a o R T c 5 igure 4 Comparison of regained stiffness characteristics of istracted tibiae at the 74th postoperative day in relation to ntact tibiae"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003246_1.4004116-Figure7-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003246_1.4004116-Figure7-1.png",
        "caption": "Fig. 7 Meshing sketch of circular gear and rack",
        "texts": [
            " Point A1 is the crossover point of the tooth profile of noncircular gear and its pitch curve, while A2 is the crossover point of the tooth profile of the shaper and its pitch circle. The angle between lines AN and P1P2 is the profile angle a of the shaper cutter. The following vector equation is derived based on the geometric relationship in Fig. 6 ON \u00bc OA\u00fe AN (24) If the pitch curve formula of deformed limacon gear is r \u00bc r\u00f0u\u00de, the vector angle and module of OA is u and r, respectively. The module of vector AN equals to the distance between the instantaneous rotation center and meshing point of tooth profile. Figure 7 shows the meshing relationship of a circle gear and a rack. The pitch circle is tangent to the pitch line at the point B. Line BN is perpendicular to tooth profile of the rack. The tooth profile of the circular gear and the pitch circle intersect at B1, while the tooth profile of the rack and the pitch line intersect at B2. Since the pitch circle rolls on the pitch line without sliding, jB2Bj equals to B1B _ .Fig. 3 Pitch curves of two kinds of gear pairs Fig. 4 Transmission ratio curves of two kinds of gear pairs 061004-4 / Vol"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002513_3.4984-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002513_3.4984-Figure1-1.png",
        "caption": "Fig. 1 Schematic of vehicle in orbit.",
        "texts": [
            " The computations were performed on an IBM 360 generously made available by the Computing Center at the University of Kentucky. This investigation was supported in part by a 1967 Summer Research Fellowship granted by the Graduate School of the University of Kentucky. * Assistant Professor of Engineering Mechanics. Member AIAA. influence of the tether on vehicle stability and some illustrative examples to demonstrate the interrelationships among system parameters. The fourth part summarizes the conclusions to be drawn from this investigation. Description of Vehicle Dynamics Figure 1 is a schematic of a vehicle in a circular orbit of radius R about a spherically symmetric body E of mass M whose center E* is considered fixed in a Newtonian reference frame N. The vehicle, whose mass center is F*, is comprised of two identical, unsymmetrical, tethered rigid bodies B and C with mass centers B* and (7*, respectively. AI, At, and A3 are mutually perpendicular orbital reference axes with origin at F* and an orbital angular rate 12 in N. AI is the radial line passing through E* and F*, A2 points in the direction of the motion of F*, and A% is normal to the plane of the orbit"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000587_2007-01-2232-Figure14-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000587_2007-01-2232-Figure14-1.png",
        "caption": "Fig. 14 Vibration Mode of Power Plant Bending Resonance",
        "texts": [
            " 13), and the mass is approximately 35 kilograms greater. There was concern that the increase in mass and overall length would cause the power plant resonance to drop into the normal engine speed range. An increase in the transmitted force was also predicted, because the increase in the load allocated to the mounts would require that the spring constant for the mounts be raised to 1.5 times that of the mounts for a 6-speed A/T. The deformation mode of the power plant resonance is such as to cause large deformation in the transmission, as shown in Fig. 14. A study was undertaken of how to improve the resonant frequency so as to separate it from the normal engine speed range. FEM analysis was used to optimize the shape of the transmission case by smoothing the outline, reinforcing the ribs, and so on. As a result, the resonant frequency was raised to 180 Hz, equivalent to that of a 6-speed A/T. Also, the mounting position was shifted 80 millimeters farther forward than in the original plan in order to set it at a nodal point of the vibration mode (Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002028_icnsc.2010.5461545-Figure6-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002028_icnsc.2010.5461545-Figure6-1.png",
        "caption": "Fig. 6. The damaged gear",
        "texts": [
            " A magnetic loading system is connected to the output shaft of the output spur gear. The magnetic loading system is controlled by a power supply and the load can be adjusted by changing the output current of the amplifier. Three accelerometers were mounted on the input drive pinion and at the locations near the output driving gears and two acoustic emission sensors were mounted at the locations near the output driving gears. In the experiments, 20% of a tooth in one of the driving gears was chipped. The damaged gear is shown in Fig. 6. The damaged gear is placed at location 2. The locations of the accelerometers and the acoustic emission sensors are shown in Fig. 7. During the experiments, the input speed was kept at 3600 rpm. The sampling rate for vibration signals was set to be 102.4 kHz. Vibration data for both the damaged gearbox and the healthy gearbox were collected. For each case, there were totally 200 datasets sampled. For AE signals, the sampling rate was set to be 5 MHz. AE data for both the damaged gearbox and the healthy gearbox were collected"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001690_iros.2009.5353974-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001690_iros.2009.5353974-Figure1-1.png",
        "caption": "Fig. 1. Coordinate system of a rigid body",
        "texts": [
            " The attitude control problems of rigid bodies such as the stabilization and navigation require the transformation of measured and computed quantities between various frames of references. The position and the attitude of a rigid-body is based on measurements from sensors attached to a rigidbody. Indeed, inertial sensors (accelerometer, gyro,. . . ) are attached to the body-platform and provide inertial measurements expressed relative to the instrument axes. In most systems, the instrument axes are nominally aligned with the body-platform axes. Since the measurements are performed in the body frame we describe in Fig. 1 the orientation of the body-fixed frame B(xm, ym, zm) with respect to the inertial reference frame RI (xa, ya, za). Various mathematical representation can be used to define the attitude of the rigidbody with respect to coordinate inertial reference frame. In this paper, we consider the Euler angles representation in which a transformation from one coordinate frame to another is defined by three successive rotations about different axes taken in turn. The Euler rotation angles used here corresponds to the following rotation sequence: yaw(\u03c8)-pitch(\u03b8)roll (\u03c6)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000597_978-3-540-73129-0-Figure6.12-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000597_978-3-540-73129-0-Figure6.12-1.png",
        "caption": "Fig. 6.12. Definition of the axes in piezo materials. a The digits 4, 5 and 6 indicate the shear on the axes 1, 2 and 3; b longitudinal (d33) effect, c transversal (d31) effect",
        "texts": [
            "2), the electric displacement density D and the mechanical strain S are combined with the mechanical stress T and the electrical field strength E: D = dT + \u03b5TE (6.1) S = sET + dtE . (6.2) In this system of equations the piezoelectric charge constant d indicates the intensity of the piezo effect; \u03b5T is the dielectric constant for constant T and sE is the elastic compliance for constant E; dt is the transpose of matrix d. The mentioned parameters are tensors of the first to fourth order. A simplification is possible by using the symmetry properties of tensors. Usually, the Cartesian coordinate system in Fig. 6.12a is used, with axis 3 pointing in the direction of polarization of the piezo substance (see below) [5, 6]. All material dependent parameters can be described by matrices, whose elements are marked with double indices. In d, the first index marks the orientation of E, the second the direction of S. The examples in Fig. 6.12b and c are based on the condition that the field strength works in the direction of the polarization. The resulting elongation in Fig. 6.12b points as well in direction 3 (longitudinal effect), in Fig. 6.12c however, it works in direction 1 (transversal effect). These two characteristics of the piezoelectric effect are quantified by means of the piezo constants d33 and d31. It is common to summarize all matrix elements in so-called coupling matrices. From the coefficients in the coupling matrix it is possible to determine an important parameter of piezo materials, the coupling coefficient k. For the coupling coefficient of the longitudinal effect k33 applies for instance k33 = d33\u221a sE33\u03b5 T 33 . (6.3) Since k2 corresponds to the ratio of stored mechanical energy to absorbed electrical energy, achieving actuators with high elongation efficiency requires substances with a large k",
            " Similarly, it is possible to connect two thin ceramic strips one of which shortens while the other expands (bimorph). One can distinguish between two designs: in the series bimorph, the polarization of the two piezo layers is inversely arranged, while it is codirectional in the parallel bimorph (see Fig. 6.18). Compared to stack translators, bending elements feature a greater deflection, lower stiffness, smaller blocking force and lower eigenfrequency. Recently, Physik Instrumente [7] started offering a line of actuators based on the strong d15-effect (shear effect). According to the definitions in Fig. 6.12a, the quantities E and S work along the axes 1 and 5, i. e. upon applying a voltage the piezo element experiences a shearing motion about its axis 2. Making use of this effect, the end surfaces of block-shaped elements without casing (cross sections of 3\u00d73 to 16\u00d716mm2) are shifted by up to 10 \u03bcm with respect to each other, while the shearing loads are limited to 300N. By stacking two such elements, a x\u2013y positioner can be created. Adding a third piezoceramic element based on the d33-effect results in a 3-axis positioning system"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000597_978-3-540-73129-0-Figure6.72-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000597_978-3-540-73129-0-Figure6.72-1.png",
        "caption": "Fig. 6.72. Digital ER motion synthesizer (concept): y direction is shown ER controlled in switched steps of equal time elements, with steady speed traverse in (say) a lattice in the x direction",
        "texts": [
            " Without getting involved with digital technology: if the x direction speed provided is constant, then the y penetration (driven by a bang-bang application of voltage and a yield stress of sufficient magnitude to give the relevant part high and instant acceleration) must be maintained over a very small time interval (fixed by the switching speed) if the resolution is not to be too crude. DC operation seems virtually mandatory, with any hysteretic and electrophoretic tendencies being arrested by a conjunction of binary switching and high \u03b3\u0307 (see Fig. 6.72). It does not seem possible to provide a figure of merit for a fluid that possesses these sundry needs, but (see Sect. 6.6) the linear traverse mechanism will demonstrably test total capability in that respect [115]. In this device, two contra-rotating, high-inertia, constant-velocity rotors provide motion sources with HT (high tension) and earth \u2018busbars\u2019, the excitation being controlled via switches. Two driven clutches, spaced from their drivers co axially by the ERF, are each connected to a pulley, both of which are connected by a belt"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003189_j.mimet.2011.09.006-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003189_j.mimet.2011.09.006-Figure1-1.png",
        "caption": "Fig. 1. (a) Schematic view of a micro chamber prepared on a glass slide. (b) A typical micrograph of Navicula sp. cells in a micro chamber. The trajectory of a cell over 60 s was over-written. \u201cO\u201d and \u201cE\u201d indicate the origin (t=0 s) and end (t=600 s) positions. The cell turned at position Nos.1, 2, 3, and 4.",
        "texts": [
            " The observed cell movements were captured as avi files (one flame per second) by a digital camera system (DP72, Olympus, Tokyo, Japan). The tracks for 39 cells were calculated from the avi files using two-dimensional video analysis software (Move-tr/2D 7.0, Library, Tokyo, Japan). The coordinate positions for individual cells in the chamber were estimated by the software, and the velocity, acceleration, and distances moved were calculated from the coordinate data. To avoid the effect of coordinate fluctuation, the velocity, acceleration, and distance values were estimated as an average of 20 s. Fig. 1a shows a schematic view of a micro chamber. Diatom cells were pre-cultured on a bare glass slide that was placed into a petri dish with 35 mL of culture medium. After 7\u20139 days of cultivation, an acryl plate with 9 holes and a glass cover slip was put onto the glass to secure the diatom cells inside. Because the diameter of a hole and the thickness of the plate were 600 \u03bcm and 1 mm, respectively, the volume of a chamber was almost 1 mm3. In this way, most of the cells in a chamber did not escape during the 10 min observation period. An example of an optical microscopy image of some diatom cells in a micro chamber is shown in Fig. 1b. The trajectory of a motile cell, which was estimated by the two-dimensional video analysis software, was written over the image. \u201cO\u201d and \u201cE\u201d indicate the original (t=0 s) and end (t=600 s) positions of the cell, respectively. No.1 to No.4 indicate the turned positions. The cell came into contact with the wall of the chamber at position No.1, and returned to a position just before the wall at position No.3 and No.4. At position No.2, the cell turned, although it did not come into close contact with the wall"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003531_j.piutam.2011.04.015-Figure6-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003531_j.piutam.2011.04.015-Figure6-1.png",
        "caption": "Fig. 6: The stance model A), the swing model B) and the double stance model C) used by the control system to compute the hip, knee and ankle torques. The actual foot contact model consists of a 2-part model that includes volumetric foot contact pads D), while the abstraction used by the control system treats the foot as a rigid link that rotates about a pin joint that is translating along a prismatic joint E)",
        "texts": [
            " Instead, an input-output feedback linearization control law is calculated for an approximate model (detailed in Sec. 3.1) which has a rigid foot. Since the control model is an approximation of the gait model, feedback control \u2014 supplied by the error terms (\u03bdx,\u03bdy and \u03bd\u03b8) \u2014 is necessary to ensure that the orientation of the torso is regulated and its position converges with the COM location of the SLIP model. x\u0308M = x\u0308S + \u03bdx (32) y\u0308M = y\u0308S + \u03bdy (33) \u03b8\u0308M = \u03bd\u03b8 (34) Hip, knee and ankle torques that satisfy Eqns. 32-34 are computed for the multibody gait model using a series of control models (Fig. 6) that have a simplified foot, making it possible to use input-output feedback linearization [17, 18]. As with the ASLIP model, a set of additional heuristic equations are introduced during double stance to permit a unique set of joint torques to be computed to satisfy Eqns. 32-34 in this overactuated pose (here the torso has 3 dof, and there are 6 joint torques that can be applied). Since the legs now have mass (in contrast to the previous models) a swing controller is required to guide the leg from its final push-off position to its contact position in a specific amount of time",
            " When the foot pads are touching the ground, but have not reached steady state compression, the foot pads behave like a spring of low stiffness. The transient low stiffness of the foot pads greatly limits the ability of the leg to apply a desirable force and torque to the hip joint. As the pads of the foot compress, their apparent stiffness significantly increases and can be approximated as being rigid giving the leg greater control authority over the torso. While a compliant foot is used for the dynamic model (Fig. 6.D), a geometrically equivalent but rigid foot is used for the control model (Fig. 6.E). During the transient contact phase when the apparent stiffness of the foot pads is quite low, the control model and the dynamic model differ. The two models are made to converge to one and other as the foot pads reach steady state compression by using feed-back control to augment the desired torso accelerations (Eqns. 32- 34). Each of the feedback error terms (\u03bdx,\u03bdy and \u03bd\u03b8) take the form of a state feedback PD controller: \u03bdx = \u2212Kx(xM \u2212 xS ) \u2212 Dx(x\u0307M \u2212 x\u0307S ) (35) \u03bdy = \u2212Ky(yM \u2212 yS ) \u2212 Dy(y\u0307M \u2212 y\u0307S ) (36) \u03bd\u03b8 = \u2212K\u03b8(\u03b8M \u2212 \u03b80) \u2212 D\u03b8(\u03b8\u0307M). (37) Input-output feedback linearization [17, 18] is used to compute the hip, knee, and ankle torques required to accelerate the torso of the multibody model such that Eqns. 32-34 are satisfied. The input-output feedback linearization control expressions are not formulated using the multibody model (Fig. 5) \u2014 due to the difficulties the full foot model introduces \u2014 but with an approximate single stance model (Fig. 6A) that includes a rigid foot. To form the control law, we first begin with the equations of motion of the stance control model (Fig. 6A) in functional form (using square brackets to denote matrices). \u0308\u03b3S S = [MS S ]\u22121 4\u00d74 ( \u2212 CS S + [PS S ]4\u00d73 { \u03c4S S } 3\u00d71 + [QS S ]4\u00d73 { FS W \u03c4S W } 3\u00d71 ) (38) In Eqn. 38 \u03b3S S is the vector of joint angles (and respective derivatives) of the single stance (SS) control model (Fig. 6A), [MS S ] is the mass matrix, CS S the vector of Coriolis, centripetal and gravitational forces; [PS S ]4\u00d73 is the matrix that transforms the joint torques { \u03c4S S } 3\u00d71 into generalized forces; and [QS S ]4\u00d73 is the matrix that transforms the reaction force and torque vector { FS W , \u03c4S W } 3\u00d71 (that the swing limb applies to the pelvis) into generalized forces. The variables used to describe the general multibody terms in Eqn. 38 are used throughout this chapter. The accelerations of the torso (x\u0308M , y\u0308M , \u03b8\u0308M) can be expressed as a linear combination of the joint accelerations of the stance model",
            " \u23a7\u23aa\u23aa\u23aa\u23a8\u23aa\u23aa\u23aa\u23a9 x\u0308S + \u03bdx y\u0308S + \u03bdy \u03b8\u0308S + \u03bd\u03b8 \u23ab\u23aa\u23aa\u23aa\u23ac\u23aa\u23aa\u23aa\u23ad = [TS S ]3\u00d74 [MS S ]\u22121 4\u00d74 ( \u2212 CS S + [PS S ]4\u00d73 { \u03c4S S } 3\u00d71 + [QS S ]4\u00d73 { FS W \u03c4S W } 3\u00d71 ) (40) Once the state of the multibody model has been mapped to an equivalent state of the stance model, Eqn. 40 becomes a system of three equations with three unknowns (the three components of \u03c4S S ) making it possible to compute values of the hip, knee and ankle torques that will satisfy Eqns. 32-34. The hip, knee and ankle states can be mapped directly from the multibody model to the stance control model. \u03b31 S S = \u2212(\u03b8M \u2212 \u03c0 2 + \u03b1M) (41) \u03b32 S S = \u2212\u03b2M (42) \u03b33 S S = \u2212\u03b3M (43) \u03b3\u03071 S S = \u2212(\u03b8\u0307M + \u03b1\u0307M) (44) \u03b3\u03072 S S = \u2212\u03b2\u0307M (45) \u03b3\u03073 S S = \u2212\u03b3\u0307M (46) The geometry of the foot of the control model (Fig. 6A) \u2014 the length of the link between the COM of the foot and the revolute joint attached to the ground \u2014 is adjusted so that the revolute joint attaches to the ground at a location that coincides with the COP of the foot of the multibody model (Fig. 5). The angular velocity of the stance model foot, and the translational velocity of the COP of the stance model (\u03b3\u03074 S S in Fig. 6A and x\u0307COP in Fig. 6E) are computed such that the translational velocity of the ankle joints of the stance control and multibody gait model match. Once the swing foot comes into contact with the ground, the controller changes its internal state from single stance to double stance (Fig. 7), and employs a completely different control model (Fig. 6C), for which a new control law must be derived. As before, input-output feedback linearization is applied to an approximate double stance model to yield hip, knee and ankle torques for both legs that will satisfy Eqns. 32-34. The derivation begins by computing the net force and torque that the two legs must apply to the torso to satisfy Eqns. 32-34. fMx = m(x\u0308M + \u03bdx) (47) fMy = m(y\u0308M + \u03bdy + g) (48) \u03c4M = J\u03bd\u03b8 \u2212 L( fMx sin(\u03b8M) \u2212 fMy cos(\u03b8M)) (49) Scalars fMx, fMy, and \u03c4M are the net force and torque that the two legs must apply to the torso (of mass m, inertia J at an orientation of \u03b8M as before) to satisfy Eqns",
            " As before with the ASLIP model, extra heuristic equations are introduced to divide the load between the two legs in proportion to the contact force beneath the respective foot of the multibody model. F1 DS \u00b7 x\u0302 f y1 M + F5 DS \u00b7 x\u0302 f y2 M = 0 (53) F1 DS \u00b7 y\u0302 f y1 M \u2212 F5 DS \u00b7 y\u0302 f y2 M = 0 (54) \u03c41 DS f y1 M \u2212 \u03c4 5 DS f y2 M = 0 (55) The system of six equations (Eqns. 50-52 and Eqns. 53-55) can be solved for the forces and torques that each leg must apply to hip joint of the torso to satisfy Eqns. 47-49 and Eqns. 50-52. The force and torque that each leg applies to the hip joint of the torso can be used in combination with the equations of motion of the double stance model (Fig. 6C) to compute the remaining knee and ankle torques that each leg must generate. The equations of motion of the double stance model in functional form are [MDS ]9\u00d79 { \u0308\u03b3DS } + CDS + { 05\u00d71 DDS ,4\u00d71 \u03bb } = [PDS ]9\u00d76 { \u03c4DS } 6\u00d71 . (56) Position constraint equations DDS have been used to model the hip joints (rather than using joint coordinates) to make it possible to solve for the force that the legs apply at this joint. The reaction force at the hip can now be expressed as { F1 DS F5 DS } 4\u00d71 = [BDS ]4\u00d74 { \u03bb } (57) where matrix [BDS ]4\u00d74 is a matrix that transform the Lagrange multipliers into reaction forces",
            " 56 we have [MDS ]9\u00d79 { \u0308\u03b3DS } + CDS + \u23a7\u23aa\u23aa\u23aa\u23aa\u23a8\u23aa\u23aa\u23aa\u23aa\u23a9 05\u00d71 DDS ,4\u00d71[BDS ]\u22121 { F1 DS F5 DS } 4\u00d71 \u23ab\u23aa\u23aa\u23aa\u23aa\u23ac\u23aa\u23aa\u23aa\u23aa\u23ad = [PDS ]9\u00d76 { \u03c4DS } 6\u00d71 (58) Since the two hip torques (\u03c41 DS and \u03c45 DS ) and forces ( f 1 DS and f 5 DS ) are known from the solution to Eqns. 50-52 and Eqns. 53-55, Eqn. 58 has embedded in it a set of four equations (the constraint equations) that are linear in four unknowns (\u03c42 DS , \u03c43 DS , \u03c46 DS , and \u03c47 DS ). After the state of the multibody model (Fig. 5) is mapped to the equivalent state of the double stance control model (Fig. 6) \u2014 using the same procedure detailed in Sec. 3.1 \u2014 Eqn. 57 can be solved for the remaining knee and ankle torques required to satisfy Eqns. 47-49. The swing phase has been a topic of robotics research for many years and has resulted in a number of standard approaches: active trajectory tracking [4, 30], passive swing [11, 31], and a combination of passive and active swing techniques [32]. Although a lot of research has been done on the topic of swing, little of it is directly applicable to formulating a control law that will yield a human-like swing phase",
            " Beginning the swing phase passively and finishing with trajectory tracking [32] seems like a logical approach, though care must be taken to blend the two phases in a manner that does not cause torque transients. For this preliminary investigation, optimization was used to pre-compute a human-like swing trajectory. During the multibody simulation, the swing limb was driven to follow the pre-computed optimal swing trajectory using a computed torque controller with feedback. Human-like swing kinematics that fit the swing phase of the target SLIP model were found by searching for a trajectory that minimized a convex function of joint work for the swing model (Fig. 6B). A convex function of joint work was employed to crudely emulate the increased metabolic cost of eccentric and concentric contractions relative to isometric contractions [34]. Note that the joint angles of the swing model are represented using the variables \u03c81 S W , \u03c82 S W , \u03c83 S W in the place of \u03b1M , \u03b2M and \u03b3M for convenience. min 3\u2211 i=1 \u222b t f t0 (\u03c4i S W \u03c8\u0307 i S W )2dt (59) Unlike the stance model, bandwidth-limited joint torque actuators were used during the optimization process. It was critical to use bandwidth-limited joint torque actuators to prevent the optimization algorithm from converging on a solution that required sharp changes of joint torque outside of human capabilities"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002335_detc2010-28926-Figure6-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002335_detc2010-28926-Figure6-1.png",
        "caption": "Figure 6. Example for decomposition of a linkage into Assur Graphs.",
        "texts": [
            " In order to get as small matrices as possible for analysis, the analysis is done through the decomposition graph, but this time in the composition order. First, the analysis is done on the Assur Graph that all of its outer vertices are ground vertices, i.e., vertices that their velocities are known. After calculating the velocities of inner vertices of the Assur graph, this AG is then deleted and its inner vertices are replaced with ground vertices. This process ends when all the vertices of the linkage are grounded. This analysis process is demonstrated by solving the velocities in the mechanisms appearing in Figure 6a. First, the driving link is replaced with a ground vertex and a structural scheme is constructed (Figure 6b), for which a decomposition graph with three Assur Graphs is constructed as shown in Figure 6c. a) The linkage. b) The structural scheme. c) The decomposition graph. In this example, the first Assur Graph to be analyzed is the tetrad \u2013 (B,C,D,J) where the two outer vertices, A and p4, are ground vertices thus the inner velocities \u2013 B,C,D,E,J can be calculated. The second AG that can be analyzed is (G,H,I) or the dyad F. In this example, the second AG chosen, arbitrarily, to be analyzed is the dyad F, where now the velocity of the outer vertex \u2013 J is known from the previous AG, as shown in Figure 7b,b1"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002510_epepemc.2010.5606557-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002510_epepemc.2010.5606557-Figure1-1.png",
        "caption": "Fig. 1. Inverter switching states and resulting voltage vectors",
        "texts": [],
        "surrounding_texts": [
            "The stator voltage equation (4) is: us s = Rsi s s + \u03c8\u0307s s (4) The flux linkage \u03c8s s of the RSM is only caused by the stator currents and varied in magnitude and orientation as a function of the the rotor angle. \u03c8s s = \u03c8s s (iss, \u03b8) (5) = T\u03c8 r s (irs) (6) = T\u03c8 r s ( T \u22121iss ) (7) The function \u03c8r s (irs) is a one-to-one assignment of a current vector to a flux linkage vector in rotor fixed frame. It can be linear or nonlinear but for an RSM it has to be anisotropic. Equation (4) requires the derivation of (7) with respect to the time. \u03c8\u0307s s = \u2202T \u2202\u03b8 d\u03b8 dt \u03c8r s + T \u2202\u03c8r s \u2202irs dirs dt (8) = L s si\u0307 s s + J\u03c9\u03c8s s \u2212 L s sJ\u03c9i s s (9) with L s s = TL r sT \u22121 and L r s = \u2202\u03c8r s \u2202irs (10) Using (9) the voltage equation (4) can be transposed to calculate the current derivative i\u0307ss. i\u0307ss = L s s \u22121 (us s \u2212Rsi s s \u2212 J\u03c9\u03c8s s + L s sJ\u03c9i s s) (11) Considering the number of pole pairs p the vector product of current and flux linkage gives the mechanical torque M of the machine. M = p is T s J \u03c8s s (12) The rotor speed and angle are obtained by torque integration. \u03c9 = p \u0398 \u222b (M \u2212ML) dt , \u03b8 = \u222b \u03c9dt (13)"
        ]
    },
    {
        "image_filename": "designv11_12_0000277_j.1751-1097.1977.tb09169.x-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000277_j.1751-1097.1977.tb09169.x-Figure2-1.png",
        "caption": "Figure 2. Plot of I l k (apparent) versus [H'] for the oxidation of flavin semiquinones by oxygen in 0.1 M sodium phosphate--I mM EDTA. x , 8x-[N(3thistidyl]-riboflavin; 0, 8a-[N(l)-histidyI]-riboflavin; and 0, 8a-[N(3)-(N acetyl-N( 1)-methylhistidyl)]-tetraacetylriboflavin. For each of the above flavin analogs, the flavin concentration was 10 p M in 10 cm path length cells; the oxygen concentration was 250 pM.",
        "texts": [
            " the intrinsic rate of anion flavin semiquinone oxidation by oxygen is independent of the pH). This behavior is observed with all 8a-substituted flavins tested and is demonstrated for 8cr-[N(3)-histidyl]-riboflavin in Fig. 1. To determine the pK for neutral semiquinone ionization as well as the second order rate constant for anion semiquinone oxidation by oxygen, experimental data such as those shown in Fig. 1 were analyzed by the method of Vaish and Tollin (1971). Plots of the reciprocal of the apparent second-order rate constants vs [H'] are linear, as demonstrated in Fig. 2 for the N(3) and N(1) isomers of Sa-histidylriboflavin and for 8a-[N(3)-(N acetyl-N(1)methylhistidyl)ltetraacetyl riboflavin. The intercepts give the reciprocal of the second-order rate constant for anion semi- *Standard error analysis of the data showed the pK values to have a 0.05 uncertainty while the values for the second-order rate constants had an uncertainty of 0.01. Standard errors for the individual pseudo-first order rate constants ranged from k 2% to 20% as the pH was increased. quinone oxidation by oxygen ( k ) and the slope is the reciprocal of the product of k and the equilibrium constant for neutral semiquinone ionization (K). A comparison of the data in Fig. 2 with the data of Vaish and Tollin (1971) for riboflavin and lumiflavin show that differences were found in both the slopes and the intercepts of the plots. This demonstrates that 8r-substitution has a substantial effect on the pK for neutral flavin semiquinone ionization and on the rate constant for radical oxidation. The pK values given in Table 3 show that, irrespective of the nature of 8a-modification, the pK for neutral serniquinone ionization is lowered by 1-1.5 pH units as compared with the value for unsubstituted flavins"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002400_1.4006324-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002400_1.4006324-Figure2-1.png",
        "caption": "Fig. 2 Thin-rimmed spur gears with inclined webs and the mating gear",
        "texts": [
            " This face-contact model is in agreement with the real contact state of teeth with deformations; therefore, this model can analyze the real surface contact stress as well as the root bending stress of the contact teeth, especially in the case where the gears have machining errors, assembly errors, and tooth modifications. It has already been proven by experiments and ISO standards that this model and the developed FEM software can perform exact analyses of tooth surface contact stresses and root bending stresses of spur gears [11\u201315]. This model and the FEM are also used here for the LTCA, deformation, and stress calculations of thinrimmed inclined web gears. 3.1 Structures and Parameters of the Gears. Three types of inclined web gears and a solid mating gear as shown in Fig. 2 are used as research objects in this paper. Figures 2(a)\u20132(c) are Contributed by the Power Transmission and Gearing Committee of ASME for publication in the JOURNAL OF MECHANICAL DESIGN. Manuscript received July 8, 2011; final manuscript received February 27, 2012; published online April 5, 2012. Assoc. Editor: Professor Philippe Velex. Journal of Mechanical Design MAY 2012, Vol. 134 / 051001-1Copyright VC 2012 by ASME Downloaded From: http://mechanicaldesign.asmedigitalcollection.asme.org/ on 01/28/2016 Terms of Use: http://www",
            "org/about-asme/terms-of-use the thin-rimmed spur gears with inclined webs on the left side of the tooth, the center of the tooth, and the right side of the tooth, and they are called the left web gear, the center web gear, and the right web gear, respectively. In Sec. 4, they are also called the left inclined web gear, the center inclined web gear, and the right inclined web gear to make a distinction from the thin-rimmed gears with straight webs. The web inclination angle (simply called web angle) is denoted by h as shown in Figs. 2(a)\u20132(c). When h\u00bc 0 deg, this indicates that the web is a straight web. Figure 2(d) is a solid gear used as the mating gear of these gears when they are engaged. The tooth numbers, the modules, the pressure angles, and the shifting coefficients of all the gears in Fig. 2 are denoted by Z1\u00bc Z2\u00bcZ3\u00bc Z4\u00bc 50, m\u00bc 4, a\u00bc 20 deg, and X1\u00bcX2\u00bcX3 \u00bcX4\u00bc 0, respectively. The structural dimensions of these gears are also shown in Fig. 2. 3.2 FEM Models and Boundary Conditions. Figure 3(a) is used to define the angle A that will be used in Sec. 6. Figure 3(b) is the FEM model used for the LTCA, deformation, and stress calculations of the thin-rimmed inclined web gears when they are engaged with the mating gear as shown in Fig. 2(d). Because the mathematical programming method [11\u201315] is used for the LTCA in this paper, it is only necessary to calculate the deformation influence coefficients and the gaps of the assumed contact point pairs on the contact tooth surfaces when conducting LTCA with the FEM. The simple procedure is as follows. First, a mathematical model used for LTCA is developed based on the principle of the mathematical programming method. Then, the deformation influence coefficients and the gaps of the assumed contact point pairs on the contact tooth surfaces are calculated with the FEM, and the models shown in Fig",
            " It has been previously reported that this method of developing the FEM models and boundary conditions can produce suitable calculation results by experiments through comparing the calculated tooth root strain of the thinrimmed gear with a straight web with the measured strain [14]. The FEM models shown in Fig. 3(b) can be produced automatically for all of the gears and calculations in this paper with the software developed through the efforts of many years. The gearing parameters, the structural parameters, as well as the web angle can be changed freely in FEM modeling. LTCA is conducted for the three types of thin-rimmed gears shown in Fig. 2, when these gears have both straight webs and inclined webs, and they are engaged with the solid mating gear at the highest point of the single pair tooth contact. The tooth contact stresses are calculated under a torque load of 294 N m for all of the cases in the paper. 4.1 Tooth Contact Stress Distributions of the Thin-Rimmed Left Web Gears. At the time when the webs are located at the left side of the tooth, the tooth contact stresses are calculated. Figure 4 shows the results of the tooth contact stresses of the left web gears",
            "org/about-asme/terms-of-use complicated than the statements mentioned above. It is necessary to consider the web deformation on the loaded tooth to understand the root stress of thin-rimmed gears with an inclined web. 7.2 Deformation-Sharing Ratios of the Tooth, the Rim, and the Web. A tooth\u2019s relative deformation of a pair of contact gears along the line of action can be calculated by LTCA under the application of a torque load. Table 4 shows the results. In Table 4, the tooth\u2019s relative deformation of a pair of solid gears (the mating gears in Fig. 2(d)) is calculated at the worst load position, and the result is shown in the first column (denoted by \u201csolid gears\u201d). The relative tooth deformations of thin-rimmed left web gears with different web angles are also calculated when the left web gears are engaged with the solid mating gear. The calculation results are shown in the columns denoted by \u201cThin-rimmed left web gear\u201d in Table 4. In Table 4, the deformations are divided into the gear deformation (the total deformation of a pair of gears), the tooth deformation (the deformation resulting only from the contact teeth), and the rim and web deformation (the deformation resulting only from the rim and the web)",
            "017 deg of misalignment errors on the plane of action of the gears, the contact stress distribution pattern can again be calculated under the condition of a misalignment error by LTCA [12,15]. Figure 25 illustrates the result and shows that the partial tooth contact occurred at the right end of the tooth because of the misalignment error. It was also found that the maximum contact stress of the teeth increased from 500 MPa to 550 MPa by comparing Fig. 25 with Fig. 6(a). To change the partial tooth contact pattern to a uniform tooth contact pattern, the thin-rimmed right inclined web gear, as shown in Fig. 2(c), with a 30-deg web angle can be used. LTCA was conducted for this gear under the condition of a 0.017-deg misalignment error with the same torque load. The calculation result of the contact stress pattern is given in Fig. 26. From Fig. 26, it was found that the uniform tooth contact pattern was obtained by using the thin-rimmed right inclined web gear even though the gear contained a 0.017-deg misalignment error. The web position and the web angle of the thin-rimmed gears have significant effects on the tooth contact stresses, the root bending stresses, and the joint stresses of the gears"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001167_1.2991128-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001167_1.2991128-Figure1-1.png",
        "caption": "Fig. 1 Geometry of journal bearings",
        "texts": [
            " When the amount of supply oil is reduced o Q=1.8 10\u22126 m3 /s, the journal motion show a gradual transi- ion to stability, as shown in Fig. 4 ii , and cavitation generation ournal of Tribology om: http://tribology.asmedigitalcollection.asme.org/ on 09/01/2017 Terms and evaporation occur alternatively with increasing in rotational speed. With further reducing the amount of supply oil to Q=0.5 10\u22126 m3 /s, the journal becomes stable and the cavitation can be observed for higher range of rotational speed, as shown in Fig. 1 iii . The established state is used to define the amount of oil under starved lubrication conditions. The same manner is applied to determine the amount of supply oil for elliptical bearing. Figure 5 shows the experimental results for circular and elliptical journal bearings in the case of bearing orientation angle of =0 deg i under flooded lubrication conditions and ii under starved lubrication conditions. Figure 5 a shows the relationship between the rotational speed, frequency, and amplitude, and Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002039_pime_conf_1968_183_282_02-Figure11.2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002039_pime_conf_1968_183_282_02-Figure11.2-1.png",
        "caption": "Fig. 11.2. Geometry of conjunction between two discs",
        "texts": [
            " Several other workers (7)-(9) have confirmed that steel discs with equal and opposite peripheral velocities will carry only a very light load before scuffing, whereas discs made of two dissimilar metals will carry quite heavy loads. During the course of the present work, loads up to 7.12 kN (1600 lbf) have been carried on two steel discs of approximately equal bulk temperatures, without marking or scuffing of the surfaces. It is thought that this was possible only because of the very good surface finish used. Explanation in principle of viscosity wedge effect A simple physical description of the viscosity wedge effect in the present work is illustrated in Fig. 11.2. It is assumed that the discs are made of the same material, and that their peripheral velocities are equal and opposite. The surface of the lower disc 2 is leaving the conjunction, and has therefore been heated by its passage through the source of frictional heat. A particle of lubricant at P, on AIAz near the surface of disc 1 will be cooler than is a corresponding particle P, near the surface of disc 2, the lengths Alpl and P,A2 being equal. The viscosity of the lubricant at P, will be greater, and the velocity gradient across the particle will be less",
            " This would be a very complicated task, and in the present state of knowledge it is not possible to arrive in this way at realistic values of the frictional traction, and the validity of the results would therefore be in considerable doubt. As an alternative approach it will be shown that it is possible by simple arguments to define theoretically two extreme conditions between which the real situation may be expected to lie in any given case. Theoretical solutions corresponding to these extreme conditions may then be compared with these solutions. MODIFICATION TO ELASTOHYDRODYNAMIC THEORY Description of the system The system is illustrated in Fig. 11.2. Disc 1 is taken to be the faster of the two discs if there is any difference in rotational speed, and the direction of its peripheral velocity, U,, is taken to be along the positive x direction. The peripheral velocity, U,, of disc 2, measured in the same direction, is assumed to be negative, so that u, > 0 r;, < 0 u,+u, > 0 u,-u2 > 0 Vol383 Pt 3P at HOWARD UNIV UNDERGRAD LIBRARY on April 27, 2016pcp.sagepub.comDownloaded from The direction of the y-axis is taken to be perpendicular to the direction of motion of the surfaces and normal to the surfaces, which are assumed to be nearly parallel",
            " Temperatures of the disc surfaces The temperatures of the surfaces will be taken as constant for each disc over a relatively narrow region near the inlet edge of the Hertzian flat. The inlet side is defined as that on which the faster surface is entering the conjunction, and it is over this region that those physical processes take place that are important in governing the minimum film thickness. The rate of generation of heat by viscous shear in this zone is assumed to be so small that the temperatures and viscosities of the oil are not affected. It is assumed that all the heat dissipated in friction arises in the Hertzian contact zone, C1C2D2D1, Fig. 11.2, where the pressures are higher. In any section, A,A,, of the inlet zone, the surface of disc 2 is leaving the zone of Hertzian contact and its temperature, 8,, is therefore higher than the temperature, 8,, of the surface of disc 1, which has yet to enter the contact zone. It will be assumed that the temperatures of the two surfaces become equal at the inlet edge, C,C,, of the Hertzian zone, the exact common value of this temperature being immaterial. This equality of the temperatures of the surfaces at the inlet edge of the Hertzian zone is required for consistency with theories of the film thickness in elastohydrodynamic lubrication of the type originated by Grubin (IO), in which the pressure gradient, dp,/dx, is assumed to vanish at this point",
            "5, has been given by Cameron Results for stepwise temperature distribution The general case of a stepwise temperature distribution with two arbitrary levels, 0, and 6\u2019,, is difficult to deal withThe symbol B\u2018, is used instead of O , to allow for the possibility that the temperature of most of the oil leaving the Hertzian contact zone may be higher than the temperature of the surface of disc 2. But it is easy to work out the extreme case where U\u2019, is assumed to be so high that the corresponding viscosity T \u2019 ~ is small compared with In the limit, when T \u2019 ~ tends to zero, it can be shown that the thickness of the part of the oil coming from the contact zone also tends to zero, i.e. the point S of zero velocity in Fig. 11.2 tends to coincide with A2. The existence of a non-zero shear stress gradient, dpldx, implies also the existence of a non-zero shear stress gradient, dT/dy. The shear stress cannot vanish everywhere over a non-zero thickness SAz; therefore, if the viscosity tends to zero, the shear rate tends towards infinity. This condition cannot be permitted over a non-zero thickness SA,, since there would then be an infinite volume rate of flow Q2 (Appendix 11.6) across the section SA,. Therefore the height of the section SA, must tend to zero",
            "1 for the viscosity of lubricant A at atmospheric pressure gives the solution: Vo1183 Pt 3P at HOWARD UNIV UNDERGRAD LIBRARY on April 27, 2016pcp.sagepub.comDownloaded from Temperature, \"C . 40 n . . -I? - 60 80 100 120 3.47 3.11 2.77 255 _ _ _ _ ~ - - _ of the momentum equation to an elementary volume of lubricant then gives the usual relation: d.r dp -- - _ dY dx or APPENDIX 11.5 F I L M T H I C K N E S S I N E L A S T O H Y D R O D Y N A M I C L U B R I C A T I O N A T H I G H S L I D E / R O L L R A T I O S F O R A L I N E A R T E M P E R A T U R E D I S T R I B U T I O N Description of the system In Fig. 11.2, the surfaces AICIDIBl and A2C2D2B2 of the two discs have shapes identical with those of the Hertzian case of dry contact, but with the addition of a constant separation, h,, at all positions. The peripheral velocities of the discs are U , and U,, and it is assumed that surface 1 is the faster surface, that the velocity of this surface is directed in the positive x direction, and that the velocity of the other surface is directed in the negative x direction. Thus u1 > 0 u2 < 0 U,+UZ 2 0 u,-u2 > 0 The idealized temperature distribution that is assumed to exist has already been described in the text",
            "39), may therefore be obtained from conventional approximate isothermal theories, of the type proposed by Grubin, by replacing the term ( U1 + U,) in those theories by the expression ( ul + U,)fdv, 4 + 3( ul - UJfdv, 4 It therefore seems appropriate to obtain a closer approximation by making the same substitution in the empirical result of Dowson and co-workers, equation (1 1.1). In this equation, the velocity term used is 0 = +(U,+ U,) and this must then be replaced by f( Ul, Vz) = 4( u, + ~ Z > . f l ( ~ . , 4 +3 u1- UJfz(v, 4 which is the expression quoted in the text. Proc lnsrn Mech Engrs 1968-69 APPENDIX 11.6 FILM T H I C K N E S S I N E L A S T O H Y D R O D Y N A M I C L U B R I C A T I O N AT H I G H S L I D E / R O L L RATIOS F O R At some point S in the section AIA, in Fig. 11.2 the velocity u of the oil must be zero relative to the conjunction between the discs. Let this point be the origin of the y co-ordinates. A S T B P W I S E TEMPERATURE D I S T R I B U T I O N Then fromy = 0 toy = h,, rl = 7 1 and from y = -hz t o y = 0, where 7 = 7'2 h,+h, = h h, 2 0 h2 2 0 Equation (1 1. 18) remains valid in both regions with the same value of C, in both, since there can be no discontinuity in the shear stress. Equations (11.18) and (11.19) give and integration with respect t oy gives u = 7-1$ g + C l y ) "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000597_978-3-540-73129-0-Figure6.76-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000597_978-3-540-73129-0-Figure6.76-1.png",
        "caption": "Fig. 6.76. Two modes of MR fluid operation: a valve-mode, b direct-shear mode",
        "texts": [
            " The best MR fluids today can sustain a LDE on the order of 107 J/cm3 before they thicken to the point where device performance is compromised. Poor MR fluids, on the other hand, may become unusable with LDEs as low as 105 J/cm3. Today, good MR fluids are capable of lasting hundreds of thousands of kilometers in automotive shock absorbers. Virtually all devices that use controllable MR fluids operate in a valvemode, direct-shear mode, or a combination of these two modes. Diagrams of the basic valve and direct-shear modes are shown in Fig. 6.76. Examples of valve-mode devices include dampers, and shock absorbers. Examples of direct shear-mode devices include clutches, brakes, chucking and locking devices, and some dampers. The pressure drop developed by a valve-mode device can be divided into two components, the pressure \u0394P\u03b7 due to the fluid viscosity and \u0394PMR due to the magnetic field-induced yield. These pressures may be approximated by [131,157,158]: \u0394P\u03b7 = 12\u03b7pQL h3w (6.34) \u0394PMR = c\u03c4MR(H)L h , (6.35) where Q is the volumetric flow rate"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000726_s11044-007-9072-4-Figure3-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000726_s11044-007-9072-4-Figure3-1.png",
        "caption": "Fig. 3 A hub-beams system with spatial motion",
        "texts": [
            " (48) The virtual power of the elastic force is given by \u03b4Wf = n\u2211 e=1 \u03b4p\u0307T e Qe = \u2212 n\u2211 e=1 \u03b4 p\u0307T e Kepe = \u2212\u03b4p\u0307T Kp, (49) where Ke = K l + Kb + K t represents the element elastic stiffness, and K represents the global elastic stiffness, which takes the form K = n\u2211 e=1 BT e KeBe. (50) The system elastic energy reads Uf = n\u2211 e=1 (Uel + Ueb + Uet ). (51) A hub-beams system undergoing three-dimensional large overall motion is composed of a rigid hub Bc , and two flexible beams, B1 and B2. As shown in Fig. 3. The length, width and height of the cubic hub are 2a, respectively. B1 is attached to the hub at the point O1 by a fixed joint, and B2 is connected to B1 at the point O2 by a spherical joint. Four coordinate systems are introduced to describe the system motion: The inertial frame O0\u2013X0Y0Z0, the body-fixed frame Oc\u2013XcYcZc of the hub with its origin located at the centroid of the hub, and the body-fixed frame O1\u2013Xb1Yb1Zb1 and O2\u2013Xb2Yb2Zb2. Let ec = [ ic jc kc]T , eb1 = [ ib1 jb1 kb1]T and eb2 = [ ib2 jb2 kb2]T be columns of the unit basic vectors of Oc\u2013XcYcZc,O1\u2013Xb1Yb1Zb1 and O2\u2013Xb2Yb2Zb2, respectively",
            " rc = eT c rc is the position vectors of the reference point Oc with respect to O0\u2013X0Y0Z0, and d = eT c d is the relative position vector of O1 with respect to Oc\u2013XcYcZc , and \u03c1O2 = eT b1\u03c1O2 is the relative position vector of O2 with respect to O1\u2013Xb1Yb1Zb1. d is given by d = [a 0 0]T . Let Ac,A1 and A2 be the transformation matrices of Oc\u2013XcYcZc,O1\u2013Xb1Yb1Zb1 and O2\u2013Xb2Yb2Zb2 with respect to O0\u2013X0Y0Z0. The system constraint equations are given by A1 = Ac, \u03981 = \u0398c, (52) r01 = rc + Acd, r02 = r01 + A1(\u03c1O2 + \u03beO2), (53) where \u0398c = [\u03b1c \u03b2c \u03b3c]T represents Tait\u2013Brian angle of Oc\u2013XcYcZc with respect to O0\u2013 X0Y0Z0. As shown in Fig. 3, O2 is on the neutral axis of B1, thus, \u03beO2 = 0, \u03c1O2 = SO2p1, where SO2 = S1(x\u0304 = le). Application of velocity variational principle yields the equations of motion 2\u2211 i=1 [\u222b V \u03b4r\u0307T i (\u2212\u03c1r\u0308 i + f i )dV + \u03b4Wif ] + \u03b4r\u0307T c (\u2212mc r\u0308c + F c) + \u03b4\u03c9T (\u2212J c\u03c9\u0307 \u2212 \u03c9\u0303J c\u03c9 + Mc) = 0, (54) where \u03b4Wif = \u2212\u03b4p\u0307T i K ipi is the virtual power of the elastic force of beam Bi , and f i = [f1i f2i f3i]T is the body force vector exerted on beam Bi , which is defined in the inertial frame. mc is the mass of the hub, and J c is the rotary inertia of the hub with respect to the body-fixed frame of the hub, and F c is the sum of the external forces defined in the inertial frame and Mc is the sum of the external torques with respect to Oc defined in the body-fixed frame of the hub, respectively"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000427_j.ijsolstr.2007.12.004-Figure7-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000427_j.ijsolstr.2007.12.004-Figure7-1.png",
        "caption": "Fig. 7. (a) Contact\u2013contact sub-domain. (b) Free body diagrams.",
        "texts": [
            " (14) can also be modified to M2 H 2 \u00bc QA sin w 2 w\u00feH 2 sin w 2 hN1 \u00fe F A cos w 2 w\u00feH 2 cos w 2 hN1 \u00f024\u00de In the case when the contact becomes distributed, the shear force QA can be calculated as QA \u00bc 1 2 F A cos w 2 w\u00fe H 2 cos w 2 hN 1 h i sin w 2 w\u00fe H 2 sin w 2 hN1 h i \u00f025\u00de If the contact occurs at point J on the inner wall, Eqs. (21) and (25) remain the same except that subscript \u2018\u2018H\u2019\u2019 is replaced by \u2018\u2018J,\u2019\u2019 and the superscript \u2018\u2018+\u2019\u2019 is replaced by \u2018\u2018 .\u2019\u2019 The equations formulated in this subsection can be used to determine the deflection curves involving clamped\u2013contact sub-domain, such as in Fig. 3(2) and (3). There are two types of contact\u2013contact sub-domains. The first case is when both the contact points are on the outer wall, as shown by points H and I in Fig. 7(a). The second case is when the sub-domain contacts the outer wall at one end and contacts the inner wall at the other end, as shown by points H and J in Fig. 7(a). The locations of points H, I, and J are signified by angles w\u00fe H 2 , w\u00fe I 2 , and w J 2 . These two scenarios are listed as cases (d) and (e) in Table 1. In our experiment, we never encounter the case when both the contact points are on the inner wall. Therefore, this scenario is omitted in this paper. Fig. 7(b) shows the free body diagrams of these two sub-domains. We establish an xy-coordinate system with origin at point H. PH, MH, and QH(l) are the longitudinal force, bending moment and shear force at point H. The subscript \u2018\u2018l\u2019\u2019 in QH(l) is to emphasize that the shear force is at a location slightly to the left of the contact point H. The rotation angle h at H is equal to w\u00fe H 2 . Following the same formulation as in Section 4.2, we can derive the equation 1 2 dh ds 2 \u00bc QH\u00f0l\u00de sin w\u00feH 2 h \u00fe P H cos w\u00feH 2 h \u00fe DH \u00f026\u00de DH is a new integration constant"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003882_9781118181249-Figure3.2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003882_9781118181249-Figure3.2-1.png",
        "caption": "Figure 3.2 Schematic illustration of the different molecular components required to create a biosensor surface. (See text for full caption.)",
        "texts": [
            " The requirement to block all surfaces to nonspecifi c binding is common, but the challenge varies with respect to how diffi cult this is to achieve in different applications. When considering the creation of the prerequisite nonfouling interface, it is instructive to divide the design of the molecular interface into three mostly independently designable parts: surface anchor, spacer, and functional unit (recognition element). The role and strategy of choice for each component will be discussed next, with particular focus on the design criteria required to generate a biofunctional pattern (see Fig. 3.2 ). ORTHOGONAL SMALL (NANO)-SCALE SURFACE MODIFICATION 77 The surface anchor is perhaps the most decisive choice to create a proper interface for both macroscopic and nanoscale biosensors. The choice of anchor strategy determines the stability of the molecular interface, for example, how strongly attached the molecules will be, under the various conditions the sensor might be subject to, if they enable ordered transitions at the interface coupled to the sensor, and/or how mobile the functional units will be"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003294_s1004-9541(13)60563-7-Figure5-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003294_s1004-9541(13)60563-7-Figure5-1.png",
        "caption": "Figure 5 Response surface and contour diagrams of CAPE selectivity in [Emim][Tf2N] as a function of: (a) X1and X2 interaction on CAPE selectivity, (b) X1 and X3 interaction on CAPE selectivity, (c) X2 and X3 interaction on CAPE selectivity. X1: reaction temperature (\u00b0C); X2: mass ratio of Novozym 435 CA (g\u00b7g 1); X3: molar ratio of PE CA (mol\u00b7mol 1)",
        "texts": [
            " The ANOVA for the reactive selectivity of CAPE synthesis using the response surface quadratic model suggests that the F-value of the model is 30.83 and the P-value is less than 0.0001. These data implies that the model is highly significant. To determine the most adequate operating conditions and analyze the improvement process of the reactive selectivity of CAPE synthesis, the response surface is plotted using Eq. (5) for three possible combinations. The response surface and contour diagrams of CAPE selectivity as a function of (a) X1 and X2, (b) X1 and X3, (c) X2 and X3 are presented in Fig. 5. The simultaneous analysis of these plots is complex if practical short cuts that take advantage of prior knowledge of the process are not adopted [24]. In Fig. 5, the main factors affecting on CAPE selectivity are reaction temperature, the mass ratio of Novozym 435 to CA and the molar ratio of PE to CA. For example, in Fig. 5 (a), low temperatures and the mass ratio of Novozym 435 to CA during the esterification process lead to low reactive selectivity of CAPE synthesis. Therefore, reaction temperature and the ratio of enzyme to caffeic acid can be selected at an appropriate value range. The optimum value is found by solving the regression equation analytically. The solution is obtained by submitting the level of the factors into Eq. (5). The optimal reaction conditions for the improvement of the reactive selectivity of CAPE synthesis are calculated as follows: reaction temperature, 84"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000427_j.ijsolstr.2007.12.004-Figure8-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000427_j.ijsolstr.2007.12.004-Figure8-1.png",
        "caption": "Fig. 8. Free body diagram in a distributed contact sub-domain.",
        "texts": [
            " (23) we have the relation M2 H 2 \u00bc P H \u00fe DH \u00f029\u00de In the special case when MH = 1 for distributed contact, DH can be calculated as DH \u00bc 1 2 P H \u00f030\u00de The analysis for case (e) in Table 1 is similar. The equations formulated in this sub-section can be used to determine the deflection curves involving contact\u2013contact sub-domain, such as in Fig. 3(4)\u2013(12). In some cases the elastica is in full contact with the outer radius over a finite segment, as listed in Table 1(f). We choose an arbitrary element HI from the fully contact segment and show the free-body diagram in Fig. 8. It can be shown that the distributed force q is a constant and is equal to PH. First of all the moment balance equation about point I can be written as 1 \u00bc 1\u00fe P H sin w\u00feH 2 sin w\u00feH 2 sin w\u00feI 2 P H cos w\u00feH 2 cos w\u00feI 2 cos w\u00feH 2 Z w\u00fe H 2 w\u00fe I 2 q cos h sin h sin w\u00feI 2 dh Z w\u00fe H 2 w\u00fe I 2 q sin h cos w\u00feI 2 cos h dh \u00f031\u00de After integration, Eq. (31) can be simplified to q = PH. This sub-domain can be found in some experimental observations, such as in Fig. 3(3), (5)\u2013(6), and (8)\u2013(12). Each deformation described in the experimental observations in Section 3 can be divided into sub-domains as listed in Section 4 and solved theoretically"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000597_978-3-540-73129-0-Figure6.127-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000597_978-3-540-73129-0-Figure6.127-1.png",
        "caption": "Fig. 6.127. Construction of an electrostatic linear actuator based on microtechnologies. (Source: PASIM Mikrosystemtechnik, Suhl in Germany)",
        "texts": [
            " The motor with integrated gear box increases the available torque and helps to open new fields of application \u2013 for example, communication and information technology as well as consumer electronics. Hybrid concepts make use of the most suitable material and the most appropriate process in the fabrication of each component. Such a heteromorphic construction is typical of many microsystems and also demonstrates a broad need for efficient construction, connection and microassembly techniques, and standardized electrical and mechanical interfaces. In the linear actuator depicted in Fig. 6.127, the slide moves over the stator supported by air. Electrostatic forces are generated between comb-like or striped electrodes located on the opposing surfaces of the stator and slide, causing motion of the slide along the x-axis, binding in the y direction and attraction in the z direction. Additional electrodes act as sensors for determining the position in the x direction and the distance of separation in the z direction. All structures are sputtered onto a glass substrate using conventional methods"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002947_17452759.2013.790599-Figure3-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002947_17452759.2013.790599-Figure3-1.png",
        "caption": "Figure 3. Scheme of the main dimension of the specimens used to identify the adhesion between layers and the direction of applied force.",
        "texts": [
            " The control factors were defined as the velocity of deposition (mm/s), the velocity of extrusion (mm/s) and the extrusion temperature (8C), and the bead height (mm) was held constant. To analyse the response factors, we defined a proof part whose geometry would allow the investigation of the influence of the process parameters for small features, finishing or dimensional accuracy, and non-supported features. Figure 2 shows the main dimensions of the proof part with the indications of the features to be analysed. Figure 3 shows the main dimension of the specimens that were used to identify the adherence between layers. This figure additionally shows the direction of the force that was applied to determine the ultimate stress of the specimens. It is important to note that the method used to determine the adherence between layers is similar to other studies, such as Ahn et al. (2002) and Sood et al. (2010). However, the use of a non-standardised specimen intends to evidence the layer adhesion. With regard to the range of parameter values that were studied, Table 2 presents the design matrix of the full experimental procedure with neither central nor faced points"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001901_1.3591479-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001901_1.3591479-Figure2-1.png",
        "caption": "Fig. 2 Spat ia l three-link mechanism",
        "texts": [
            " C - l - C - 6 ) ; numerical data for these figures is also given. Spatial Three-L ink Motion While the number of spatial three-link mechanisms is quite large [8, 14, 25], consideration of just one (other than the preceding case) will suffice to illustrate typical characteristics of this motion. We consider a three-link mechanism, which has been used in a stalor-winding machine and electric toothbrushes [32] as well as in washing-machine agitators [18]. Schematically, the mechanism is illustrated in Fig. 2. A turning joint with fixed axis CB, carries a crank BA = a, the center of the crank circle being at B. The coupling at A is a centrally slotted spherical joint, permitting the output link APQ'Q to slide through the slotted sphere and to swivel about it. The output motion is governed by the cylindrical joint at P, which permits both sliding along and rotation about the x-axis. Point Q is an arbitrary point on the output link. The common perpendicular between input and output axes is OC = c. Point 0 is the origin, OC the z-axis and OP the i-axis of a right-handed xyz coordinate system"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000893_eurcon.2007.4400406-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000893_eurcon.2007.4400406-Figure1-1.png",
        "caption": "Fig. 1. Rotor cage equivalent circuit",
        "texts": [
            " MODEL OF THE THREE-PHASE INDUCTION MOTOR The model of the induction motor takes into account the following assumptions [11]: - the saturation and the skin effect are negligible, - the air-gap is uniform, - the mmf distribution in the air-gap is sinusoidal, - the rotor bars are isolated from the magnetic circuit of the rotor, - the relative permeability of the magnetic circuit is supposed to be infinite. Although the mmf of the stator windings is supposed to be sinusoidal, other distributions of rolling up could also be considered by simply employing the superposition theorem. It is justified by the fact that the different components of the space harmonics do not act the ones on the others [12]. In order to study the phenomena taking place in the rotor, the latter is often modeled by NR meshes as shown on figure 1. The induction motor mathematical model can be written as follows: [ ] [ ] [ ] [ ][ ])I.L( dt d IRV += with: [ ] [ ] [ ]         = R S V V V , [ ] [ ] [ ]         = R S I I I and: [ ] [ ]t 3S2S1SS VVVV = , [ ] [ ] t )1N(1R R 0000V +\u00d7= [ ] [ ] t 3S2S1SS IIII = [ ] [ ] t )1N(1eRNRk1RR RR II..I..II +\u00d7= The resistance global matrix can be written as: [ ] [ ] [ ] [ ] [ ] [ ] [ ] [ ] [ ] [ ]                   \u2212 \u2212= \u00d7\u00d7\u00d7 \u00d7\u00d7\u00d7 \u00d7\u00d7\u00d7 11eN1 R e 31 1N e NNR3N 13N333S 1R1 N R0 1 N RR0 00R R R R R RRR R where: [ ]           = \u00d7 S S S 33S R00 0R0 00R R The global inductance matrix can be represented by: [ ] [ ] [ ] [ ] [ ] [ ] [ ] [ ] [ ] [ ]                     \u2212 \u2212= \u00d7\u00d7\u00d7 \u00d7\u00d7\u00d7 \u00d7\u00d7\u00d7 11eN1 R e 31 1N R e NNR3NRS 13N3SR33S 1L1 N L 0 1 N L LM 0ML L R RRRR R where: [ ]             \u2212+\u03b8\u2212 \u2212+\u03b8\u2212 +\u03b8\u2212 = \u03c0 \u03c0 "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000199_2006-01-0427-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000199_2006-01-0427-Figure1-1.png",
        "caption": "Figure 1. Single cylinder research engine",
        "texts": [
            " Many research and development projects have focused on the reduction of piston friction or NVH, but rarely together in the same activity. It is thought to be extremely important to control both aspects with balance since the behavior of each is so closely linked. For this study a single cylinder gasoline engine equipped with a real time piston friction measurement device (floating liner) was used. For the detection of piston skirt slap noise, a piezo electric accelerometer was attached to the side of the engine block. The single cylinder research engine used is shown in Figure 1. Table 1 and 2 show the main specifications of the engine and the piston respectively. Figure 2 shows the geometry of the piston rings. Dermot Madden and Kwangsoo Kim Copyright \u00a9 2006 SAE International Table 2. Piston Specifications Piston alloy Eutectic base alloy Pin offset 0.5 mm to thrust side Skirt roughness 18Rz Skirt treatment No coating Skirt ovality 0.65 mm Piston length 53.9 mm Figure 2. Geometry of piston rings During the tests, the same piston ring and cylinder were used after full break-in"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002055_s10811-009-9401-5-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002055_s10811-009-9401-5-Figure1-1.png",
        "caption": "Fig. 1 Schematic diagram of the photosynthetic activity measurement system. 1 Reaction cell, 2 magnetic bar, 3 cooling water jacket, 4 dissolved oxygen electrode, 5 inlet of cooling water, 6 outlet of cooling water, 7 inlet of sample, 8 outlet of sample, 9 convex lens, 10 quantum sensor, 11 waste water, 12 quartz halogen illuminator, 13 water bath, 14 peristaltic pump, 15 sample reservoir, 16 DO meter, 17 computer, 18 magnetic stirrer",
        "texts": [
            " In each experiment, the volumetric oxygen evolution rate was measured from the linear slope between dissolved oxygen (DO) and time. The specific oxygen evolution rate was calculated by dividing the volumetric oxygen evolution rate by the cell concentration (Jeon et al. 2005). Two replicates were conducted for each toxicant but three replicates for controls, because the control is the basis for evaluation of toxicity and if the control is not precise the EC50 may contain a large error. To obtain EC50 values, 7\u201310 levels of dosage were tested for each toxicant. Construction of the experimental device Figure 1 shows a schematic diagram of the toxicity estimation system, which is reproduced from our previous article (Cho et al. 2008b). The reaction vessel (1) was a double-jacket cylinder made of Pyrex\u00ae glass, which can remove the scattering effect. The working volume and light path length of the reaction cell were 3.58 mL and 1.8 cm, respectively. A small bar magnetic stirrer (0.5 cm in length) (2) was placed inside the vessel to make oxygen partial pressure uniform over the whole medium in the chamber as well as allow the microalgal suspensions to become homogeneous"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000388_14644193jmbd97-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000388_14644193jmbd97-Figure2-1.png",
        "caption": "Fig. 2 Loci of raceway groove curvature radii centres before loading",
        "texts": [
            " The objective here is to calculate the net force on a bearing due to the displacement of the spindle centre since this force can then be used in the equations of motion to observe the motion of the spindle centre. In order to calculate the total force, the deflection at the ith ball in Fig. 1 will be calculated first and this will be used in the calculation of the total force. As seen in Fig. 1, the ball is rotating between the inner and outer rings. During this rotation, the ball is continuously in contact with different points in the circular grooves in each race. In the initial position, without any preload, the loci of raceway groove centres of curvature will produce circles as shown in Fig. 2. The figure is a three-dimensional representation of the inner and outer ring raceway groove curvature loci, with two-dimensional crosssection of an inner and outer race overlaid on it. Proc. IMechE Vol. 222 Part K: J. Multi-body Dynamics JMBD97 # IMechE 2008 at UNIV PRINCE EDWARD ISLAND on August 5, 2015pik.sagepub.comDownloaded from If a ball is compressed by a force, for example the weight of the spindle, since the centres of the curvature of the raceway grooves are fixed with respect to the corresponding raceway, the distance between the centres is increased in proportion to the amount of the normal approach between raceways. It can be determined from Fig. 2 that the loci of the centres of the inner and outer ring raceway groove curvature radii are, respectively, expressed by <i \u00bc 1 2 dm \u00fe rgi db 2 cos\u00f0a0\u00de \u00f07\u00de <o \u00bc <i Bdb cos\u00f0a0\u00de \u00f08\u00de or <o \u00bc dm 2 \u00fe rgi \u00fe db 2 cos\u00f0a0\u00de \u00f09\u00de If the outer ring of the bearing is considered fixed in space as load is applied to the bearing, then the inner ring will be displaced. It can be shown that the distance between the centres of curvature of the JMBD97 # IMechE 2008 Proc. IMechE Vol. 222 Part K: J. Multi-body Dynamics at UNIV PRINCE EDWARD ISLAND on August 5, 2015pik"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002764_speedam.2010.5544902-Figure3-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002764_speedam.2010.5544902-Figure3-1.png",
        "caption": "Fig. 3. Vibration model",
        "texts": [
            " stat m,m cc (10) This dynamic magnetic spring constant cm can be subsequently implemented in the rotor dynamic model to take into consideration this dynamic magnetic force. This approach of linearization is of course only valid for small static eccentricities ( < 0.2) and small movements. If no homopolar flux would occur in the two-pole induction machine, no oscillating magnetic force would occur and the magnetic spring constant would be half [2, 3]: 0r F and 2 p 00 mstat m, \u02c6 4 BlRcc (11) The rotor dynamic model is described in Fig. 3. This model is a simplified rotor model based on the LAVAL rotor [13], taking in consideration electromagnetic effects [1, 10]. In this paper the model is now advanced for considering excitation by an oscillating magnetic force Fr. The total rotor mass m is concentrated as a lumped mass on an elastic shaft with rotor stiffness c symmetrically arranged between two identical sleeve bearings. The moment of inertia is not taken into consideration \u2013 and there is therefore neither lateral buckling nor any gyroscopic effects",
            " As a worst case calculation \u2013 without considering field damping effects, saturation, influence of the slots and assuming maximum homopolar flux (section III) \u2013 a maximum amplitude of the radial magnetic force is computed for: e = 225 m ( = 0.05) Simulating a constant magnetization over the whole speed range, the amplitude of the magnetic force can be assumed to be constant for different rotor speeds. The oscillating magnetic force excites the rotor to vibrate, leading to orbital movements of shaft centre point W and of shaft journal points V*. These orbits can be computed for different rotor speeds and for different directions of the magnetic forces, which is described by the angle F (Fig. 3). If F = 0\u00b0 the magnetic force Fr only acts in vertical direction (z-direction). If F = 90\u00b0 the magnetic force Fr is only acting in horizontal direction (ydirection). To show the orbits of the shaft centre W and of the journal shaft V* the semi-major axes, the semi-minor axes and the angles of the semi-major axes are computed. By changing the angle F of the magnetic force, which means here changing the position of the static eccentricity, also the direction of the static magnetic pull will change"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000715_s11249-007-9287-9-Figure14-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000715_s11249-007-9287-9-Figure14-1.png",
        "caption": "Fig. 14 Shape parameters used in the calculation",
        "texts": [
            " These results show that, in addition to the groove depth and pitch, the radius of curvature of the convex portion is also an important parameter that should be controlled for improving the traction coefficient while suppressing metal-to-metal contact. 4.4 Effect of Radius of Curvature of Convex Portion on Anti-flaking Performance The effect of the radius of curvature of the convex portion and the longitudinal groove depth on anti-flaking performance was investigated by elastic contact analysis using the conditions shown in Table 5. The longitudinal grooves were modeled in the calculation as shown by the curves in Fig. 14. The groove pitch was kept constant at 200 lm, R2 at 200 lm, R3 at 50 lm, and the groove depth at 2.7 lm, and the radius of curvature R equaled R1. Figure 15 shows the ratio of the contact stress that developed in the longitudinal grooves relative to that of the smooth surface. The results indicate that stress developed at the edges of the convex portion when R1 was larger than 4 mm and that the edge stress increased with an increasing radius of curvature. In contrast, stress did not develop at the edges when R1 was smaller than 4 mm, but contact stress increased at the ends of the convex portion"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003571_s0263574711000324-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003571_s0263574711000324-Figure1-1.png",
        "caption": "Fig. 1. Idealized motion of elastic body according to Bernoulli.",
        "texts": [
            " Section 4 analyzes a simulation example for a dynamic movement of a multiple DOF elastic robotic pair with elastic gear and flexible link in the presence of the second mode and external force. Section 5 gives some concluding remarks. \u201cOriginal form of Euler\u2013Bernoulli equation\u201dof the elastic line of beam bending has the following form: \u03b21,1 \u00b7 \u22022y\u03021,1 \u2202x\u03022 1,1 + m\u0302el1,1 \u00b7 d2y\u03021,1 dt2 \u00b7 (x1,1 \u2212 x\u03021,1) = 0, (1) where M\u03021,1 = m\u0302el1,1 \u00b7 d2y\u03021,1 dt2 \u00b7 (x1,1 \u2212 x\u03021,1) (Nm) is the load moment, in these source equations encompassing only inertia, \u03b5\u03021,1 = \u03b21,1 \u00b7 \u22022y\u03021,1 \u2202x\u03022 1,1 (Nm) is bending moment (see Fig. 1). Equation (1) is defined under the assumption that the bending moment is opposed only by the proper inertial moment. Besides, it is supposed by definition that the motion in Eq. (1) is caused by an external force F1,1, suddenly added and then removed. Euler\u2013Bernoulli Eq. (1) is expanded in refs. [5, 6] from several aspects in order to be applicable in a broader analysis of elasticity of robot mechanisms. The load moment is composed of all moment acting on the first mode of the link and these are perturbation moments, inertial moments (single and coupled moments), centrifugal, gravitational, Coriolis moments (single and coupled), coupled bending moments of the other modes, as well as the external force (which can be defined as static or dynamic force), which is via Jacobian matrix transferred to the motion of the first mode that come out directly from the motion dynamics of elastic bodies",
            " Elastic deformation defined in this way as in (29) is entered in the total dynamic system model. This is essentially different interpretation of the Eq. (29) in comparison with the interpretation of the same equation in this manuscript. In this paper, as explained above, Eq. (29) has completely new meaning. Equation (29) is a solution of dynamic models, i.e., form of elastic lines in space of Cartesian coordinates. The motion of the considered robotic system mode is far more complex (see ref. [5, 6]) than motion of the body presented in Fig. 1. This means that the equations that describe the robotic system (its elements) must be also more complex than Eq. (1), formulated by Euler and Bernoulli. This fact is overlooked, and the original equations are widely used in the literature to describe the robotic system motion. This is very inadequate because valuable pieces of information about the complexity of the elastic robotic system\u2019s motion are thus lost. Hence, we should emphasize the necessity of expanding the source equations for the purpose of modeling robotic systems and this should be done as in ref"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001032_j.jelechem.2008.07.020-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001032_j.jelechem.2008.07.020-Figure2-1.png",
        "caption": "Fig. 2. Comparison of quasireversibility produced by slow heterogeneous kinetics or by uncompensated resistance: For simulations all curves a = 1.0 cm2, jdEapp j \u00bc 1:0 10 3 V, v = 0.001 V s 1, Cdl = 1.0 10 5 farad cm 2, n = 1 For curve (a): ks = 1.0 cm s 1, a = 0.5 and Ru = 5 105 X; for curve (b): Butler\u2013Volmer kinetics (see Appendix Eqs. (37) and (38)) with ks = 1.0 10 7 cm s 1, a = 0.5 and Ru = 0; for curve (c) Marcus\u2013Hush kinetics (see Appendix Eqs. (39) and (40)) with ks = 1.0 10 7 cm s 1, k = 1.0 eV and Ru = 0.",
        "texts": [
            " The time required to achieve steady-state, sss, can be approximated by: sss P 4aRuCdl: \u00f016\u00de The values of Estart and Erev that will ensure that steady-state charging of the double layer capacitance is effectively achieved before the faradaic process becomes significant are defined by: The analysis I present here assumes that the conditions of Eq. (17) have been met. The focus of the present work is on the effects of the uncompensated resistance on reversible electron transfer. Aficionados of cyclic voltammetry will immediately recognize that the shapes of the widely separated peaks produced by large values of iRu (Fig. 2a) cannot be explained by slow heterogeneous electron transfer (using either the Butler\u2013Volmer formalism (Fig. 2b) or Marcus\u2013 Hush (Fig. 2c) [3] (see Appendix). For very large values of aFjv jRuCdl/RT and/or aF3jvjCtotal Ru/(RT)2 virtually all of the ohmic potential drop occurs across the resistance Ru. The observed response will then be ohmic \u2013 manifested by a virtually linear relationship between current and voltage. For very small peak separations, however, the peak shapes for electron transfer control and uncompensated-resistance control are almost identical \u2013 an analogous congruence was recognized long ago for solution based species [9]",
            "00) appears to be truly independent of the double layer capacitance, Cdl, over a very large range RTCdl/(n2F2 Ctotal) as shown in Table 1. There is a subtler indication that DEpeak/DipeakRu reaches a lower limit of 1.0 as Ru)1 (or (nF/RT)DEpeak)1). However, it is possible to offer a graphical argument that when nFDEpeak/RT)1, DEpeak/DipeakRu) 1.0. Consider the voltammograms shown in Fig. 4 which show simulations for Ru = 2 108 X, aF3jvjCtotal Ru/(RT)2 = 2.92 104 with (a) Cdl = 0; (b) Cdl = 1.0 10 6 farad cm 2; and (c) Cdl = 1.0 10 5 farad cm 2. Other parameter values are the same as those used in Fig. 2. The values of DEpeak/DipeakRu are virtually unchanged. Inspection of the responses reveals that regardless of the value of the capacitance (and as long as the \u2018\u2018RC\u201d time constant for the capacitive charging is sufficiently short \u2013 see Eqs. (15)\u2013(17) and accompanying discussion), the current response curves in the vicinity of the peak are well represented by a straight line with a slope 1/Ru and an intercept with the zero-current line approximately (but not exactly) where E E00 = 0 (the variation in interfacial potential required to drive the redox process is very small compared to the very large potential drop associated with iRu)",
            " Because of the weak dependence on the dou- ble layer capacitance, the expression is adequate as long as RTCdl/ (n2F2Ctotal) 6 0.1. The key question that should accompany interpretation of any cyclic voltammetric response is: Are non-zero values of jEpeak E00j caused by uncompensated resistance or by slow electron transfer, i.e., by a small heterogeneous rate constant ks (s 1) [11]? For large values of nF Epeak E00 =RT produced by large Ru (with) or by small ks (with Ru = 0 and implementing either the Marcus\u2013Hush or Butler\u2013Volmer formalisms \u2013 see Appendix) the distinction is visually obvious (see Fig. 2). For small values of nFjEpeak E00 j/RT the responses predicted by the Marcus\u2013Hush formalism and Butler\u2013Volmer formalism (with a = 0.5) become identical (see Appendix). Furthermore, the distinction between \u2018\u2018Ru-control\u201d and \u2018\u2018ks-control\u201d) is subtle (just as it is for a soluble redox species [9]). For nFjEpeak E00j/RT 1 (assuming a = 1/2): \u00f0Epeak E00\u00de ks v ffi F\u00f0Epeak E00\u00de=RT 1 0:750 or DEpeak ks v ffi F\u00f0Epeak E00\u00de=RT 1 1:50: \u00f025\u00de Note that the parameter (Epeak E00)ks/v is independent of Ctotal. (The relationship expressed in Eq",
            ") experimental ploy, it is virtually impossible to distinguish \u2018\u2018Ru-control\u201d from \u2018\u2018ks-control\u201d when FDEpeak/RT 1 as shown in Fig. 5. For the voltammograms shown in Fig. 5A (left ordinate) the input val- (right ordinate): iRu iks . : 50 S.W. Feldberg / Journal of Electroanalytical Chemistry 624 (2008) 45\u201351 ues of Ru and ks are set to: 0 and 1 to produce the reversible response with DEpeak = 0.0 V, 2663.3 X and 1 or 0 X and 0.15 s 1to produce the quasireversible response with DEpeak = 0.010 V as deduced from Eqs. (25)\u2013(27). For all voltammograms Cdl = 2 10 5 farad cm 2 and other parameter values are the same as for Fig. 2. Fig. 5B (right ordinate) shows the difference, iRu iks , between the quasireversible responses for Ru = 2663.3, ks =1 and Ru = 0, ks = 0.15 s 1 is less than 1% of the peak current \u2013 i.e., the responses are effectively indistinguishable. The analysis presented here is based on the presumption that the uncompensated resistance is constant and that the current density and potential are uniformly distributed over the electrode surface - this will not be the case for an ultramicroelectrode (ume) (e"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000047_s1560354707040028-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000047_s1560354707040028-Figure1-1.png",
        "caption": "Fig. 1. Vagner\u2019s realization of the Suslov problem.",
        "texts": [
            " The inertia tensor I can be represented as a diagonal 3 \u00d7 3 matrix with positive entries by choosing the body frame to be aligned with the principal axes of inertia of the body. We consider the evolution of the system subject to the (reduced) constraint a \u00b7 \u03a9 = c. REGULAR AND CHAOTIC DYNAMICS Vol. 12 No. 4 2007 3 is a constant vector in the body frame, c is a real number and \u03a9 \u2208 g \u223c= R 3 is the angular velocity expressed in body coordinates, that at all time satisfies \u03a9\u0302(t) = g\u22121(t)g\u0307(t) where g(t) \u2208 SO(3) gives the orientation of the body at time t. The homogeneous system where c = 0 can be realized as shown in Fig. 1, (this realization is due to Vagner [16] and was taken from [8]). The rigid body rotates about the fixed point O. The vector a is fixed and attached to the body. We imagine that there is an axis passing through O and perpendicular to a with sharp wheels on its ends, and with the vector a lying in the plane of the wheels. The body is then placed inside of a sphere with center at O with the wheels touching the inside of the sphere. The wheels are imagined to be so sharp that they forbid rotations about the axis determined by a"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002060_s11746-009-1511-3-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002060_s11746-009-1511-3-Figure1-1.png",
        "caption": "Fig. 1 Schematic diagram of chemiluminometer used in this work (description in the text)",
        "texts": [
            " The reaction of lipid hydroperoxides with sodium hypochlorite leads to the formation of oxygen singlet molecules in an electronic excitation state, with the subsequent light emission occurring after the electron relaxation process. This method could be used as a very fast, cheap and easy procedure for determining lipid hydroperoxides formed during lipid oxidation [7, 14, 15]. The aim of this study was to develop an inexpensive and easy CL reaction that could be used for the determination of lipid hydroperoxides in a non-aqueous medium and without the necessity of using expensive protein catalysts. Apparatus A static chemiluminometer built by the first-named author was used for the CL measurement. The chemiluminometer (Fig. 1) was equipped with a piston pump (2). The chemiluminometer was also equipped with a mechanical agitator (5) and a thermostat (8). The sample solution to be analyzed was injected with a syringe (3) into the Reodyne valve sample loop (4). The solvent from the reservoir (1) was used to pump the analyte to a test tube (7) of a chemiluminometer (6). Photons emitted from the test tube were counted by a PMT located in the second section of the chemiluminometer (11). Both sections were connected to a Teflon optical port (9)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003751_icuas.2013.6564678-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003751_icuas.2013.6564678-Figure2-1.png",
        "caption": "Fig. 2. Pure yawing motion",
        "texts": [
            " CD, CL and CM are the aerodynamic coefficients for drag force, lift force and pitch moment respectively The lateral dynamic generates the roll motion and, at the same time, induces a yaw motion (and vice versa), then a natural coupling exists between the rotations about the axes of roll and yaw [11]. In our case, to solve it, we have considered that there is a decoupling of yaw and roll movements [4]. Thus, each movement can be controlled independently. Generally, the effects of the engine thrust are also ignored [11]. In the Figure 2, the yaw motion is represented, which can be described with the following equations: \u03c8\u0307 = r (9) r\u0307 = N Izz (10) V\u0307y = Fy m \u2212 rVx (11) V\u0307x = Fx m + rVy (12) where \u03c8 represents the angle of yaw and r denotes the yaw rate, with respect to the centre of gravity of the airplane, N is the yawing moment and Izz represents the inertia in the z-axis. \u03b4r is the rudder deflection. Vx corresponds to the speed of the airplane in the longitudinal x-axis, Vy is the speed in the lateral y-axis, Fx describes the thrust force in the longitudinal x-axis and Fy denotes the component of the resultant lateral force on the yaxis"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003977_cdc.2012.6427069-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003977_cdc.2012.6427069-Figure1-1.png",
        "caption": "Fig. 1: Measurement of relative orientation angle.",
        "texts": [
            " Further, x(t) asymptotically converges to a finite point in EnN as t \u2192 \u221e. Consider N single-integrator modeled agents in the plane: p\u0307i = ui, i = 1, . . . , N, (2) where pi \u2208 R2 and ui \u2208 R2 denote the position and the control input, respectively, of agent i \u2208 {1, . . . , N} with respect to the global reference frame, which is denoted by g \u2211 . We assume that agent i \u2208 {1, . . . , N} maintains its own local reference frame with the origin at pi and the orientation angle \u03b8i \u2208 (\u2212\u03c0, \u03c0] with respect to g \u2211 as illustrated in Fig. 1. We denote the local reference frame of agent i by i \u2211 . By adopting a notation in which superscripts are used to denote reference frames, the agents are described as p\u0307ii = ui i, i = 1, . . . , N, (3) where pii \u2208 R2 and ui i \u2208 R2 denote the position and control input, respectively, of agent i with respect to i \u2211 . Given a graph G = (V, E , A) with V = {1, . . . , N}, which is referred to as the interaction graph of the agents, we assume that agent i \u2208 V measures the relative positions of its neighboring agents with respect to i \u2211 . Thus the following measurements are available to agent i: piji := pij \u2212 pii, j \u2208 Ni, i \u2208 V. (4) Defining \u03b8ji as \u03b8ji := PV (\u03b8j \u2212 \u03b8i) , j \u2208 Ni, (5) where PV (\u03b8j \u2212 \u03b8i) := [(\u03b8j \u2212 \u03b8i + \u03c0) mod 2\u03c0] \u2212 \u03c0, we additionally assume that \u03b8ji, j \u2208 Ni, are available to agent i \u2208 V . The relative orientation angles can be measured as follows. As depicted in Fig. 1, agent j measures the angle \u03b4ij and transmits the measured value to agent i by communication. Then agent i measures the angle \u03b4ji and calculates the relative orientation angles based on the following formula: PV (\u03b8j \u2212 \u03b8i) = PV (\u03b4ji \u2212 \u03b4ij + \u03c0) . We further assume that the agents estimate their orientation angles with respect to g \u2211 and their positions with respect to a common reference frame c \u22111 and they receive 1The common reference frame c \u2211 will be defined in Section III. the estimated values of their neighbors through communication",
            " Denoting the position of agent i with respect to c \u2211 by pci , the formation control problem is then stated as follows: Problem 2.1: For N -agents modeled by (2), let G be the interaction graph of the agents. Given p\u2217 = [p\u2217T1 \u00b7 \u00b7 \u00b7 p\u2217TN ]T \u2208 R2N , design a formation control strategy such that \u03b8\u0302i\u2212 \u03b8\u0302j \u2192 0, p\u0302ci \u2212 p\u0302cj \u2192 pci \u2212 pcj , and pci \u2212 pcj \u2192 p\u2217i \u2212 p\u2217j , based on measurements (4), (5) and (6). The orientations of g \u2211 and c \u2211 are not necessarily aligned while their origins are located at the same position as depicted in Fig. 1. Thus the objective of Problem 2.1 is to stabilize p to p\u2217 only up to congruence. For the agents under the assumption of Problem 2.1, an orientation estimation law can be designed as follows [5]: \u02d9\u0302 \u03b8i = k\u03b8\u0302 \u2211 j\u2208Ni aij(\u03b8\u0302ji \u2212 \u03b8ji), i \u2208 V, (7) where k\u03b8\u0302 > 0 and \u03b8\u0302i(0) = 0. Suppose that maxi\u2208V \u03b8i(0)\u2212 mini\u2208V \u03b8i(0) < \u03c0 for all i \u2208 V . Defining \u03b8\u0303i := \u03b8\u0302i \u2212 \u03b8i, we have maxi\u2208V \u03b8\u0303i(0) \u2212mini\u2208V \u03b8\u0303i(0) < \u03c0 for all i \u2208 V . Thus the orientation estimation error dynamics is written as \u02d9\u0303 \u03b8 = \u2212k\u03b8\u0302L\u03b8\u0303, (8) where \u03b8\u0303 = [\u03b8\u03031 \u00b7 \u00b7 \u00b7 \u03b8\u0303N ]T "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000506_j.engfracmech.2008.05.004-Figure4-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000506_j.engfracmech.2008.05.004-Figure4-1.png",
        "caption": "Fig. 4. Experimental gears, (a) unmodified gear and width profile, (b) modified gear and width profile.",
        "texts": [
            " The load distribution in a single meshing area (BD) of an unmodified spur gear with a constant width is a maximum along the tooth profile, whereas along the meshing entrance points (AE), the value drops to the minimum. Effort was made (by modifying the gear width) to correct the variable pressure distribution brought about by single and double gear meshes. In every point on the gear, care was taken to make sure that the ratio F/b (Force to gear width) remains constant by increasing the width proportionally according to the load applied. Modified and unmodified gears are shown in Fig. 4. To minimize the harmful effects of overload at the meshing region, the tooth width of the region exposed to the overload was widened. The specimen used in the experiment was AISI 8620 steel that had been carburized, quenched, and oil annealed, measured using a Rockwell hardness tester with 150 daN load, HRC 58 at the spur gear surface, and with a case depth of 0.8\u20131.1 mm. After manufacturing the gear wheels, the surface roughness was modified to Ra values between 0.3 lm and 0.6 lm by rasp. All gears were cleaned with solvents to remove the dirt and oil from the gear surface before the experiment"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000597_978-3-540-73129-0-Figure6.21-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000597_978-3-540-73129-0-Figure6.21-1.png",
        "caption": "Fig. 6.21. Resonant travelling wave ultrasonic motor",
        "texts": [
            " Piezomotors can produce elliptical motions either at the mechanical resonance (leading to ultrasonic motors) or in quasistatic (leading to stepping piezoelectric motors, so-called Inchworm\u00ae) [14]. The use of this motor in direct drive means that the complete function is obtained without any additional gear mechanism (for speed reduction, or for converting rotation in translation). Optics is probably the domain where the use of the piezoelectric motors is the most advanced. The most famous example, is the Canon camera, which includes an auto focus zoom based on a piezoelectric ultrasonic motor (USM) since 1992 (Fig. 6.21) [12]. Several other concept have been developed since then; few of them have found industrial applications. The motor from Elliptec is using a multilayer component, encased in a structure to couple two flexural modes of the beam (Fig. 6.22a) [13]. The stator includes a play recovering mechanism in the form of a spring that: \u2013 applies the preload force between the vibrating stator and the moving member; \u2013 guides the stator; \u2013 decouples the vibrations in the stator from the ground. Such a vibrating stator can be implemented in various ways (Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001265_tmag.2009.2012576-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001265_tmag.2009.2012576-Figure2-1.png",
        "caption": "Fig. 2. Magnets for measuring eddy current losses. (a) Bulk and (b) segmented.",
        "texts": [
            " Next, the total ac loss (including the ac copper loss in the solenoidal coil and the eddy current loss in the permanent magnet) is measured using a high precision wattmeter (YOKOGAWA Company, WT3000, resolution: 0.1 [mV]). The eddy current loss in the magnet is obtained by subtracting the ac loss when there is no magnetized magnet in the coil from the above-mentioned total loss. The justification of the measurement method will be shown in Fig. 4. The measurement frequency is 500 Hz 1800 Hz. The flux density B in the magnet is 13.5 mT (constant). Fig. 2 shows the measured magnets. Two kinds of fully magnetized magnets are used. One is a relatively large magnet (cross section is 20 mm 20 mm, and the length is 40 mm) as shown in Fig. 2(a). The other is a segmented magnet which is composed of four magnets (each cross section is 10 mm 10 mm, and the length is 40 mm) as shown in Fig. 2(b). The direction of magnetization of adjacent magnet is different from each other. This means that the compressive (attractive) force between magnets is a few MPa. The flux distribution (dc) produced by four magnets, of which the magnetization of adjacent magnet is opposite (N or S) each other, is different from that of usual construction of magnets. But we can examine the eddy current loss of magnets, because the dc flux produced by magnets does not affect the eddy current loss due to the ac field"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000427_j.ijsolstr.2007.12.004-Figure9-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000427_j.ijsolstr.2007.12.004-Figure9-1.png",
        "caption": "Fig. 9. Deformation pattern 4 0.",
        "texts": [
            " Similar to deformation pattern 8, we are unable to produce deformation patterns 8 0 in the experiment. Since the calculation methods involved in these asymmetric deformations are basically the same as those used in their symmetric counterparts, the unknowns to be solved are omitted in these Tables. The reason why these asymmetric deformations coexist with their symmetric counterparts for exactly the same pushing force can be explained with a specific example, for instance, deformation patterns 4 0 as shown in Fig. 9. It is noted that the location of the first contact point H from clamp A is determined once the pushing force FA is specified. This statement is valid only when the contact near point H is distributed (Lu, 2007). However, the location of the point I, the other boundary point of the distributed contact region, can be determined only when the equilibrium and the geometric condition on segment CI is considered at the same time. In the calculation, if we specify that the point C of the middle folding segment lies on the central radius, we will obtain the symmetric deformation 4 as before. However, if we relax this symmetric condition and allow the middle point of the folding segment to be away from the central radius, such as at position C 0 in Fig. 9, then the equilibrium condition as well as this new geometric condition can still be satisfied with a different location of point I, denoted now as I 0. As to the left half of the elastica, since the pushing force on the left end is the same as the one on the right end, the location of the first contact point from the left will be symmetric with respect to the one on the right. However, the contact region on the left will be of different size. As a consequence, the elastica deformation exhibits asymmetric pattern 4 0",
            " The location of the middle point C of the folding segment can be in any place within the range 0 < 1 2 wC0 < 1 2 \u00f0w\u00feH w\u00feI \u00de, where w\u00fe H 2 and w\u00fe I 2 are angles of point H and I in symmetric deformation pattern 4. In other words, in the solution procedure, if we specify wC0 \u00bc 0, we will obtain deformation pattern 4. On the other hand, if we specify a non-zero wC0 within the range, we will obtain asymmetric deformation pattern 4 0 with the same equations. The pushing force FA producing the deflection curve in Fig. 9 is 200. The same argument applies to the calculations of variations 7 0, 8 0 and, 11 0, with more complexities. Table 6 shows three different types of asymmetric deformation 7 0 for the pushing force FA = 200. Deformation patterns 7 0a and 7 0c are the variations of 7a and 7b in Table 3. Besides these two variations there is one additional asymmetric deformation 7 0b, which has a slanted folding segment. The dashed curves in 7 0b and 7 0c show the left extreme position of the folding segments, while the solid curves represent the right extreme positions of the folding segment"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003415_j.mechmachtheory.2011.03.006-Figure3-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003415_j.mechmachtheory.2011.03.006-Figure3-1.png",
        "caption": "Fig. 3. The construction of the generalised Goldberg 5R linkage. (a) Two Bennett linkages are firstly superposed with the common link a/\u03b1 in the middle; (b) then kink angle is set to be any value. The new link is added to connect two joints at the bottom, two offsets appear correspondingly; (c) last, after removing three links and one joint inside, the generalised Goldberg 5R linkage is obtained.",
        "texts": [
            " The techniques developed by Goldberg can be summarised as the summation of two Bennett linkages to produce a 5R linkage, or the subtraction of a primary composite linkage from another Bennett linkage to form a syncopated linkage. Here we only consider the summation case of Goldberg 5R linkage. Fig. 1. The Bennett linkage. A more generalised 5R linkage was also proposed by Goldberg briefly [5] and later derived in detail by Wohlhart [8]. In the general case, two links which form the rigidified link are not collinearly posed. A variable \u201ckink angle\u201d was introduced, see Fig. 3. Therefore, Goldberg 5R linkage is a special case of the generalised 5R linkage when the kink angle is zero. When the kink angle in the generalised Goldberg 5R linkage equals to \u03c0, the two links adjacent to the common link are overlapped and the resultant linkage is in fact formed by subtracting Bennett linkage B from Bennett linkage A as shown in Fig. 4, which is the first variation of the Goldberg linkages given in [5]. Therefore, we call such linkage as the subtractive Goldberg 5R linkage. Consider two Bennett linkages shown in Figs"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000846_iros.2008.4650783-Figure3-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000846_iros.2008.4650783-Figure3-1.png",
        "caption": "Fig. 3. Two effects of bisecting hip mechanism on leg-swinging motion",
        "texts": [
            " Let E [J] be the robot\u2019s total mechanical energy, which is defined as the sum of kinetic and potential energies: E(\u03b8, \u03b8\u0307) = 1 2 \u03b8\u0307 T M (\u03b8)\u03b8\u0307 + P (\u03b8), (10) and its time derivative satisfies the following relation: E\u0307 = ( \u03b8\u03071 \u2212 \u03b8\u03073 ) u1 + ( \u03b8\u03072 \u2212 \u03b8\u03073 ) u2. (11) By using Eq. (4), this can be rearranged as E\u0307 = \u03b8\u0307Hu1 2 \u2212 \u03b8\u0307Hu2 2 , (12) where \u03b8H := \u03b81 \u2212 \u03b82 [rad] is the relative hip-joint angle. This implies that each joint torque actuates the hip-joint alternately or these two control inputs are redundant actuations for one joint. Here, we analyze the effect of the upper body incorporating the BHM as a counterweight. We consider two cases: (a) the torso is put to the ceiling, and (b) the stance leg is put to the level floor, as shown in Fig. 3. 1) Case (a): The dynamic equation of the legs is [ I +mb2 0 0 I +mb2 ] [ \u03b8\u03081 \u03b8\u03082 ] + [ mbg sin \u03b81 mbg sin \u03b82 ] = [ 1 1 ] \u03bb, (13) where \u03bb \u2208 R is the constraint force of the BHM, and it can be solved as \u03bb = mbg (sin \u03b81 + sin \u03b82) 2 . (14) Considering \u03b81 = \u2212\u03b82, we can find \u03bb = 0. Thus, we can conclude that the BHM does not destroy the natural swinging motion of the legs at all. 2) Case (b): The dynamic equation in this case is [ I +mb2 0 0 IT +mT l 2 T ][ \u03b8\u03082 \u03b8\u03083 ] + [ mbg sin \u03b82 \u2212mT lT g sin \u03b83 ] = [ 1 \u22122 ] \u03bb"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000729_0020-7403(77)90022-4-Figure6-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000729_0020-7403(77)90022-4-Figure6-1.png",
        "caption": "FIG. 6.",
        "texts": [
            "; they were fired at approximately V0 = 130 m/sec-' and at 00 = 15 \u00b0 and v0 --- 110 m/sec l and 00 = 10 \u00b0 respectively. Similar results were obtained for the range of firing speeds 50-175 m/sec -~. Photographs associated with the penetration phase are those shown in Figs 5(a), (b) and (c); Fig. 5(d) pertains to the ascending phase and Fig. 5(e) to the emergent phase. Fig. 5(f) shows several superimposed profiles during emergence. Immediately after impact the sand is seen to be scattered in a well-defined (in plan) horse-shoe shaped pattern. The forward angle of sand dispersion, 8, in the plane of symmetry see Fig. 6, depends on the initial speed and the projectile position in sand. In the penetration phase of ricochet, 6 is found to be greater than 45 \u00b0 . This suggests that the \"wetted area\" of the projectile is not symmetrical about its direction of motion. A typical photograph used for determining the angle and speed at exit is shown in Fig. 5(f). The film was exposed to six successive illuminations at 4 \u00d7 10 -~ second intervals. In all the tests reported here the angles of firing and exit were found to be approximately equal"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003136_1.3645806-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003136_1.3645806-Figure2-1.png",
        "caption": "Fig. 2 Microsl ip patterns for a ' /2- in-dia bal l w i t h a load of 10 l b / showing effect of geometric conformi ty at var ious values of R/r",
        "texts": [
            " a = R/r. The values of y are now 0 and \u20142R c r /R . The values of y in each of these cases may be greater than y= B so t ha t the contact region may contain only one region of sticking. In view of the large number of geometric and load variables it is not possible to produce an exhaustive summary of the results from equations (12) to (14). Nonetheless the basic significance of these equations ma}' be demonstrated by the calculations leading to the microslip pat terns shown in Figs. 2, 3 and 4. Fig. 2 Bhows the microslip pat terns for a value of a \u2014 R/r with a -J-in-dia ball and a normal load of 10 lb. I t is clearly seen t ha t one value of y is always zero, the other being greater than B. The effect of decreasing conformity, i.e., increasing RT, is to increase the area of sticking and the effect of increasing R/r is again to increase the area of sticking. For R/r = co the maximum stick area occurs when RT = m, the slip being confined to a vanishingly small area a t the trailing edge of the contact zone"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000474_gt2008-51179-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000474_gt2008-51179-Figure2-1.png",
        "caption": "Figure 2. Cross section of the turboexpander TC 400/90",
        "texts": [
            " The bearing support structure is relatively stiff compared to the oil film and can therefore be regarded as rigid. The bearing carrier also includes the labyrinth seals and supports the expander nozzle assembly and booster diffuser. The wheels are connected to the shaft by a Hirth radial spur tooth coupling and tightening bolt. The assembly of the bearing carrier with nozzle assembly, diffuser and shrouds is called plug-in unit. Expander and compressor housings are assembled on each side of the flanged bearing carrier (see Fig. 2). Calculated basic rotor dynamic behavior The Campbell diagram (natural frequencies and damping ratios as a function of speed) of the shaft is shown in Fig. 3 together with the mode shapes of the forward whirling modes. The speed range from the nominal speed to maximum continuous speed in the diagram is grey shaded. There are two forward and backward whirling rigid body modes below nominal speed with a very high damping (damping ratio >20%). The forward whirling 1 st and 2 nd bending modes are above the speed range with a comfortable separation margin",
            " With this execution both turbo-expanders and their spare plug-in units were successfully FAT tested with the customer and all accepted without reservations. After this new exhaustive experience with the vibration induced hot spots, based on predictive analysis and testing, the unexplained historical cases from the years 1996 and 2001 were analyzed and the unexpected excessive vibrations during FAT were re-evaluated. Both cases with still available ADRE vibration records could be clearly identified as the spiral vibration problem (see Fig. 22 and 23). Both cases happened on turbo-expanders with similar shaftbearings configurations as shown in Fig. 2 and 11. The first case (Fig. 22) was identified by slight labyrinth rubbing marks on shaft and was corrected by increasing the labyrinth clearance. The second case was tested up to 104 m/s bearing journal 7 Copyright \u00a9 2008 by ASME Terms of Use: http://asme.org/terms velocity (Fig. 23) and was corrected during FAT by modification of the bearing clearance. But the vibration problem solving was a kind of a try and error solution, the predictive analytical method was missing. In another case of integral gear expander-compressor machine, the predictive hot spot stability analysis was performed during the design stage to eliminate any possible vibration problems in advance"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003880_apcase.2015.44-Figure9-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003880_apcase.2015.44-Figure9-1.png",
        "caption": "Fig. 9. Principle of tilt sensor",
        "texts": [
            " The tilt sensor is mounted on the caregiver\u2019s back, and the pressure sensor is mounted on the back of the caregiver\u2019s right hand. Overview of each sensor is shown in Fig. 8. When the caregiver performs \u201c4) Lifting\u201d, he/she holds the care-receiver by both hands with putting the palm of the left hand on the back of the right hand. At this moment, the pressure sensor reacts. Therefore, the caregiver\u2019s contact to the care-receiver can be detected by the pressure sensor. In this study, RAS-2C produced by Kondo Kagaku co.,ltd. is used as a tilt sensor. The principle of the tilt sensor is illustrated in Fig. 9. As shown in Fig. 9, the tilt sensor is always outputting the reference voltage to the direction of gravitational acceleration. When the sensor inclines, the voltage variation depending on the degree of inclination occurs. Then, the following equation holds. (2) Therefore, the ascending vertical angle can be obtained from the following equation by measuring and . (3) where the measurement range of the ascending vertical angle is restricted to 0 to 90 degrees. In addition, an angular velocity is obtained by the difference value of the ascending vertical angle"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000168_iros.2006.282537-Figure3-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000168_iros.2006.282537-Figure3-1.png",
        "caption": "Fig. 3. The Importance of Preview Control System",
        "texts": [
            " In this study, control variable v(t) is determined so as to bring joint angles at current time q(t) close to joint angles of imaginary manipulator q\u0303d(t\u2217) at time t\u2217 to satisfy non-collision requirement, and is given by Eq.(4), v(t) = Kpr(q\u0303d(t \u2217) \u2212 q(t)) (4) where, Kpr is a positive definite diagonal matrix representing gains, that is, Kpr = diag[kv1, kv2, \u00b7 \u00b7 \u00b7 , kvn]. Substituting Eq.(4) into Eq.(3) constitutes preview control system to use the future possible configuration, that is, the joint angles q\u0303d(t\u2217) satisfying non-collision requirements obtained by configuration planning at time t\u2217 is utilized to control for current configuration q(t). For example, as shown in Fig.3, when the hand reaches the position B1, two kinds of the manipulator\u2019s shapes denoted by P1 and P1 \u2217 both can avoid collision. However, when the hand reaches the position B2, only the shape of P2 \u2217 in the two shapes denoted by P2 and P2 \u2217 shown in Fig.3 can avoid collision. If the manipulator\u2019s shape is selected as P1 at hand point B1, the angular velocities of joints will be high values to change its shape like P2 \u2217 near the corner B. This poses a possibility that the manipulator crashes to corner B when the required high angular velocity is over specified maximum velocity of the joint. Therefore, the manipulator\u2019s shape must be prepared to the shape P1 \u2217 that is similar configuration to P2 \u2217 rather than P1 to ensure non-collision. This requires that the current manipulator\u2019s shape should be determined in a consideration of future possible configuration such as P2 \u2217, this is so-called preview control system which is depicted in Fig",
            " In addition, in this study, the process of using camera to discern the instantaneous trajectory tracking will be finished in 33 milliseconds. However, simulation indicates that the maximun AMSIP value can be obtained with GA method through about 30 generations. This process will cost 61 milliseconds or so. So, when the best manipulator\u2019 shape with GA method is obtained, the object orbit will move to other position and manipulator will not follow desired tracking. According to above discussion, 1-step GA method is adaptable for the realtime optimization. The desired hand trajectory on the surface of working object is shown in Fig.3, that is the line connecting from A to F. The distributions of AMSIP value 1S and the manipulator\u2019s shapes at the maximun peak when the hand is fixes at A, B, C, D, E and F respectively are shown from Fig.7 to Fig.12. The whole simulation time is set by 50 seconds and preview time t\u0303 denoted by (t\u2217 \u2212 t) is set by 10 seconds, which leads to predictive avoiding function all through the desired hand trajectory. Then, using preview control system to make the actual manipulator\u2019shape at A close to desired shape of imaginary manipulator at B to satisfy the real-time optimization"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002188_bj1220079-Figure5-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002188_bj1220079-Figure5-1.png",
        "caption": "Fig. 5. Reaction of 5,um-ferroperoxidase with cyanide in 50mm-borate buffer, pH9.1, at 200C. The pseudo-firstorder rate constant, k', is plotted against the concentration of NaCN.",
        "texts": [
            " Kinetic8 ofthe binding ofcyanide toferroperoxidase. In the reaction: k+i Peroxidase2+ +HCN peroxidase2+.HCN k-t the rate equation is: -d [peroxidase2+] - k+ - [peroxidase2+]. [HCN] -k-I * [peroxidase2+ - HCN] The rate of approach to equilibrium under conditions where [HCN] is much greater than [peroxidase2+] is k' = k.. +Ak+1 - [HCN], where k' is the pseudo-first-order rate constant measured under such conditions. Thus a plot of k' against [RCN] should give a straight line with intercept equal to k-4 and a slope equal to k+1. Fig. 5 shows the graphical summary of three series of experiments in which 2,um-peroxidase was mixed rapidly at pH9.1 and 2000 with various concentrations of cyanide, the concentration of this latter being expressed as the total cyanide species present. The results indicated: k+j = 29M-1 *s-1 and k-I = 2.5x10-2s-1 The affinity constant K, i.e. k+l/k-., is 1.16x 103M-1, which agrees well with the statically determined value of 1.05 X 103 M-1 from titration at the same pH. As a further confirmation of the validity of these values, the total extinction change that occurs during the rapid-mixing experiments can be plotted to Vol"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003761_9781118818428-Figure6.1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003761_9781118818428-Figure6.1-1.png",
        "caption": "Figure 6.1 Notation for algebra of time of decay of laminar velocity profile based on instantaneous removal of driving pressure difference",
        "texts": [
            " Rex Britter suggests decay time should be proportional to l\u2215u, where l [m] is characteristic eddy dimension and u [m/s] is speed. The \u2018back of envelope\u2019 approach which follows leads to the estimate of \u2018half-life\u2019 \u0394t1\u20442 = rhRe\u2215u. With hydraulic radius rh [m] for characteristic linear dimension, this is in line with Britter\u2019s suggestion. To acquire a feel for numbers it is assumed that the half-life of an eddy is of the same order as half-life for collapse of the laminar velocity profile on instantaneous removal of the driving pressure difference. In Figure 6.1, which shows an element of cylindrical duct: \u03c4w = \u2212\u03bc\u2202u\u2215\u2202r (6.1a) = 1\u20442\u03f1u2Cf (6.1b) \u0394pAff = \u22122u\u03bcpwdx\u2215rh Sudden removal of \u0394p is equivalent to imposition of a decelerating force balanced by the product of mass element \u03f1Affdx with its deceleration du\u2215dt: du\u2215u = \u2212 ( 2\u03bc\u2215\u03f1rh 2 ) dt (6.2) Integrating: loge(u 1 \u2215u 0 ) = \u2212 ( 2\u03bc\u2215\u03f1rh 2 ) (t1 \u2212 t0) Time \u0394t1\u20442 for u to fall from u0 to half of that value is given by inverting Equation 6.3: \u0394t1\u20442 = \u2212 ( \u03f1rh 2 / 2\u03bc ) log e 1\u20442 = \u2212(rh\u22158u)Re \u22c5 loge 1\u20442 (6.3) Hydraulic radius rh of the exchanger slots of the Philips MP1002CA air engine is 0"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000329_j.electacta.2008.09.020-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000329_j.electacta.2008.09.020-Figure1-1.png",
        "caption": "Fig. 1. Cross-section of oxygen electrode equipped with carbon felt electrode on the oxygen gas permeable membrane. (A) recorder, (B) potentiostat, (C) potentiostat, ( e w o",
        "texts": [
            " n order to detect hydrogen peroxide produced by oxygen reducion, a rotating ring-disk electrode system (Nikko Keisoku, RRDE-1, otor speed controller SC-5, Dual potentiogalvanostat DPGS-1, otential sweeper NPS-2) using glassy carbon disk-platinum ring lectrode was used to detect of hydrogen peroxide produced at the arbon disk surface. o o r r D) Pt counter electrode, (E) Pt lead wire, (F) dialysis membrane, (G) carbon felt lectrode, (H) oxygen permeable membrane, (I) Ag/AgCl reference electrode, (J) gold orking electrode of oxygen electrode, (K) inner electrolyte, (L) counter electrode f oxygen electrode. .3. Measurements of oxygen consumption In order to confirm the consumption of oxygen by the electrode eduction, a porous carbon felt (CF) was used by combining with an xygen electrode. Fig. 1 shows the cross sectional view of the oxyen electrode equipped with CF electrode. A polarographic oxygen lectrode equipped with a carbon felt electrode was fabricated by mmobilizing the CF by an electrode cap. A CF consisting of carbon bers (about 10 m diameter) have very large porosity more than 0%, and oxygen can easily penetrate this carbon material. Then, the ffect of the electrode potential of CF on the oxygen consumption n the CF can be measured, because the dissolved oxygen transfers hrough the carbon felt and reaches to the oxygen electrode"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003713_978-1-4614-2419-2-Figure42.5-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003713_978-1-4614-2419-2-Figure42.5-1.png",
        "caption": "Fig. 42.5 Ballistic impact test specimen, hatched region represents the clamped portion",
        "texts": [
            " The ballistic impact tests were performed using an in house designed gas-gun setup equipped with an environmental chamber. Inside the chamber is a fixture which holds the specimen to be impacted. A 22 caliber copper bullet fitted in to a plastic sabot was used as the projectile. The sabot is very important and provides a tight fit of the projectile in the gun barrel. The gun consists of helium tank, temporary gas storage vessel, solenoid valve and stainless-steel barrel. Figures 42.1, 42.2, 42.3 and 42.4 show the gas gun with the environmental chamber fixture and bullet with sabot. Figure 42.5 shows the dimensions and clamped areas of the specimen to be impacted. The boundary conditions of the clamped sides are assumed to be fixed-fixed. Fig. 42.2 Ballistic gas gun and environmental chamber Fig. 42.3 Fixture 414 Y. Budhoo et al. Compressed helium was used to launch the projectile. The velocity of the projectile can be varied by changing the pressure used to launch the projectile or by changing the location of the bullet in the gun barrel. Placing the bullet closer to the exit of the gun barrel lowers the velocity of the bullet since the pressure used to launch the bullet will be smaller as compared to when the bullet is placed further away from the exit"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002335_detc2010-28926-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002335_detc2010-28926-Figure1-1.png",
        "caption": "Figure 1. Two configurations with the same topology of a determinate truss (rigid graphs).",
        "texts": [
            " Since the paper deals only with the topology of planar linkages and all the mathematical foundation of this paper is based on graph theory, the terminology used in the paper is from graph theory and can be found in any basic textbooks on the subject, such as (Swamy and Thulasiraman, 1981). For example, joints are referred to as vertices, links as edges and structures as graphs. Moreover, to avoid other terminologies used in the rigidity theory community and not in mechanical engineering, the definitions appearing in the paper are slightly modified by giving them more physical than combinatorial meaning. To clarify the terminology used in the paper let us define the structure depicted in Figure 1 in both terminologies. In the terminology of engineering this is a determinate truss with four rods/bars, two joints \u2013 A and B, three pinned joints connecting rods 1,2 and 4 to the ground, while each rod has its specific geometry (length, inclination angle, etc.). Therefore, in engineering terminology there is a difference between the two determinate trusses in Figure 1. In the terminology of rigidity theory the graph in Figure 1a is a rigid graph with four edges, two inner vertices, three ground vertices, three ground edges \u2013 1,2 and 4 and there is no notion of geometry of the elements. Thus, from the rigidity theory point of view there is no difference 2 Copyright \u00a9 2010 by ASME Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 08/12/2018 Terms of Use: http://www.asme.org/about-asme/terms-of-use 3 Copyright \u00a9 2010 by ASME between the two graphs in Figure 1. Now, we shall define Assur graphs and outline what distinguishes them from other rigid graphs. Assur Graph \u2013 is a minimally rigid graph with e(G)=2*v(G) where e(G) and v(G) stand for the number of edges and inner vertices of graph G, respectively. The main property of the graph is that removal of any vertex with its incident edges makes the graph non-rigid. The graph, appearing in Figure 2(a) is an Assur Graph since the number of the edges is twice the number of the inner vertices, it is rigid and all its sub-graphs are not rigid"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003202_0022-2569(66)90030-9-Figure6-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003202_0022-2569(66)90030-9-Figure6-1.png",
        "caption": "FIG. 6.",
        "texts": [],
        "surrounding_texts": [
            "4\n3 2\nFtG. 4.\nthe corresponding spherical mechanism (Fig. 4) we arrive at the deduction that motion is possible only in the event of coincidence of axis one with axis 4 (or with axis 2). Assuming that axis 1 is parallel to axis 4, we obtain as necessary conditions:\n~o =0 ; (15)\nand the existence of constant angles 0 o and Xo. These relations are determined from the principal parts (4'), (5') and (6') of equations (4), (5), and (6).\nEquation (4') will have the form\n{ [cos(ao + ? o ) - cos rio] + [cos(~o - Yo)- cos flo]\u00a2o2}Vo 2 + 4 sm ~o sin Yo*oVo\n+[cos(ao-Yo)-COSflo]+[cos(o~o+Yo)-COSflo']\u00a2~2=O. (4**)\nfrom which a function relating ipo and ~Po is obtained. Equation (5') results in the form\n{[cos(#o + ~o)- cos ~o] + [cos(#o +~,o)- cos ~o]O~}X~o\n+[cos(/~o-~o)-COS~o]+[cos(/~o-ro)-COS~o]Oo ~=o. (5**)\nThe left side of the equation, as can be shown, is identically zero. In fact, by the formulas of spherical trigonometry we have\ncos to cos ?o - sin to sin Yo cos Xo = cos ao,\no r\n1 - X o ~ cos to cos ?o - sin to sin ? o 1--~-~o 2 = cos ao.\nOn substitution of the latter expression into equation (5\"*) we will find that this equation reduces to an identity.\nEquation (6') has the form\n{[cos(,~o- Bo)- cos ~o] + [cos(~o- Po)- cos ~'o]\u00a2',]}Oo ~\n+[cos(o~o+flo)-COSyo]+[cos(ao+flo)-COSYo]02=O. (6**)",
            "From the equations of spherical trigonometry:\n1 cos ~o cos fie + sin 0~o sin fie cos 0o = cos 0~ o cos/7o + sin ~o sin fie i +\u00aeo 2-v\n= COS ~o \u2022\nWe find, for confirmation, that equation (6**) also reduces to an identity. The amount of sliding on axes 2, 3, 4 is determined from the dual parts (4\"), (5\") and (6\") of equations (4), (5) and (6). The lack of contradiction in these dual parts proves the sufficiency of condition (16). Thus we find that together with the pure sliding motion on axis 3, pure sliding also takes place on axis 2, i.e. the mechanism has two excess constraints. In this mechanism pure sliding occurs simultaneously in two adjacent pairs.\nThus the satisfaction of the requirement for pure translational motion along one of the axes unavoidably leads to the existence of another axis along which pure translation takes place. Therefore in four-link mechanisms we see it is necessary to have two axes with prismatic pairs. In one case these axes occupy diagonally opposite positions, and in the other case they are adjacent. Both of these cases are represented in Figs. 5 and 6.",
            "REFERENCES\n[1] F. M. DIMENTBERG, Determination o f the Motions o f Spatial Mechanisnts (in Russian). IZD-V() Akad. Nauk. USSR (1950).t\n[2] V. V. DOBROVOLSKII, Spherical representation of spatial four-link mechanisms. Transactions oJ the Institute o f Machine Design Seminar on Theory o f Machines and Mechanisms (in Russian), Vol. [I, p. 7. IZD-VO Akad. Nauk. USSR (1947),\n[3] A. P. KOTELNIKOV, Screw Calculus and Some Applications to Geometry and Mechanics (m Russian). Kazan (1895).\nt [Translator's Note: A shorter account of the same subject matter which has been translated into English by I. E. Morse, Jr., R. Sridhar and J. W. Moore is \"A General Method for the Investigation of Finite Displacements of Spatial Mechanisms and Certain Cases of Passive Joints\" by F. M. Dimentberg, which is available as Purdue Translation No. 436 from the Purdue University Library, West Lafayette, Indiana.]"
        ]
    },
    {
        "image_filename": "designv11_12_0002309_s11012-010-9380-2-Figure3-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002309_s11012-010-9380-2-Figure3-1.png",
        "caption": "Fig. 3 Planar 2RPR/RP PKM with DOF 2",
        "texts": [
            " This is obvious from the closed loop dynamics Ge\u0308 + Ce\u0307 + KDe\u0307 + KPe + Gq\u03082 + Cq\u03072 + Q \u2212 S(Gq\u0308d 2 + Cq\u0307d 2 + Q) \u2212 AT NAT c0 = 0, (25) when (23) is applied to the uncertain system (21), and from G(e\u0308 + KDe\u0307 + KPe) + Gq\u03082 + Cq\u03072 + Q \u2212 S(Gq\u0308d 2 + Cq\u0307d 2 + Q) \u2212 AT NAT c0 = 0, (26) when the computed torque controller (24) is applied. Now the linear feedback acts freely on the uncertain system, in contrast to (23) and (24). Therewith the uncertainties affect the dynamics of the controlled PKM, but not the way the controls affect the system. The second and third lines in (25) and (26) embody the uncertain dynamics that is not balanced by the controller. For illustration purpose the effect of geometric uncertainties of the planar RP/2RPR PKM in Fig. 3 is analyzed (underlines denote the actuated joints). This is a fully-parallel but not symmetric PKM. There is no moving platform, and the EE is mounted on one of the limbs. The EE is connected to the base by one RP and two RPR chains. The PKM is obtained from a nonredundant RP/RPR by adding one RPR chain. The base joints are mounted on the base at the corners of an equilateral triangle. A disturbance frequently encountered in setting up a PKM is the misplacement of joints. Now assume that one of the base joints is displaced on the ground plane with \u03b4x and \u03b4y as indicated in Fig. 3. This leads to a perturbed plant with input matrix AT . The control forces are deduced from the nominal model with AT . Consequently, the inverse dynamics solution (8) applied to the perturbed system (21) cannot perfectly reproduce the desired control forces, due to AT (AT )+ = I. This leads to desired forces in the null-space of AT becoming effective, due to AT NAT = 0. For a quantitative analysis the drive unit has been displaced by 5% of the triangle side length, as shown in Fig. 4. The perfect model and the perturbed plant are evaluated along the indicated EE path"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001242_s12239-009-0050-0-Figure16-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001242_s12239-009-0050-0-Figure16-1.png",
        "caption": "Figure 16. Dynamics model of full vehicle.",
        "texts": [
            " The steering angle of the 1st axle is controlled by the driver, but the articulation angle is geometrically determin- ed according to the steering state of each wheel. Therefore, a multi-body dynamic simulation was performed to investigate the steering angles and the trajectory of the vehicle. To verify the AWS algorithm, the commercial software ADAMS was used to validate the dynamic model and algorithm. After modeling each component, such as suspension, steering system, and tire, the dynamic model of the full vehicle was assembled, as shown in Figure 16. The data used for the model were based on the data from a real vehicle, such as that required for the dimensions of the vehicle, characteristics of the dampers, air-springs, and tires. Dry asphalt with the friction coefficient of 0.8 was applied to the road surface and the lateral forces on the tire were obtained from the Magic Formula model (Pacejka, 2002). was applied to the existing AWS control algorithm has Rmin= v 2 alat_max -------------- v 20\u2264 , P=Pmax 20 v 46< < , P=Pmax\u2212 v 20\u2013 46 20\u2013 ----------------- Pmax\u00d7 v 46\u2265 , P=0 been verified by comparison with field tests of the Phileas vehicle"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003854_978-3-319-10723-3_1-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003854_978-3-319-10723-3_1-Figure1-1.png",
        "caption": "Fig. 1 3-RPS manipulator design",
        "texts": [
            " The OM are separated by constraint singularities. Using Study coordinates (x0, x1, x2, x3, y0, y1, y2, y3) (see e.g. Husty et al. 2007) to describe the spatial Euclidean displacements, a complete analysis of the motion capabilities of a special 3-RPS PMwas performed in Schadlbauer and Husty (2011). The 3-RPSPM is a 3-dof PMproposed by Hunt (1983). This LMPM is often referred to as \u201c1T2R manipulator\u201d but this designation does not mean that the 3-RPS LMPM has two rotational degrees of freedom about two fixed axes. The 3-RPS PM, shown in Fig. 1, is composed of three limbs Li = Ri Pi Si , i = 1, 2, 3 such that: (i) The axis of Ri is directed along fic such that f1c, f2c and f3c are three independent unit vectors parallel to the plane of the fixed base, namely, orthogonal to n, (ii) The Pi -joint is directed along fia , (iii) The Si -joint is centered at point Bi , (iv) points Ai , Bi are vertices of equilateral triangles and fic have directions tangent to the circumcircle of Ai . In Schadlbauer and Husty (2011) it was found out that the workspace splits into two different components that are characterized by either x0 = 0 or x1 = 0. In Eq. (1) the set of constraint equations for the component x0 = 0 is displayed.1 The set for the other component is equally simple but is omitted here because of lack of space. 1 The design parameters h1 and h2 (Fig. 1) have been set to h1 = 1 and h2 = 2. Note that all following computations can be done without specifying these parameters, but the equations become longer. r00 := [0, x1y3 + 2x2x3 + x2 y0 \u2212 x3y1, \u22122x1y2 + 2x22 + 2x2 y1 \u2212 2x23 + 2x3y0, \u2212 R1x21 \u2212 R1x22 \u2212 R1x23 + 9x21 + 12x1y2 + 9x22 \u2212 12x2 y1 + x23 + 4x3y0 + 4y20 + 4y21 + 4y22 + 4y23 , (R1 \u2212 R2)(x21 + x22 + x23 ) + 18(x2 y1 \u2212 x1y2) + 6(x23 \u2212 x22 \u2212 x3y0) + 6 \u221a 3(x3y1 \u2212 x1y3) + 2 \u221a 3(x2 y0 + 2x2x3), (R1 \u2212 R3)(x21 + x22 + x23 ) + 18(x2 y1 \u2212 x1y2) + 6(x23 \u2212 x22 \u2212 x3y0) \u2212 6 \u221a 3(x3y1 \u2212 x1y3) \u2212 2 \u221a 3(x2 y0 + 2x2x3) x1y1 + x2 y2 + x3y3, x21 + x22 + x23 \u2212 1], (1) where Ri are the quadrances (squares) of the leg lengths ri "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002550_s0219455412500186-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002550_s0219455412500186-Figure1-1.png",
        "caption": "Fig. 1. A rotating \u00b0exible cylindrical shell partially \u00aflled with ideal liquid.",
        "texts": [
            " By using the rotating Euler Bernoulli beam model, Firuz-Abadi and Haddadpour23,24 presented analytical models for the \u00b0exural vibrations of long spinning cylinders partially \u00aflled with inviscid and viscous \u00b0uid and examined the e\u00aeect of the viscoelastic and hysteresis damping of the cylinder material on the stability boundaries. The aim of the present paper is to investigate the instability boundaries of a rotating cylindrical shell partially \u00aflled with ideal liquid. The structural dynamic model of a shear-deformable cylindrical shell is combined with the quasi 2D model of the \u00b0uid to obtain the coupled-\u00afeld model. The obtained model is used to determine the instability conditions of the rotating shell. 2.1. Fluid dynamics equations Consider a rotating \u00b0exible shell partially \u00aflled with ideal liquid as shown in Fig. 1. The cylinder spins with a constant angular velocity about its axis and is held by elastic supports at both ends. In practice, after a few time of starting the cylinder spinning, the liquid is formed into an annular shape that its motion is synchronized with the cylinder. Based on the literature results, for a cylinder with h=r > 2 the small perturbations of the liquid motion in the rotating frame r (Fig. 1) can be 1250018-3 In t. J. S tr . S ta b. D yn . 2 01 2. 12 . D ow nl oa de d fr om w w w .w or ld sc ie nt if ic .c om by J O H N S H O PK IN S U N IV E R SI T Y o n 01 /0 3/ 15 . F or p er so na l u se o nl y. described with a 2D model that is governed by the linearized Navier Stokes equations as follows: @v @t \u00fe 2 vr \u00bc 1 fr @P @ ; \u00f01\u00de @vr @t 2 v r 2 \u00bc 1 f @P @r ; \u00f02\u00de @\u00f0rvr\u00de @r \u00fe @v @ \u00bc 0; \u00f03\u00de where vr and v stand for the radial and tangential components of \u00b0uid velocity, respectively, and f is the liquid density"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003713_978-1-4614-2419-2-Figure16.1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003713_978-1-4614-2419-2-Figure16.1-1.png",
        "caption": "Fig. 16.1 COMBIN14 force displacement relationship",
        "texts": [
            " The base of each column was connected to a 600 600 1/200 steel plate using a 200 200 1/800 steel angle and four, \u00bc00 bolts for each column leg; the corresponding plate was then secured to an 80 40 1/200 steel plate to serve as a rigid base. The numerical model was created in the finite element software program ANSYS v. 12.0 using BEAM188 elements. These are two node 6 degree of freedom linear elements that take into account the cross section and orientation of the member. Rotational springs were used for all connections. The spring elements are the COMBIN14 element in ANSYS v.12.0 and operate under linear rotational spring principles (see Fig. 16.1. where M is the moment, krot is the rotational stiffness, and y is the rotation of the joint in the direction of interest). Three springs are used for each beam to capture rotations in all three directions. The stiffness coefficient of these springs is calibrated with experimental data to obtain the highest fidelity model. The translational degrees of freedom for each connection are coupled so that there is no relative motion between translation of the beam and translation of the column at the connection"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003048_1.b34104-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003048_1.b34104-Figure2-1.png",
        "caption": "Fig. 2 Displacement variables and coordinate system of a rotorbearing system.",
        "texts": [
            " The coefficients are used with a finite element model belonging to a C1-class formulation for the study of whirl speeds and unbalance response analyses. A transient analysis is performed to evaluate the transition of the system through resonance. Vibration amplitudes are evaluated during turbine startup andwhen it is operating in steady state conditions. At the end of the paper, a spectral map plot is presented to identify possible vibration problems when considering excitation frequencies from 0 to 500 Hz. The configuration of a simple rotor-bearing system is illustrated in Fig. 2.We assumed that, as compared with the translational motions, the axial motion is small enough to be reasonably neglected. A typical cross-section of the shaft located at a distance s from the left end, in a deformed state, can be described by the translations V s; t and W s; t in the y- and z-directions as well as the small rotations s; t and s; t about the y- and z-axes. The relationships can be expressed as V s; t Vb s; t Vs s; t (1a) W s; t Wb s; t Ws s; t (1b) s; t @Wb s; t @s (1c) s; t @Vb s; t @s (1d) whereVb,Vs, andWb,Ws are translations due to bending and shear in the y- and z-directions, respectively"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001209_ssp.147-149.542-Figure3-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001209_ssp.147-149.542-Figure3-1.png",
        "caption": "Fig. 3. The pressure distributions caused by the rotation in circumferential direction in parabolic micro-bearings. Left side presents the view from the film origin, right side shows the view from film end",
        "texts": [
            " Friction coefficients are as follows [7]: )p( tot pRpR p C FF jj   ee , (9) where p , j ee are the unit vectors in parabolic coordinates. We determine the pressure distributions and load carrying capacity values in HDD micro-bearing for parabolic journal in the lubrication region F, which is defined by the following inequalities: 0  j  jk, p1=p/bp, bp  p  bp where 2bp  micro-bearing length. Numerical calculations are performed in Mathcad 14 Program by virtue of the equation (3), (8) by means of the finite difference method (see Fig. 3). If grooves length is situated in p and j direction then gap height of the parabolic micro-bearing has the following form respectively [1]: ),,,()n5,0(H)1(cos)t,(1)t,,( p k 0n T n 1gpppT j       jjjj   (10) ),,,()n5,0(H)1(cos)t,(1)t,,( p k 0n Tp n 1gpppT j       jj   for 0  j < 2p, bp  p  bp where p  eccentricity ratio in parabolic micro-bearing,   radial clearance in parabolic micro-bearing, g1g/, g  ridge height, H  Heavisidea unit function. Symbols jT,T denote periods of grooves sequence about 65nm in j and  directions respectively, k  number of ridges about 1000. Symbol  denotes the dimensional random part of gap height changes resulting from vibrations, unsteady loading and surface roughness measured from the nominal mean level. The symbol  describes the random variable, which characterizes roughness arrangement. We show in Fig. 3 the results of numerical calculations of pressure without magnetic field influences and stochastic changes. The grooves and ridges are now neglected. We assume the largest radius of the journal a=0.001 m, the smallest radius of the journal a1= 0.0008 m, length/radius ratio Lp1=bp/a=1, dynamic viscosity of the oil o=0.03 Pas, angular velocity w=565.5 s 1 , characteristic dimensional value of hydrodynamic pressure po=wo/ 2 =16.96 MPa,, relative radial clearance =T/a=0.001, eccentricity ratio p=0.4; p=0"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003713_978-1-4614-2419-2-Figure7.3-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003713_978-1-4614-2419-2-Figure7.3-1.png",
        "caption": "Fig. 7.3 Layout of the parametric prototype: (a) guide, (b) moving magnet, (c) preload magnets, (d) copper windings, (e) reels, (f) spacer, (g) endstroke bumpers, (h) inner lids, (i) outer lids, (j) chuck jaws, (k) connection bolts",
        "texts": [
            " In the case of smaller magnets (Fig. 7.2b) it is possible to notice a wider dispersion of experimental data and the agreement between magnetic model and experimental data is lower. The main reason is probably the strongest influence of friction between chucks and guide and imprecision in the structure when the applied loads are smaller. The main purpose of this prototype is to test many different configurations in order to validate the simulation model and to experimentally verify its predictions. The prototype (Fig. 7.3) is completely dismountable and most of its components are tunable and replaceable. The guide is made of PMMA with the aim of make it possible to see, end eventually acquire, the moving magnet dynamic if the device is not mounting coils (thus, without most of damping). The external diameter of the guide is constrained by the corresponding interface on the inner lids and itself constrains the diameter of the floating magnet. Otherwise the lengths of the guide and of the moving magnet are not so strictly constrained"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002915_iros.2011.6094845-Figure5-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002915_iros.2011.6094845-Figure5-1.png",
        "caption": "Fig. 5. The 2D model of a deflection sensor based on the ray tracing. Every light ray carries an intensity dependent on the emission angle \u03b1 and the angular intensity profile of a light source fsource(\u03b1); while every light ray perceived by a light sensor is weighted according to angle \u03b3 and an angular light sensitivity profile fsensor(\u03b3). The model is based on determining the angular ranges for the light emission angles (\u03b1dmin, \u03b1dmax) and light reception angles (\u03b3dmin, \u03b3dmax) for given geometrical arrangement of the source-sensor (d1, d2, h) around the point of deflection and angle \u03b2, following (3).",
        "texts": [
            " Therefore, a 2D model is created using ray tracing [13], [14] which captures the relation between the DS characteristics and the signal strength, for every desired deflection \u03b2. The model allows to simulate the performance of the device for a particular source - sensor pair, based on the datasheet information and the geometrical parameters. In addition to the previously described characteristics of the DS, namely, fsource(\u03b1), fsensor(\u03b3) and d1 and d2, the height h of the photosensitive area of the sensor is added, as presented in Fig. 5. Based on these, we calculate four angles for every given deflection \u03b2: \u03b1dmin , minimal ray emission angle, as: \u03b1dmin = 90\u25e6 \u2212 asin ( (d2\u2217sin(\u03b2)) c1 ) \u03b1dmax , maximal ray emission angle, as: \u03b1dmax = 90\u25e6\u2212\u03b1dmin + asin ( (h\u2217sin(90\u2212\u03b3dmin)) c2 ) \u03b3dmin , minimal ray reception angle, as: \u03b3dmin = 180\u2212 \u03b2 \u2212 \u03b1dmin \u03b3dmax , maximal ray reception angle, as: \u03b3dmax = 90\u2212 \u03b3dmin \u2212 asin ( (h\u2217sin(90\u2212\u03b3dmin)) c2 ) where: c1 = \u221a d21 + d22 \u2212 2 \u2217 d1 \u2217 d2 \u2217 cos(\u03b2) c2 = \u221a c21 + h2 \u2212 2 \u2217 c1 \u2217 h \u2217 cos(90\u2212 \u03b3dmin) In the simulation, 9000 rays are emitted from the source Re (from \u03b11 = 0\u25e6 to \u03b19000 = 90\u25e6) with \u2206\u03b1 = 0"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003602_icmech.2013.6518547-Figure8-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003602_icmech.2013.6518547-Figure8-1.png",
        "caption": "Fig. 8 shows parts of the robot. Fig. 9 shows an outline of the robot parts. Fig. 10 shows a cross-sectional view of the air duct line [8].",
        "texts": [
            "56 mm/s in the pipe without lubricant. These results indicate that for two-pattern motion, the robot speed is greater in the lubricated pipe. We attribute this to the reduction in friction between the robot and pipe wall. The friction occurs to allow sagging under the robot\u2019s own weight when the robot extends. In addition, the friction inhibits the robot\u2019s motion when it is extended. This friction is reduced with lubrication, and the speed is therefore greater in the lubricated pipe for each pattern. Fig. 8. Robot parts Fig. 7. Mechanism of the air tube slide robot (4 units) Fig. 19. Vertical pipe Fig. 20. Robot speed in vertical pipe -10 40 90 140 190 0 10 20 30 40 50 60 70 \u6f64\u6ed1\u7121\u3057 \u901f\u5ea62.30 [mm/s] \u6f64\u6ed1\u6709\u308a \u901f\u5ea60.94 [mm/s] Time (s) \u2014 With lubricant \u2015 Without lubricant D is ta nc e (m m ) B. Bent pipe Driving Test Second, we performed an experiment in a bent pipe with or without lubricant using two movement patterns. The inner diameter of the bent pipe was 50 mm, and the outer curvature radius was 90 mm. The bent pipe is shown in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003887_j.procir.2015.06.103-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003887_j.procir.2015.06.103-Figure2-1.png",
        "caption": "Fig. 2. The linear superposition of two screws",
        "texts": [
            " The meshing function is a relation among the geometric and kinematic parameters of the two mating spiral bevel gears to ensure the common tangent mating. The meshing parameters are the geometric and kinematic parameters of the two mating spiral bevel gears. In order to obtain the meshing function, screws 1$ and 2$ are proposed to present the known screw of two mating spiral gears respectively, and the screw 3$ is the sum screw. To calculate the linear combination of the two screws, coordinate system is established and shown in Fig. 2. By setting Z-axis along the direction of the base tangent of two screws, the coordinate axis X and Y and coordinate origin O can freely choose. The axis of 1$ and 2$ are intersect with Z-axis at point A and B . The angle present the angle between two known screws. Based on the screw theory, a screw i$ can be divided into real unit S and dual unit 0 iS of dual vector. According to the algorithms of screw, the real unit and dual unit of sum screw are the sum of original vector and dual vector of two screws respectively. The linear combination of two Screws can be obtained from [9] as 213 $$$ (1) For any screw, can be given from as )3,2,1)(( 0 0 ih iiiiiii SSSSS$ (2) and then the original vector and dual vector of sum Screw are 213 SSS (3) 221102013303 0 3 SSSSSSS hhh (4) Where ih ( 3,2,1i ) is the pitch of these screws, i0S )3,2,1(i is the moment of line. From Fig.2, the line moment can be obtained as )3,2,1(120 ia iiiii SaSrS (5) Where )3,2,1(iai is the distance of screws to origin of coordinate. The unit vector 12a is the common normal of 1$ and 2$ with a unit value. The vector )3,2,1(iir prestent the the distance )3,2,1(iai , respectively. Because the unit vector is the common normal of the two screws, so the sum of two screws is vertical to the common normal. And 0123 aS 21 21 12 SS SS a (6) The pitch of the sum screw can be obtained by the dot product between the 3S and equation (4) and expressed as cos2 sin)(coscos 21 2 2 2 1 2112212 2 22211 2 11 3 SSSS SSaaSShShSShShh (7) where represents the angle between 1$ and 2$ . The cross product of between the 3S and equation (4) can be given as below cos2 cos)(sin)( 21 2 2 2 1 2 22 2 1121122112 3 SSSS SaSaSSaaSShha (8) Then the magnitude, location, pitch and direction of the sum screw can be determined by equations (3), (7) and (8). Particularly, in the nutation drive, the cone vertex of the bevel gear pair is coincidence, and consequently the point A , B and C is coincidence, i. e. 01a and 012 aa . Compare with Fig. 1 and Fig. 2, the screw system of nutation drive are further illustrated in Fig. 3. In nutation gear drive system, bevel gear pair meshes only for pure rotation, and the pitch of the screw is 021 hh . Thus, the pitch of the screw and the distance between sum screw 3$ and coordinate system ),,( ZYXS is 03h and 03a . The meshing between the external and internal spiral bevel gears in the nutation drive can be considered as the external and internal spiral bevel gear meshing with crown gear (an imaginary gear), which has a pitch cone angle of 90 , and the pitch cone is at right angle to its axis"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000177_tro.2006.882921-Figure4-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000177_tro.2006.882921-Figure4-1.png",
        "caption": "Fig. 4. Interval along the side at an acute corner.",
        "texts": [
            " Therefore, is the worst case for side one (here, we choose the point on side one as far as possible from ). For side two, we have . Thus, is the worst change for side two. Finally, we take the maximum of the worst changes for sides one and two. To cover the pie segment defined by , , and , note that the potential changes for each side are monotonic between the bounding lines of the pie segment, and thus, the maximum of the two worst changes bounds the entire pie segment. 3) Convex Acute Angle Corners: We consider an acute angle corner, as shown in Fig. 4. The procedure is the same as the case of obtuse angle corner, i.e., first we find an interval along the side and a depth into the interior, then extend them to make a region where we can guarantee nonexistence of a pivot point. Take along side one. The potential of the point which is closer to point on side two decreases, while it increases after a certain point as the sink is moved from to . Thus, we can find a segment on side two such that the potential change for this segment is zero when the sink is moved from to by ",
            " The bound for the region around a concave acute corner is analyzed in the same way as the concave obtuse case. Fig. 6 shows the geometry, although we omit the details for brevity. These bounds provide the means for an algorithm to peel off the boundary, as given in Algorithm 2. Algorithm 2: Guaranteeing no pivot points near the boundary Require Ensure perimeter {if not, reduce } if CORNER then if OBTUSE then % Step: where is found by (Fig. 2). % Region: if Convex then Side 1: where satisfies (Fig. 2) Side 2: (Fig. 2) else {Concave} Side 1: (Fig. 3) Side 2: (Fig. 3) end if else {ACUTE} % Step: (Fig. 4) % Region: if Convex then Side 1: (Fig. 5) Side 2: (Fig. 5) else {Concave} Side 1: (Fig. 6) Side 2: (Fig. 6) end if end if else {SIDE} (Fig. 1) end if For a given object, we have checked each vertex and the boundary, and peeled off the boundary all around the object with a finite thickness everywhere, except points on the boundary where the force is zero. The analysis on the existence of a finite resolution (away from critical points) is guaranteed, and suggests a constructive algorithm. The arguments so far prove the existence of a finite resolution where we can guarantee no pivot point exists, except in regions near actual critical points, i",
            " A more intuitive way to look at this idea is that a picture frame hanging on a nail hangs down, as long as the pivot point is not coincident with the center of gravity. We have constructed a hardware testbed capable of demonstrating this sensorless manipulation algorithm. An aluminum plate with a fine grid of small holes powered by wall air provides the air bearing, and a plexiglass plate on standoffs with tapped holes is the \u201cair palm\u201d which is placed approximately over an object, attracts it to its center, and orients it without sensors using a predesigned air field using flow sinks (see [25, Fig. 4]). Hand valves individually activate the flow sinks which are powered by a shop-vac (see Fig. 9). The system manipulates small pieces of sandpaper-covered plexiglass (for additional surface drag). The damping in the actual system is quite low, thus it takes quite a long time for the object to settle down at equilibrium. To increase the damping in the sense of dissipating energy, we generate intermittent \u201chopping\u201d of an object by applying a pulsed flow in the air bearing. To prevent the intervening of the lifting air with the manipulating flow, the flat object is fixed to a thin disk large enough to cover all the sinks"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002194_bf02116431-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002194_bf02116431-Figure1-1.png",
        "caption": "Fig. 1. Externally pressurized bearing.",
        "texts": [
            " It is shown that the load capacity due to viscous forces increases with an increase in the Hartmann number while the load capacity due to inertia forces decreases with an increase in the Har tmann number and the inertia forces can be neglected at large values of the Har tmann number. The externally pressurized bearing with an axial current is considered in \u00a7 4. It is shown that the load capacity due to inertia forces in this case is independent of axial current and an additional load proportional to the square of the axial current is obtained by applying the axial current. \u00a7 2. Bearing with axial magnetic field. Referring to fig. 1, the equations of conservation of mass and momentum are given by 1) -~ (ru) -=-- 0, (1) pu ~r -- dr + # ~rr -~r (ru) + Oz 2 j Equations (1) and (2) result into ~2u aB~u 1 dp pr ~u - - + - - - - ( 3 ) 8z 2- t~ # dr # & We will obtain an approximate solution of the non-linear equation (3), with the boundary conditions u = 0 at z = ~ h , by the iteration method used by Jackson and Symmons 6) and by the averaged inertia method suggested by Slezkin and TargT) and used by Osterle and SaibelS). (A) M e t h o d of i t e r a t i o n "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002371_c1ay05023k-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002371_c1ay05023k-Figure1-1.png",
        "caption": "Fig. 1 Colorimetric Cu2+ sensing assembly.",
        "texts": [
            "9 250 mL 10 5 mol L 1 CdTe QDs were added into a quartz cuvette containing Anal. Methods, 2011, 3, 1471\u20131474 | 1471 D ow nl oa de d by B ro w n U ni ve rs ity o n 21 J ul y 20 12 Pu bl is he d on 2 1 Ju ne 2 01 1 on h ttp :// pu bs .r sc .o rg | do i:1 0. 10 39 /C 1A Y 05 02 3K different concentrations of Cu2+ solution, and then pure water was added to keep the final volume at 2.5 mL. Under the different Cu2+ concentrations, different fluorescence quenching of CdTe QDs occurred. A glass slide immobilized Ru(bpy)3Cl2 layer was inserted in the cuvette as shown in Fig. 1 and an LEDwith amaximum emission wavelength of 396 nm was used as the emission source. After mixing the solution for 3 min, the resulting colors were recorded using the Nikon camera. Generally, inorganic QDs are of narrow and symmetric photoluminescence, tunable color, wide excitation wavelength and low cost to synthesize. Water-soluble CdTe QDs are ideal fluophors due to their unique optical properties, low cost and ease of scale-up, and have beenwidely used formulticolor visualized sensing and encoding"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000427_j.ijsolstr.2007.12.004-Figure6-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000427_j.ijsolstr.2007.12.004-Figure6-1.png",
        "caption": "Fig. 6. Clamped\u2013contact sub-domain.",
        "texts": [
            " The deformation pattern as described in this section and as observed in Fig. 3(1) is called deformation pattern 1 in this paper. In this sub-section, we describe in detail the calculation procedure adopted in analyzing the elastica deformation. In other sub-domains as to be discussed later, we skip the detailed calculations and list only the main characteristics pertaining to the associated boundary conditions. Near the clamped end A, the elastica may contact the outer wall or the inner wall. Fig. 6 shows two possible scenarios; the elastica may contact the outer wall r = 1 at point H, or the elastica may contact the inner wall r = g at point J. The locations of point H and J may be denoted by the angles w\u00fe H 2 and w J 2 as measured from the central radius, respectively. The superscripts \u2018\u2018+\u2019\u2019 and \u2018\u2018 \u2019\u2019 signify that the contact points are on the outer and inner walls, respectively. These two sub-domains are listed as cases (b) and (c) in Table 1. In the case when the elastica is in point contact with the walls, the bending moments at points H and J are denoted as MH and MJ, respectively"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003877_apcase.2015.45-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003877_apcase.2015.45-Figure1-1.png",
        "caption": "Fig. 1. The twin rotor MIMO system",
        "texts": [
            " The rest of this article is organized as follows: Section II describes the mathematical model of TRMS, Section III presents the design of the linear and the nonlinear PID controllers, in Section IV describe a nonlinear observer design, for both vertical and horizontal directions. Then, in Section V presents the analysis results and performance evaluation of control algorithms and dynamic nonlinear observers is performed. Finally, conclusions are summarized in Section VI. II. TRMS MODELING The TRMS mechanical unit is composed by two rotors placed on a beam, due to its design allows it to operate as a real helicopter. The beam is attached to a counterweight arm in the pivot; it determines a stable equilibrium position as shown in Fig.1 [1]. The entire TRMS system dynamics could be approximated in state space form as indicated in [12 -13], as follows: 978-1-4799-7588-4/15 $31.00 \u00a9 2015 IEEE DOI 10.1109/APCASE.2015.45 214 2 21 1 1 1 1 1 1 1 1 2 1 1 1 1 1 1 1 0.0326sin sin(2 )( ) 2 cos( ) cos( ) v v v g v v v h gy gy v v h v h h h d dt Ma bd dt I I I I B k k a b I I I d dt 1(cos ( 12 22 2 2 2 1 1 1 1 2 2 2 2 2 10 1 1 1 1 11 11 20 2 2 2 2 21 21 1.75 1.75h h h c c Ba bd k a k b dt I I I I I T kd u dt T T T kd u dt T T (1) The output is described by: [ ]hv Ty (2) Donde, 1 2 1 2 : Pitch (elevation) angle : Yaw (azimuth) angle : Momentum of main rotor : Momentum of tail rotor : Control signal of main rotor : Control signal of tail rotor v h u u In advance, a state representation 1 2 3 4 5 6[ ]x x x x x x replaces the corresponding physical sign of the TRMS 1 2[ ] v v h h "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000587_2007-01-2232-Figure21-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000587_2007-01-2232-Figure21-1.png",
        "caption": "Fig. 21 Vibration Mode of Longitudinal Transmission",
        "texts": [
            " In order to reduce the 24th-order electromagnetic force of revolution that is generated in MG2, the permanent magnets were arranged in a V shape and the angle of the magnet arrangement was optimized, as in the THS II for FWD vehicles.(3) In order to improve the radiated noise characteristics of the transmission, FEM analysis was used to analyze in detail the vibration modes that contribute to radiated noise. The results indicated that two vibration modes exist. One vibration mode couples the bending resonance of the transmission case with a resonance in which the MG1 and MG2 rotors serve as mass elements and the support bearings serve as spring elements. The other vibration mode, shown in Fig. 21, couples the bending resonance of the transmission case. To dissipate these resonant frequencies, the resonant frequency of the transmission case was changed, thereby improving the radiated noise characteristics. The radiated noise characteristics were also improved by installing a dynamic damper on the back end of the transmission, where the amplitude was high, and by adding ribs in radiating areas of the transmission case. In the LS600h, a sound proof cover was added on the surface of the transmission"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002822_c1sm05282a-Figure6-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002822_c1sm05282a-Figure6-1.png",
        "caption": "Fig. 6 Surfactant balance in a control box of typical sizeR located at the trailing end of the slug, shown in dotted line. fin is the incoming flux of surfactant through the lubricating film of water and fout is the outgoing diffusive flux.",
        "texts": [
            " 5b), the motion is initially much slower (due to the large oil viscosity, eqn (2)), and the amount of This journal is \u00aa The Royal Society of Chemistry 2011 surfactant transported through the much thinner wetting film can be small enough to get simultaneously diluted by diffusion in fresh water, making possible a continuous motion. Then the velocity should be given by eqn (2), where Dg < DgM is the surface tension difference between the slug ends in this stationary regime. We can compare the amounts of surfactant carried by the film and lost by diffusion in a control box of typical size R located at the trailing end of the slug, as shown in Fig. 6. On the one hand, if the slug moves continuously at a velocityV (Fig. 5b), the water film of thickness h transports per unit time a flux fin CRhV. On the other hand, the diffusive flux fout of surfactant in a tube of characteristic radius R scales as CDR, where D is the diffusion coefficient of SDS in water, typically of order 10 10 m2 s 1. The Peclet number Pe \u00bc fin/fout hV/D compares the magnitude of these two fluxes. If this ratio is smaller than unity, diffusion is dominant and we expect a uniform motion"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002794_iros.2011.6094984-Figure4-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002794_iros.2011.6094984-Figure4-1.png",
        "caption": "Fig. 4: An illustrative sketch of a robotic serial manipulator. The unknown external wrench wext can be computed through the information provided by the six-axis force/torque sensor mounted in the middle of link 1.",
        "texts": [
            " Taking in account this relationship, the torque of the shoulder pitch joint can be expressed as: 4 4 4 5 5 5( ) ( )pitch rollc s o c s o\u03c4 \u03c4= \u22c5 + + \u22c5 + \u2212 (8) Also in this case two calibration coefficients (c4, c5) and two offsets (o4, o5) are required to obtain a calibrated measurement. These parameters are determined with the procedure described in section III. C. Computed joint torques approach A model based approach [13] has been developed for the iCub humanoid robot, which exploits a graph formulation to compute the inverse dynamics of the system. The method makes use of graph theory to introduce in the computation of the dynamics the measurements from the six-axis force/torque sensors and the contact point information retrieved from the iCub artificial skin. Figure 4 shows an illustrative example of a robotic serial mechanism equipped with a force/torque sensor in one of its links. Through this sensor it is possible to measure both the components (forces and moments) of the wrench which are due to the internal dynamics of the robot, but also a single external wrench applied distally with respect to the sensor. Let us represent the serial mechanism of Figure 4 with an oriented graph formulation, in which each vertex vi represents the i-th link and each outgoing edge represents the reference frame constrained to the i-th link (Figure 5a). Let us consider that the unknown wrenches of this robotic mechanism are the externally applied wrench wext=(fi, \u00b5 i) and the reaction forces on the base w0=(f0, \u00b50), both graphically represented by a white rhombus (\u25ca) . Let us also consider the known quantities ws=(fs, \u00b5 s), graphically indicated by a black rhombus (\u2666), which correspond to the wrench measured by the force/torque sensor"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000264_3-540-36224-x_4-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000264_3-540-36224-x_4-Figure1-1.png",
        "caption": "Fig. 1. Underactuated robots appear in a variety of environments. From left to right, a planar vertical take-off and landing (PVTOL) aircraft model, a horizontal model of a blimp and the snakeboard.",
        "texts": [],
        "surrounding_texts": [
            "The research in Robotics is continuously exploring the design of novel, more reliable and agile systems that can provide more efficient tools in current applications such as factory automation systems, material handling, and autonomous robotic applications, and can make possible their progressive use in areas such as medical and social assistance applications. Mobile Robotics, primarily motivated by the development of tasks in unreachable environments, is giving way to new generations of autonomous robots in its search for new and \u201cbetter adapted\u201d systems of locomotion. For example, traditional wheeled platforms have evolved into articulated devices endowed with various types of wheels and suspension systems that maximize their traction and the robot\u2019s ability to move over rough terrain or even climb obstacles. The types of wheels that are being employed include passive and powered castors, ball-wheels or omni-directional wheels that allow a high accuracy in positioning and yet retain the versatility, flexibility and other properties of wheels. A rich and active literature includes (i) various vehicle designs [38,41,44,46], (ii) the automated guided vehicle \u201cOmniMate\u201d [2], (iii) the roller-walker [15] and other dexterous systems [17] that change their internal shape and constraints in response to the required motion sequence, and (iv) the omni-directional platform in [19]. A. Bicchi, H.I. Christensen, D. Prattichizzo (Eds.): Control Problems in Robotics, STAR 4, pp. 59--74, 2003 Springer-Verlag Berlin Heidelberg 2003 Other types of remotely controlled autonomous vehicles that are increasingly being employed in space, air and underwater applications include submersibles, blimps, helicopters, and other crafts. More often than not they rely on innovative ideas to affect their motion instead of on classic design ideas. For example, in underwater vehicle applications, innovative propulsion systems such as shape changes, internal masses, and momentum wheels are being investigated. Fault tolerance, agility, and maneuverability in low velocity regimes, as in the previous example systems, are some of the desired capabilities. A growing field in Mobile Robotics is that of biomimetics. The idea of this approach is to obtain some of the robustness and adaptability that biological systems have refined through evolution. In particular, biomimetic locomotion studies the periodic movement patterns or gaits that biological systems undergo during locomotion and then takes them as reference for the design of the mechanical counterpart. In other cases, the design of robots without physical counterpart is inspired by similar principles. Robotic locomotion systems include the classic bipeds and multi-legged robots as well as swimming snake-like robots and flying robots. These systems find potential applications in harsh or hazardous environments, such as under deep or shallow water, on rough terrain (with stairs), along vertical walls or pipes and other environments difficult to access for wheeled robots. Specific examples in the literature include hyper-redundant robots [13,16], the snakeboard [32,40], the G-snakes and roller racer models in [26,27], fish robots [23,25], eel robots [21,36], and passive and hopping robots [18,35,42]. All this set of emerging robotic applications have special characteristics that pose new challenges in motion planning. Among them, we highlight: Underactuation. This could be owned to a design choice: nowadays low weight and fewer actuators must perform the task of former more expensive systems. For example, consider a manufacturing environment where robotic devices perform material handling and manipulation tasks: automatic planning algorithms might be able to cope with failures without interrupting the manufacturing process. Another reason why these systems are underactuated is because of an unavoidable limited control authority: in some locomotion systems it is not possible to actuate all the directions of motion. For example, consider a robot operating in a hazardous or remote environment (e.g., aerospace or underwater), an important concern is its ability to operate faced with a component failure, since retrieval or repair is not always possible. Complex dynamics. In these control systems, the drift plays a key role. Dynamic effects must necessarily be taken into account, since kinematic models are no longer available in a wide range of current applications. Examples include lift and drag effects in underwater vehicles, the generation of momentum by means of the coupling of internal shape changes with the environment in the eel robot and the snakeboard, the dynamic stability properties of walking machines and nonholonomic wheeled platforms, etc. Current limitations of motion algorithms. Most of the work on motion planning has relied on assumptions that are no longer valid in the present applications. For example, one of these is that (wheeled) robots are kinematic systems and, therefore, controlled by velocity inputs. This type of models allows one to design a control to reach a desired point and then immediately stop by setting the inputs to zero. This is obviously not the case when dealing with complex dynamic models. Another common assumption is the one of fully actuation that allows to decouple the motion planning problem into path planning (computational geometry) and then tracking. For underactuated systems, this may be not possible because we may be obtaining motions in the path planning stage that the system can not perform in the tracking step because of its dynamic limitations. Furthermore, motion planning and optimization problems for these systems are nonlinear, non-convex problems with exponential complexity in the dimension of the model. These issues have become increasingly important due to the high dimensionality of many current mechanical systems, including flexible structures, compliant manipulators and multibody systems undergoing reconfiguration in space. Benefits that would result from better motion planning algorithms for underactuated systems. From a practical perspective, there are at least two advantages to designing controllers for underactuated robotic manipulators and vehicles. First, a fully actuated system requires more control inputs than an underactuated system, which means there will have to be more devices to generate the necessary forces. The additional controlling devices add to the cost and weight of the system. Finding a way to control an underactuated version of the system would improve the overall performance or reduce the cost. The second practical reason for studying underactuated vehicles is that underactuation provides a backup control technique for a fully actuated system. If a fully actuated system is damaged and a controller for an underactuated system is available, then we may be able to recover gracefully from the failure. The underactuated controller may be able to salvage a system that would otherwise be uncontrollable."
        ]
    },
    {
        "image_filename": "designv11_12_0003713_978-1-4614-2419-2-Figure31.5-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003713_978-1-4614-2419-2-Figure31.5-1.png",
        "caption": "Fig. 31.5 Rendering of shaft assembly and instrumentation",
        "texts": [],
        "surrounding_texts": [
            "Both torsional and lateral loads can be applied to a shaft during testing. The test rig is depicted in Fig. 31.2 and consists of a"
        ]
    },
    {
        "image_filename": "designv11_12_0002794_iros.2011.6094984-Figure3-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002794_iros.2011.6094984-Figure3-1.png",
        "caption": "Fig. 3: Top: the sensorised mechanical parts. Finite element simulations are performed in order to choose the optimal location of the strain gauges. Bottom: each couple of sensors is configured as an half Wheatstone bridge.",
        "texts": [
            " This is obtained through the placement of semiconductor strain gauges (SSGs) on ad-hoc redesigned mechanical parts. More in particular, the location for the four joint torque sensors was chosen to be as follows: 1) Elbow joint: this is the simplest case, since the elbow joint takes no part in the coupling mechanism of the shoulder and is controlled by its own independent motor. The torque of this joint is measured by two SSGs mounted on a spoke-structure, which has been conveniently introduced on a modified mechanical part (Figure 3). The two strain gauges are configured as an half Wheatstone bridge (the other half of the bridge is directly mounted on the acquisition board) and the output signal is acquired by a 16-bits ADC converter. The resulting joint torque \u03c4elbow can be thus obtained from the raw measurements of the sensor (s1) accordingly to the following formula: 1 1 1( ) elbow c s o\u03c4 = \u22c5 + (5) where c1 represents the calibration coefficient required to convert from arbitrary units (i.e. the 16-bits integers returned by the analog-to-digital converter) to metric measurements (Nm), and o1 represents the sensor offset which has to be canceled out so that the sensor measurement is zero when no torque is applied to the joint"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001522_detc2009-86119-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001522_detc2009-86119-Figure1-1.png",
        "caption": "Figure 1. Machine-tool setting for pinion tooth-surface finishing on cradle-type hypoid generator",
        "texts": [
            " The pinion is the driving member. The convex side of the gear tooth and the mating concave side of the pinion tooth are the drive sides. The modifications are introduced into the pinion tooth-surface by applying machinetool setting variations in pinion tooth generation. As a result of these modifications the spiral bevel gear pair becomes mismatched and point contact of the meshed teeth surfaces appears instead of line contact. The machine-tool setting used for pinion teeth finishing is given in Fig. 1. The concave side of pinion teeth is in the coordinate system 1K (attached to the pinion, Fig. 1) defined by the following system of equations )T( Tp1p2p3 )1( 1 1 1 rMMMr rr \u22c5\u22c5\u22c5= (1a) 0)T( 1m )1,T( 1m 11 =\u22c5 ev rr (1b) where ( )1 1 T Tr r is the radius vector of tool-surface points, matrices p1M , p2M , and p3M provide the coordinate transformations from system 1TK (rigidly connected to the cradle and headcutter 1T ) to system 1K (rigidly connected to the being generated pinion). Equation (1b) describes mathematically the generation of pinion tooth-surface by the head-cutter [28]. The Downloaded From: http://proceedings.asmedigitalcollection.asme.org/pdfaccess.ashx?url=/data/conferences/idetc/cie2009/70876/ on 03/25/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use 4 Copyright \u00a9 2009 by ASME matrices and vectors of system of Eqs. (1a) and (1b) are defined as it follows. The surface of the head-cutter used for finishing the concave side of pinion teeth is in the coordinate system 1TK (attached to the tool) defined by the following equation (based on Fig. 1):             \u03b8\u22c5\u03b1\u22c5+ \u03b8\u22c5\u03b1\u22c5+ \u2212 =\u03b8 1 sin)tgur( cos)tgur( u ),u( 1T 1T)T( T 1 11 1 r r (2) On the basis of Fig. 1 and Eq. (2), for the relative velocity vector )1,T( 1m 1v r of tool 1T to the pinion, and for the unit normal vector of the tool-surface, )T( 1m 1e r , it follows ( ) ( ) ( )[ ]           \u2212\u03b3\u22c5\u2212\u2212\u03b3\u22c5\u22c5 \u03b3\u22c5+\u22c5\u2212 \u03b3\u22c5+\u22c5 \u22c5\u03c9= )T( 1m1 1T( 1m1 )T( 1mgp 1 )T( 1mgp )T( 1m 1 )T( 1mgp )c()1,T( 1m 111 11 1 1 ycoscxsinyi singziz cosgzi v r (3)             \u03b8\u22c5\u03b1 \u03b8\u22c5\u03b1 \u03b1 \u22c5=\u22c5= 0 sincos coscos sin 1 1 1 p1 )T( 1Tp1 )T( 1m 11 MeMe rr (4) where ( ) ( )1 1 1 T Tp1 T 1m rMr rr \u22c5= . For matrices p1M , p2M , and p3M , providing the coordinate transformations, on the basis of Fig. 1, it follows 1T cppcpcp cppcpcp 1Tp11m 1000 sinesincos0 cosecossin0 0001 rrMr rrr \u22c5             \u03c8\u22c5\u2212\u03c8\u03c8 \u03c8\u22c5\u03c8\u2212\u03c8 =\u22c5= (5) 1m 111 111 1mp201 1000 g100 sincf0cossin cosc0sincos rrMr rrr \u22c5             \u03b3\u22c5\u2212\u2212\u03b3\u03b3 \u03b3\u22c5\u2212\u03b3\u2212\u03b3 =\u22c5= (6) 01 11 11 01p31 1000 0cos0sin p010 0sin0cos rrMr rrr \u22c5             \u03c8\u03c8\u2212 \u2212 \u03c8\u03c8 =\u22c5= (7) while, for the traditional cradle-type hypoid generators ( )0cpcpgp101 i \u03c8\u2212\u03c8\u22c5+\u03c8=\u03c8 . Angles 10\u03c8 and 0cp\u03c8 correspond to the generation of the \u201cinitial\u201d contact point on the pinion tooth-surface. Because of the mismatch of the gear pair, only in one point of the path of contact, called as the \u201cinitial\u201d contact point, the basic mating equation of the contacting pinion and gear tooth-surfaces is satisfied, producing the correct velocity ratio based on the numbers of teeth",
            " The corresponding coordinate transformations are defined by the following equations:             \u03b8\u22c5\u03b1\u22c5+ \u03b8\u22c5\u03b1\u22c5+ \u2212 \u22c5             \u03be\u22c5\u2212\u03be\u22c5\u2212\u03be\u03be \u03be\u22c5+\u03be\u22c5\u2212\u03be\u2212\u03be =\u22c5= 1 sin)tgur( cos)tgur( u 1000 0100 cosYsinX0cossin sinYcosX0sincos CNC1T CNC1T c 1 1 T1c rMr rr (10) cc11 rrMr rrr \u22c5             \u03b7\u22c5\u2212\u03b7\u03b7\u2212 \u03b7\u22c5\u2212\u03b7\u03b7 =\u22c5= 1000 cosZcos0sin 0010 sinZsin0cos (11) The coordinate transformation from system 1TK to system 1K , performs the following equation: T1c11 rMMr rr \u22c5\u22c5= (12) 10 1y , y Cy Z T1xY Cx 10x T1y CO 1OT1O X 1x T1z 1z 10 cz ,z Head-Cutter Pinion Figure 3. Machine-tool setting for pinion tooth-surface finishing on CNC generator The location and the orientation of the tool with respect to the pinion/gear are given in coordinate systems that are represented for a conventional, cradle-type generator (Fig. 1). The goal is to develop the algorithm for the execution of motions of the CNC machine using the relations in the cradletype machine. Since the tool is a rotary surface and the pinion/gear blank is related to a rotary surface, too, it is necessary to ensure that the relative position of the two axes, 1Tx and 10y , and the axial relative position of the head cutter and the pinion, to be the same in both machines. The origo of Downloaded From: http://proceedings.asmedigitalcollection.asme.org/pdfaccess"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002570_016918610x552187-Figure6-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002570_016918610x552187-Figure6-1.png",
        "caption": "Figure 6. Conical object grasped by three fingers.",
        "texts": [
            " Eigenvalues and eigenvectors for the sliding contact grasp Eigenvalue Eigenvector \u03bbfs 1 = 1000 N/m [1/2,0, \u221a 3/2,0,0,0] \u03bbfs 2 = h22 = \u22122000fx 15\u22122fx N/m [0,1,0,0,0,0] \u03bbfs 3 = \u22122000fx 15\u22122fx N/m [\u2212\u221a 3/2,0,1/2,0,0,0] \u03bbfs 4 = h44 = 0 Nm/rad [0,0,0,1,0,0] \u03bbfs 5 = h55 = 0 Nm/rad [0,0,0,0,1,0] \u03bbfs 6 = h66 = 0 Nm/rad [0,0,0,0,0,1] Figure 5. Eigenvalues \u03bbfs 2 and \u03bbfs 3 . Human beings mainly experience the property of eigenvalues \u03bbfs 2 and \u03bbfs 3 , and so expect that this frictionless grasp is always unstable. Next, we explore the case of a conical object grasped by three fingers as shown in Fig. 6. Fingers 1 and 2 grasp the conical surface defined by (39), and the initial contact points are located at (u1, v1) = (a, \u03b8) and (u2, v2) = (a, \u03b8 + \u03c0), respectively: opLo = 1\u221a 2 [u cosv,u sinv,u]T (0 < u < 0.02 \u221a 2,0 v < 2\u03c0). (39) In this example, we use a = (50 \u221a 2)\u22121 and \u03b8 = 0. Finger 3 grasps the top surface defined by (40) and the initial contact point is located at (u3, v3) = (0,0): opLo = [u, v,0.02]T (u2 + v2 < 0.022). (40) The metric tensor MC, curvature KC and torsion TC are calculated as shown in Table 5. The grasping force in fi is assigned with: f1 = f2 = [fx,0,0]T (41) f3 = [\u221a2fx,0,0]T. D ow nl oa de d by [ U ni ve rs ity o f N ew E ng la nd ] at 0 7: 08 1 3 Ja nu ar y 20 15 T. Yamada et al. / Advanced Robotics 25 (2011) 447\u2013472 461 Table 5. Geometrical parameters of the grasp shown in Fig. 6 Curvature KC (m\u22121) Torsion TC (m\u22121) Metric tensor MC Contact point 1 [ 0 0 0 50 \u221a 2 ] [0,50 \u221a 2] [ 1 0 0 1/100 ] Contact point 2 [ 0 0 0 50 \u221a 2 ] [0,50 \u221a 2] [ 1 0 0 1/100 ] Contact point 3 [ 0 0 0 0 ] [0,0] [ 1 0 0 1 ] Finger 1, 2, 3 [ 200 0 0 200 ] [0,0] [ 1/200 0 0 1/200 ] First, we evaluate the case when all the three contact points are sliding contact. From Section 4.3, the elements of H are given by: h11 = 500, h15 = h51 = 10, h55 = 200 + 5(2 + 5 \u221a 2)fx \u2212 8f 2 x 1000 h22 = 2000fx \u22125 \u2212 10 \u221a 2 + 2fx , h24 = h42 = (20 \u2212 5 \u221a 2 + 2 \u221a 2fx)fx 5 + 10 \u221a 2 \u2212 2fx (42) h44 = {225 \u221a 2 + 100 \u2212 (130 \u221a 2 + 20)fx + 8f 2 x }fx 1000(5 + 10 \u221a 2 \u2212 2fx) h33 = 1000",
            " The other elements are all zero. The eigenvalues and eigenvectors of H are calculated as shown in Table 6. The eigenvalues \u03bbfs 3 and \u03bbfs 4 are obtained from: \u03bb2 \u2212 (h11 + h55)\u03bb + h11h55 \u2212 h15h51 = 0, (43) and the eigenvalues \u03bbfs 5 and \u03bbfs 6 are obtained from \u03bb2 \u2212 (h22 + h44)\u03bb + h22h44 \u2212 h24h42 = 0. (44) D ow nl oa de d by [ U ni ve rs ity o f N ew E ng la nd ] at 0 7: 08 1 3 Ja nu ar y 20 15 462 T. Yamada et al. / Advanced Robotics 25 (2011) 447\u2013472 Table 6. Eigenvalues and eigenvectors of the sliding contact grasp shown in Fig. 6 Eigenvalue Eigenvector \u03bbfs 1 = h33 = 1000 N/m [0,0,1,0,0,0] \u03bbfs 2 = h66 = 0 Nm/rad [0,0,0,0,0,1] \u03bbfs 3 mainly x-directional translation \u03bbfs 4 mainly y-directional rotation \u03bbfs 5 mainly y-directional translation \u03bbfs 6 mainly x-directional rotation (a) (b) (c) (d) Figure 7. Eigenvalues of the sliding contact grasp shown in Fig. 6. Figure 7 shows the last four eigenvalues. The units (N/m) and (Nm/rad) are mixed together in these eigenvalues, because the corresponding eigenvectors are not strictly decomposed into translation and rotation. Hence, the units along the vertical axis are not shown. As each contact point is sliding contact, the object can freely rotate around the z-axis. Hence, we have \u03bbfs 2 = 0 and rotation in the z-direction is always unstable. From Fig. 7b and d, eigenvalues \u03bbfs 4 and \u03bbfs 6 have the maximum values"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002248_690114-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002248_690114-Figure1-1.png",
        "caption": "Fig. 1 - Journal position for unidirectional load",
        "texts": [
            " An automotive engineer, aided by a com puter, can use the included computation procedure to get a reasonable picture of bearing performance. He can discover the effect of changes in speed, power, dimensions, and oil on potential trouble spots, and he can see the difference in behavior of main bearings, crankpin bearings, and wrist pin bushings. Typical orbit diagrams, based on this procedure, are included for comparison with photographs of test bearings for which the journal paths were drawn. 548 techniques intended for day-to-day use are concerned with the unidirectional, constant load situation described in Fig. 1 . The designer must usually be content with the adaptation of one of these techniques for approximate solutions to the dynamic case illustrated in Fig. 2. This gross approach to predicted bearing performance is not consistent with motor industry trends which find sleeve bearing technology being strained to its limits and marginal bearing operation increasing. Recently, several methods have been advanced for a more direct approach to the prob lems found in reciprocating engines. They include hand, graphical, and computer solutions",
            " Co-operative participation with other firms in this development is invited by Clevite Corp. The journal orbit analysis offers prospects for achieving a greatly im proved journal bearing design tool that can be applied on a routine basis. 560 J . M. ROSS AND R. R. SLAYMAKER APPENDDC A The action of a piston engine journal in its bearing is complicated because so many things happen at once, but if the variables are examined one at a time the total process is much less confusing. Consider, for example, the steady unidirectional load shown in Fig. 1. Here the load is fixed in magnitude and direction, the bearing does not rotate, and the shaft turns at a constant speed. An equilibrium position is reached with the line of centers making an angle 0 with the load line, the shaft center displaced a small distance e from the bearing center and the load is supported by the oil film pressure. Now imagine that the bearing as well as the shaft in Fig. 1 rotates at a constant speed with respect to the load. This will cause both the attitude angle 0 and the eccentric posi tion e to change but the system is still in static equilibrium. In a reciprocating engine, however, the magnitude of the load, its direction with respect to the bearing and the shaft speed with respect to the bearing all change with time. This presents a situation shown schematically in Fig. 2 in which all speeds are referred to the load direction as being fixed. The fact that the shaft speed with respect to the load w g^ L > the bearing speed with respect to the load w and the load B/L magnitude are continually changing at known rates means that there must be corresponding changes in 0 and e in order that the oil film pressure can support the load"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001196_0094-114x(72)90004-3-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001196_0094-114x(72)90004-3-Figure2-1.png",
        "caption": "Figure 2. Link i in obl ique project ion.",
        "texts": [
            " For convenience, the revolute joint is assumed to be at the input joint (joint 1) and the fixed link is looked at as link 4. Each link i of the chain has a length, a~, which is the shortest distance between its two axes and a twist angle, m, which is a measure of the relative directions of its two axes. The rotations, 0i, and the translations, s~, define the relative positions of the links ( i - 1 and i) which are connected at joint i. The quantities are measured with respect to coordinate systems attached to the four links. Coordinate system i + 1 is fixed to link i. Figure 2 illustrates link i, its two parameters, a~ and a~, and the attached i + I coordinate system. The two axes of link i are represented by the unit vectors z~ and z~+l. The unit vector xf+~ is perpendicular to both zi and z~+1 and can be written as z~ X z~+t (1) xi+~ = [zi \u00d7 zi+t[ if z~ and z~+t are not parallel. If the two axes of link i are parallel, x~_~ is directed from z~ to z~+~. The unit vector y~,~ is defined as Yi+l = Z~+l \u00d7 xi+l. (2) The twist angle, a~, is measured from +z~ to +zi\u00f7, in a positive sense about \u00f7x~+, according to the right-hand rule"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003357_icc.2012.6364935-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003357_icc.2012.6364935-Figure1-1.png",
        "caption": "Figure 1. Sensors embeded in a smart device.",
        "texts": [
            " But, as far as we know there is not any smart collaborative system like the one described in this paper. III. SYSTEM PROPOSAL In this section we describe the embedded sensors used in our system and the decision algorithm procedure. Our system is proposed to be implemented in a group of disabled or elderly people. Each person will carry out a smart device that integrates multiple transducers, such as proximity/light sensor, magnetic sensor, acoustic sensor or microphone, compass, gyroscope, accelerometer, GPS locator or others, as a barometer (see Figure 1). Initially, we define the threshold for each sensor, from which, the algorithm will decide when there is an emergency situation or not. Compass: Electronic compass bases its operation on the uptake of terrestrial electromagnetic fields, always marking the magnetic north, which does not correspond to the geographic North. The difference between both norths is known as magnetic declination. Its value is expressed in degrees. This sensor let us know if a person walks in the same direction as the rest of the group"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002757_icsens.2010.5690907-Figure5-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002757_icsens.2010.5690907-Figure5-1.png",
        "caption": "Fig. 5. Flat cylinder connected to base in middle of unit by using bearing and rotary potentiometer sensing angle of the flat cylinder",
        "texts": [
            " Then we confirmed that we could readout stiffness of a wire and started to control a joint stiffness of the humanoid robot, Kenzoh to achieve interaction with people and human life support for the future. II. STRUCTURE OF ROTARY NONLINEAR SPRING UNIT I will explain rotary mechanism that is the main original point of these nonlinear spring units we have developed [14] (I am waiting for a Japanese patent grant [15]). A cylinder is placed in the middle of the unit and connected to a base by upper and lower bearings to rotate smoothly. A rotary potentiometer is attached on top of the cylinder to measure angle of the cylinder rotation (Fig. 5). Upper part of the cylinder is a pulley through which a wire passes (Fig. 6 right). Lower part are connected to the base by using compression-type springs (Fig. 6 left). Cross section of the spring made square to realize high kinetic coefficient in a small size. To avoid bucking of the spring we insert a shaft through the spring, and to move it smoothly we used spring housing with bearings (Fig. 6 left). Upper Section Lower Section Initial State State After Adding Force Fig. 7. Structure change of upper and lower section before and after adding force to wire Next, I will explain the behavior of the unit when we apply tension to the wire which passing through the unit"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000243_jae-2007-772-Figure3-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000243_jae-2007-772-Figure3-1.png",
        "caption": "Fig. 3. The squirrel cage induction motor.",
        "texts": [
            " The common boundaries of the two methods are two circles covering the air-gap. Boundary conditions between the two solutions are as in the following: Br (Analytical) = Br (MEC) H\u03b8 (Analytical) = H\u03b8 (MEC) (7) Figure 2 shows the interface between MEC and Analytical methods where the above boundary conditions are used to couple the two methods. The scalar potentials of nodes in MEC modelling, and constants of the analytical solution are the variables that should be obtained. The considered induction motor is shown in Fig. 3. For simplicity of MEC modelling, the slots and teeth in stator and rotor are shown as sectors of a circle. The torque versus speed calculated from the proposed method is shown and compared with FEM results in Fig. 4. Figures 5 and 6 show the radial and circumferential components of magnetic flux density along with the FEM results in the middle of air-gap, respectively. One can deduce from Figs 4 to 6 that the results of the proposed method compare very accurately with FEM. This paper has presented a new method to couple an analytical magnetic field solution with the magnetic equivalent circuits method"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002790_mesa.2012.6275561-Figure5-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002790_mesa.2012.6275561-Figure5-1.png",
        "caption": "Fig. 5. a) The guidance system with aligned fiber. b) Circling leads to different lengths of guidance system and fiber. L is the length of the guidance system; r is the radius of the circled guidance system; d is the distance of the fiber to the center of the guidance system.",
        "texts": [
            " The big advantage of this method compared to the other two is the increased number of points of measurement: each fiber can use the full wavelength range of the interrogator, while in the other cases the wavelength range has to be portioned to all of the fibers. In general, using an n-fold optical switch multiplies the number of fibers and with it the total number of possible sensing points by n, but also divides the sample rate by n. The maximum and minimum measurable bending radiuses rmax and rmin are calculated as follows (cf. Fig 5): Circling the guidance system of length l, the difference between its length and the length of the attached bended fiber lf is: dl = lf \u2212 l = 2\u03c0(r + d)\u2212 2\u03c0r = 2\u03c0d, (4) where d is the distance between the center of the guiding system and the fiber. The relative change in length \u03b5 of the fiber is \u03b5 = dl l = 2\u03c0d 2\u03c0r = d r . (5) This equals to r = d \u03b5 (6) This means that the maximum and minimum measurable bending radiuses are dependent on two parameters: the distance d of the fiber to the center of the guidance system, which is constant; and the measurable strain \u03b5, which is directly dependent on the wavelength precision and range"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003531_j.piutam.2011.04.015-Figure3-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003531_j.piutam.2011.04.015-Figure3-1.png",
        "caption": "Fig. 3: Free body diagram of the ASLIP. The leg has been drawn to emphasize that it behaves like a massless telescoping force actuator",
        "texts": [
            " Poulakakis and Grizzle introduced an asymmetric monopedal (running) SLIP model (ASLIP, denoted with a subscript \u2018A\u2019) that included a torso, with the linear leg actuators terminating at a hip joint (Fig. 1B). The equations of motion of the ASLIP, Eqns. 8-10, are very similar to those of the SLIP but include torso (of mass m, inertia J, length L, oriented at angle \u03b8A relative to the inertial frame) dynamics, which are a critical component for an anthropomorphic gait model. x\u0308A = 1 m fAx (8) y\u0308A = 1 m ( fAy \u2212 mg) (9) \u03b8\u0308A = 1 J (L( fAx sin(\u03b8A) \u2212 fAy cos(\u03b8A)) + \u03c4A) (10) A statics analysis (Fig. 3) of the massless leg can be used to obtain the expressions for the forces and torques applied to the torso. In the local (x\u2032, y\u2032) axis parallel to the leg, with unit vectors ((x\u0302\u2032, y\u0302\u2032)), we have \u2211 F \u00b7 x\u0302\u2032 = 0, qA \u2212 fA2 = 0 (11)\u2211 F \u00b7 y\u0302\u2032 = 0, pA \u2212 fA1 = 0 (12)\u2211 \u03c4 = 0, \u2212\u03c4A + qAlA = 0. (13) Solving Eqns. 11-13 for the net force applied to the torso ( fA2 x\u0302\u2032 + fA1y\u0302\u2032) as a function of the actuator outputs (pA and \u03c4A), and resolving the result into the global coordinate frame yields fAx = n\u2211 i=1 \u2212pA,i sin(\u03b8A + \u03b1A,i) + \u03c4A,i lA,i cos(\u03b8A + \u03b1A,i) (14) fAy = n\u2211 i=1 pA,i cos(\u03b8A + \u03b1A,i) + \u03c4A,i lA,i sin(\u03b8A + \u03b1A,i)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000418_j.engfracmech.2007.01.016-Figure4-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000418_j.engfracmech.2007.01.016-Figure4-1.png",
        "caption": "Fig. 4. Modelling of the lubricant driven in the crack by the hydraulic pressure mechanism.",
        "texts": [
            " The hydraulic pressure mechanism has been adopted in the presented model, where the lubricant is driven into the crack by contact loading pressure. This results in crack faces being separated by the lubricant pressure, thus implying increased mode I and II separation at the crack tip due to additional normal pressure and lack of friction between crack faces [10,11]. The lubricant pressure in the crack can be simply approximated with a uniform pressure distribution along the crack faces [10], where its level equals the pressure level determined at the crack mouth. Fig. 4 illustrates crack face pressure determination and distribution for two consecutive contact loading configurations. For more realistic simulation of the fatigue crack propagation on gear teeth flanks it is necessary to consider the influence of moving gear teeth contact in the vicinity of initial crack. The moving contact can be simulated with different loading configurations as it is shown in Fig. 5. Five contact loading configurations have been considered, each with the same normal p(x) and tangential q(x) contact loading distributions, but acting at different positions in respect to the initial crack"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002163_cec.2009.4983203-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002163_cec.2009.4983203-Figure2-1.png",
        "caption": "Fig. 2. Current (solid line) and next position (dashed line) of the box after translation.",
        "texts": [],
        "surrounding_texts": [
            "I. INTRODUCTION\nOOPERATION of mobile robots is an interesting area of modern research in multi-agent robotics [13], [14], [15]. Since 1990\u2019s researchers took active interest in formulating and solving the box-pushing problem by different techniques. Some of the well known works in this regard include adaptive action selection by the robots without communication [10], mutual cooperation by intension inference [9], cooperative conveyance by velocity adaptation of robots [11], and role of perceptual cues in multi-robot box-pushing [5]. The principle of subsumption architecture proposed by Brooks [16] was realized in a recent work [17] of cooperative box-pushing problem by mobile robots. The architecture proposed in this work combines the coordination principle of subsumption with motor schemas to obtain an efficient controlled movement of the box.\nThe work proposed in this paper, however, is different from the existing works on multi-agent robotics, as it\nJayasree Chakraborty is with Department of Electronics and TeleCommunication Engineering, Jadavpur University, Kolkata 700032, India (phone: +91 33 25142396; e-mail: jayasree2@gmail.com).\nAmit Konar is with Department of Electronics and Tele-Communication Engineering, Jadavpur University, Kolkata 700032, India (e-mail: konaramit@yahoo.co.in).\nAtulya Nagar is with Intelligence and Distributed Systems Lab, Liverpool Hope University, Liverpool L169JD (email: nagara@hope.ac.uk).\nSwagatam Das is with Department of Electronics and TeleCommunication Engineering, Jadavpur University, Kolkata 700032, India (email: swagatamdas19@yahoo.co.in).\nattempts to satisfy multiple objectives concerning minimization of both time and energy in local trajectory planning of the box by employing an evolutionary algorithm. The NSGA-II algorithm used here provides a Pareto optimal solution, concerning different parameters of the box-pushing problem in each distinct step of local planning.\nIn this paper, we consider a special version of the boxpushing problem, called box-shifting, where two similar robots have to locally plan the trajectory of motion of the box from a predefined starting position to a fixed goal position in a complex terrain with non-liner boundary, containing one or more static obstacles. We presume that the robots do not have any background knowledge about their environment; consequently, the problem of box-shifting is solved here heuristically. Here, the robots jointly attempt to shift (both push and pull) a large box by applying forces at specific locations perpendicular to the edge of the box. The shifting of the box is performed by turning and translating the box in each step of local trajectory planning. The turning involves both push and pull operations, while translation requires only push operation by the robots [12]. In both the operations, the robots stand by one side of the box, and apply force perpendicularly to an edge of the box. A centralized local planning scheme has been adopted to determine the necessary turning angle and displacement of the box, and the magnitude of forces to be applied by the robots on the box. Sufficient spacing between the box and the obstacle is maintained during turning and translation of the box.\nThe most interesting issue of this paper is the formulation of box-shifting as a multi-objective optimization problem. The primary objectives of the box-shifting problem in this context are to minimize the time consumed by the robots for complete traversal of the planned trajectory, and to minimize the exploitation of the robots. In other words, we expect the robots to apply forces efficiently, so that the box is shifted from a given position to the next position (sub- goal) in a time and energy optimal sense without colliding with obstacles or the boundary of the world-map (robot\u2019s environment). To ensure the objective of minimizing time consumption for traversal of the box, we require maximizing the forces applied by the two robots. On the contrary, for minimum energy consumption, the robots have to apply minimum forces. So, there is trade-off between these two objectives. Consequently, the problem of box-shifting here, has been formulated as a multi-objective optimization problem, and has been solved using the well known and most popular multi-objective optimization algorithm, called Non-dominated Sorting Genetic Algorithm-II (NSGA-II) proposed by Deb et al. [2].\nC\n2120978-1-4244-2959-2/09/$25.00 c\u00a9 2009 IEEE",
            "The rest of the paper is organized into 5 sections. In section II, we provide a formulation of the problem. In section III, we provide an overview of NSGA-II and its application in box-pushing. In section IV, we demonstrate the experimental issues and computer simulations for the said problem. Conclusions are listed in section V.\nII. FORMULATION OF THE BOX SHIFTING PROBLEM In this section we use the basic problem formulation undertaken in [12] with slight extension in the nomenclature. We attach a notion of time t to current two dimensional positions of the box and the linear distance and angular rotation selected for motion of the box.\nLet ABCD be the initial position of a box, at time )1( \u2212t , represented by solid line in Fig. 1. Suppose two robots 1R and 2R are applying forces perpendicularly at points E and F on the front edge BC of the box.\nLet O be the centre of gravity of the box at time )1( \u2212t and the coordinates of the points E, F, and O be ( )1(),1( \u2212\u2212 tytx ee ), ( )1(),1( \u2212\u2212 tytx ff ) and\n( )1(),1( \u2212\u2212 tytx cc ) respectively. Suppose the robots 1R and 2R together maneuvered the\nbox around the point ))1(),1(( \u2212\u2212 tytxI ii by an angle\n)(t\u03b1 and centre of gravity after rotation becomes\n( )(),( tytx crcr ), and the corresponding new position of the\nrobots E and F become ( )(),( tytx erer ), ( )(),( tytx frfr ) respectively.\nBy using the principle of static\u2019s we derive the new positions of the robots and centre of gravity of the box, the x- and y- coordinates of which are explicitly given in (1) at time t .\n)1(cos)1())1(cos1)(1()(\n))1()1()(1(sin )1(cos)1())1(cos1)(1()(\n))1()1()(1(sin )1(cos)1())1(cos1)(1()(\n))1()1()(1(sin )1(cos)1())1(cos1)(1()(\n))1()1()(1(sin )1(cos)1())1(cos1)(1()(\n))1()1()(1(sin )1(cos)1())1(cos1)(1()(\n\u2212\u2212+\u2212\u2212\u2212=\n\u2212\u2212\u2212\u2212\u2212\n\u2212\u2212+\u2212\u2212\u2212= \u2212\u2212\u2212\u2212\u2212\n\u2212\u2212+\u2212\u2212\u2212= \u2212\u2212\u2212\u2212\u2212\n\u2212\u2212+\u2212\u2212\u2212= \u2212\u2212\u2212\u2212\u2212\n\u2212\u2212+\u2212\u2212\u2212= \u2212\u2212\u2212\u2212\u2212\n\u2212\u2212+\u2212\u2212\u2212=\nttyttyty\ntytyt\nttxttxtx txtxt\nttyttyty tytyt\nttxttxtx txtxt\nttyttyty tytyt\nttxttxtx\nfifr\nif\nfifr\nie\neier\nie\neier\nic\ncicr\nic\ncicr\n\u03b1\u03b1 \u03b1\n\u03b1\u03b1 \u03b1\n\u03b1\u03b1 \u03b1\n\u03b1\u03b1 \u03b1\n\u03b1\u03b1 \u03b1\n\u03b1\u03b1\n(1)\nConsider another position of the box ABCD with its edge BC at an angle )(t\u03b8 with respect to x-axis. The box is now displaced by a magnitude )(td .The new position of the centre of gravity of the box is now given by,\n)(txc = )(txcr + )(cos)( ttd \u03b8\n)(sin)()()( ttdtyty crc \u03b8+= (2) We, now form an objective function concerning minimization of time, which has three components. The 1st component refers to the time required for rotation, denoted by t1, where\nT\nJtt )(2 1 \u03b1= (3)\nwhere, J=mass moment of inertia T= Torque= 2211 dFdF rr + = 112 dF r , since 2211 dFdF rr = , and, rF1 =force applied by 1R to turn the box,\nrF2 =force applied by 2R to turn the box,\n1d and 2d are the perpendicular distance from the rotational axis to the line of action of the forces.\nThe 2nd time component refers to the time needed for translation of the box to the next position, while the 3rd time refers to the predicted time cost required for transportation of the box from the next position to the goal position. Let 2t and 3t be the respective times defined above. Evaluation of t2 and t3 follows from (4) and (5).\n2009 IEEE Congress on Evolutionary Computation (CEC 2009) 2121",
            "tt FF tmdt 21 2 )(2 + = (4)\nand St \u221d3 or Skt t=3 (5) where, m=mass of the box and Kt is a constant.\ntF1 = force applied by 1R to transport the box, tF2 = force applied by 2R to transport the box,\nfor translation only tF1 = tF2 and S is the distance between the next centre of gravity\nand the goal position of the centre of gravity, 2))((2))(( cgytcycgxtcxS \u2212+\u2212=\n= 2})(sin)()({2})(cos)()({ cgyttdtcrycgxttdtcrx \u2212++\u2212+ \u03b8\u03b8\nHere (3), (4), and (5) are derived from the relations given below: \u03c9\u00d7= JT , where =\u03c9 angular acceleration,\n2 2 1)( tt \u03c9\u03b8 = , 2 2 1 atS = and maF = , where, a=\nlinear acceleration. So, the first objective function is, 3211 tttf ++= (6)\nOur 2nd objective function concerning minimization of energy consumption has also three components, energy consumption for rotation, and energy consumption for translation of the box to the next position and the predicted energy for transportation of the box from the next position to the goal position. If these energy consumptions are denoted by 321 ,, EEE respectively, then the total energy consumption 2f is obtained as 3212 EEEf ++= (7)\nwhere, )(2)( 111 tdFtTE r \u03b1\u03b1 == ,\n)()( 212 tdFFE tt += = )(2 1 tdF t and\nSkE e=3 where, ek is a constant. In our problem, it is also desired that the distance of the nearest obstacle in the direction of movement is as high as possible. For this, we introduce one penalty function. Thus, the 2nd objective function becomes, obsst disfEEEf /3212 +++= (8)\nHere, the objectives are the functions of\nrii Ftytx 1)),1(),1(( \u2212\u2212 , ,1tF )(td and )(t\u03b1 , which we have to determine to optimize the objective functions.\nIII. SOLVING THE BOX SHIFTING PROBLEM USING NSGA-II In this section, we first briefly outline NSGA-II for convenience of the readers, and then present the pseudo code for the entire scheme.\nA. Non-Dominated Sorting Genetic Algorithm (NSGA-II) In a multi-objective optimization problem, we usually need to optimize more than one conflicting objectives [1], [18], [19]. Naturally, finding a true optimal solution satisfying all the objective functions is not feasible. The general trend of solving multi-objective optimization is to determine a Pareto optimal solution set [1] to the problem. Several formulations for determining Pareto-optimal solutions to multi-objective optimization problem, employing evolutionary/swarm optimization algorithms are addressed in the current literature [1], [2], [3], [4], [6], [20], [21]. One of such evolutionary algorithms was proposed by Deb et al. in [2], which is well known as Non-dominated Sorting GA-II (NSGA-II). Due to its good spread of solutions and convergence near the true Pareto-optimal front, low computational requirements, elitism, and parameter less-niching, simple constraint handling strategy, it is widely used.\nLike many other evolutionary algorithms, in NSGA-II also an initial population called parent population P0 (at time t=0) of size N is randomly generated. Then, the population is sorted according to non-domination. Subsequent generations can be represented by discrete time steps: t = 1, 2, ... etc. After initialization, an iterative optimization process begins, where at the first step, using genetic operations i.e. binary tournament selection, recombination, and mutation operations child population tQ of the same size N, is generated from the parent population Pt. Next, the parent and the child populations are combined to form the merged population tR i.e. ttt QPR \u222a= , which is of size 2N. Then, the next population 1+tP is constructed by choosing the best N solutions from the merged population tR . Each solution is evaluated by using its rank as primary criterion and crowding distance as secondary.\nThe ranking is done based on the non-domination. All the non-dominated solutions in the merged population are assigned rank 1. The rank 1 solution set is called front set F1. We now remove these solutions from the merged population, and again look for non-dominated solutions, if any, from the reduced merged population, and then assign rank 2 to these non-dominated solutions. The list of non-dominated solutions thus obtained is called front set F2. In this way, rank is assigned to all the solutions. The members of the population 1+tP are chosen from subsequent non-dominated fronts in order of their ranking. Let lF be the set, beyond which no other set can be accommodated. If by adding set\nlF to 1+tP , size of 1+tP exceeds the population size then to select some solutions (N- 1+tP ) from lF , the set will be sorted based on the crowding distance, and the solutions with higher crowding distance are chosen. For maintaining good spread of solutions in the obtained set of solutions, the crowding distance concept has been introduced instead of choosing random solutions from lF . Crowding distance of a solution is the sum of the difference between the function values of two adjacent solutions for all objectives i.e., to determine crowding distance of a solution, we have to sort\n2122 2009 IEEE Congress on Evolutionary Computation (CEC 2009)"
        ]
    },
    {
        "image_filename": "designv11_12_0002733_0954410012464002-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002733_0954410012464002-Figure1-1.png",
        "caption": "Figure 1. Fuselage angle of attack and sideslip.",
        "texts": [
            " The mathematical model of the maneuvering helicopter has 9 fuselage equations, 16 blade flapping and lead\u2013lagging equations, 16 blade flexibility equations, 3 static main rotor downwash equations, and 1 algebraic flight path angle equation (see Oktay17 for details). These governing equations of motion in implicit form are f \u00f0 _x, x, u\u00de \u00bc 0 \u00f01\u00de at University of Bath - The Library on June 11, 2015pig.sagepub.comDownloaded from where f 2 R 45, x 2 R 41(x is the nonlinear state vector), and 2 R 4 ( is the nonlinear control vector). In this study, level banked turn without sideslip and helical turn without sideslip are examined.35\u201339 For maneuvering flight, the aircraft linear velocities are (Figure 1) u v w T \u00bc VA cos\u00f0 F\u00decos\u00f0 F\u00de VA sin\u00f0 F\u00de VA sin\u00f0 F\u00decos\u00f0 F\u00de T \u00f02\u00de where fuselage angle of attack, F, and sideslip, F, are given by F \u00bc tan 1 w=u\u00f0 \u00de; F \u00bc sin 1 v=VA\u00f0 \u00de \u00f03\u00de Level banked turn is a maneuver in which the helicopter banks towards the center of the turning circle. For helicopters, the fuselage roll angle, A, is in general slightly different than the bank angle, B. For coordinated banked turn A \u00bc B. A picture describing these angles for a particular case ( A \u00bc 0) is given in Figure 2, where Fresultant is the sum of the gravitational force (W) and the centrifugal force (Fcf)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000759_la801006g-Figure4-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000759_la801006g-Figure4-1.png",
        "caption": "Figure 4. This phase diagram presents the preferred orientations for the nanorod as a function of and 2\u03b5 . The color map is used to present the equilibrium orientation, with blue regions denoting 0\u00b0, and dark red regions denoting 90\u00b0. Dark blue regions indicate areas where no particular orientation predominates. A value of \u03c30 ) 50 Oe was used in these simulations as well.",
        "texts": [
            ", 2\u03b5 > 20), the substrate\u2019s field at the center of each particle in the chain becomes weaker, due to the exponentially decaying field distribution produced by the micromagnet array. With further increases in 2\u03b5 , the orientation-dependent potential energy again dominates at lower values of the external field strength. It is worth noting that the jagged characteristic shape of the \u201cpockets of stability\u201d is caused by the discretization of the relative dimensions of the nanorod with respect to the periodicity of the array. Figure 4 presents the most probable orientation of the nanorod as a function of the control parameters, 2\u03b5 and . As clearly seen in Figure 4, nanorod alignment along the external field direction occurs for large and large 2\u03b5 . For fixed in the range of 2-4, a continuous transition between parallel and perpendicular orientation occurs as 2\u03b5 is allowed to increase. This transition is further evidence that the field variation across the nanorod becomes more influential to its total potential energy as the nanorod size increases with respect to the fixed size of the micromagnets. A third distinct region was observed for low values of 2\u03b5 and , for which no particular orientation dominates"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000613_detc2007-35045-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000613_detc2007-35045-Figure2-1.png",
        "caption": "Fig. 2 Experimental test gears, (a) 23T chemically polished gear, (b) 23T ground gear, (c) 40T chemically polished gear, and (d) 40T ground gear.",
        "texts": [
            "org/about-asme/terms-of-use apparatus was disengaged from the flexible coupling, and any electronic drift of the torque-meter was recorded. Electronic drift in the torque signal occurs partly due to thermal expansion of the highspeed spindle during the test, and was subtracted from the measured in order to obtain the actual input torque. This drift torque was typically less than 0.3-0.4% of the measured value. TT TT 2.2 Gear Specimens, Test Matrix and Repeatability A total of four sets of spur gears, shown in Fig. 2, were used in the test matrix shown in Table 1. Each test gear set is formed by two identical gear pairs, or four gears total. The scope of the test program was limited to the investigation of the impact of gear module m, surface finish amplitude, and lubricant type on spur gear efficiency. For the study on gear module, 23-tooth (23T) ground gears with 3.95 mm were compared to 40-tooth (40T) ground gears with 2.32 mm. Table 2 lists basic design parameters of these gears. In order to match the bending strength of the gear teeth, the gear face width for 40T gears was slightly larger than for the 23T gears as shown in Table 2"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003603_0954406212466479-Figure8-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003603_0954406212466479-Figure8-1.png",
        "caption": "Figure 8. Geometric models: (a) rack-cutter and pinion and (b) shaper-cutter (pinion) and internal gear.",
        "texts": [
            " Based on the kinematic relationships shown in Figure 3, Figure 7 demonstrates the generating processes of the pinion and the internal gear by their at University of Bristol Library on January 6, 2015pic.sagepub.comDownloaded from corresponding cutting tool. Here, the rack-cutter is located outside the pinion, while the shape-cutter is located inside the internal gear. Obviously, the generated profile is the envelope of the families of the cutter surfaces. Moreover, the geometric models can be built with the help of the computer-aided software SolidWorks, as shown in Figure 8. These geometric models are helpful for researchers in the design of these gears, such as, finite element analysis, kinematics and dynamics analysis, assembly and interference analysis, etc. Overcutting is a phenomenon where the gear profile is damaged by the cutting tool in the generating process, which usually occurs to gears with a small tooth number, and this decreases the tooth strength. According to some literatures,7,22 overcutting can be determined by the appearance of singular points on the generated surfaces"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003761_9781118818428-FigureA.1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003761_9781118818428-FigureA.1-1.png",
        "caption": "Figure A.1.1 Sequence of four process paths forming the ideal, reciprocating Carnot cycle: 1\u20132 isothermal compression at temperature TC; 2\u20133 adiabatic compression; 3\u20134 isothermal expansion from and at TE; 4\u20131 adiabatic expansion to starting conditions. (The sequence described by Carnot starts with isothermal expansion from and at TE)",
        "texts": [
            " \u00a9 2014 John Wiley & Sons, Ltd. Published 2014 by John Wiley & Sons, Ltd. The ideal Carnot cycle is universally taken for granted. It is re-stated here (in Appendix A-1) firstly to rectify the deficiency of the original Re\u0301flexions, and secondly to introduce the reality that the volume ranges occupied by the respective phases \u2013 isothermal and adiabatic \u2013 are not arbitrary, but require to be pre-set in terms of temperature ratio NT, compression ratio rv, and isentropic index \u03b3. With reference to Figure A-1.1 (Appendix A-1): Initial compression from V1 to V2 is isothermal: T2 = T1 (2.1) The adiabatic phase between volumes V2 and V3 can be defined: T3\u2215T2 = (V2\u2215V3)\u03b3\u22121 (2.2) Combining Equations 2.1 and 2.2: (T3\u2215T1)1\u2215(\u03b3\u22121) = (V2\u2215V1)(V1\u2215V3) The term V1\u2215V3 is volumetric compression ratio rv, and T3\u2215T1 is characteristic temperature ratio NT: V2\u2215V1 = rv \u22121NT 1\u2215(\u03b3\u22121) (2.3) A viable cycle requires V2 < V1, a condition which an arbitrary combination of NT, rv, and \u03b3 does not necessarily satisfy. The selected indicator diagrams illustrate: all p-V loops of Figure 2",
            " The solution strategy can be explored without compromise by working with an engine of 5 cc displacement \u2013 or 0.5 cc \u2013 or 0.05 cc and subsequently converting to full-size by formal scaling. Storage requirement accordingly reduces by further orders of magnitude. Processing power: CPU time in generating Figure 21.2 was 3.59 sec per revolution of the crank-shaft using an 8-year old PC with Windows XP, Pentium P 845 processor and macroscopic counterpart. Appendix 1 The reciprocating Carnot cycle In Figure A.1.1 a cylinder contains a fixed mass of gas enclosed by a piston. Ideal gas behaviour is assumed, but the assumption is not essential. The cylinder and piston are perfect thermal insulators. Three interchangeable cylinder heads are provided, one a perfect insulator, one permanently hot at absolute temperature TE, and one permanently cold at TC. The latter two are of unlimited thermal capacity and of infinite thermal conductivity. Switching between heads can be carried out without loss of gas or thermal leak",
            " Expansion at uniform TE to volumeV4 pre-calculated so that subsequent adiabatic expansion will lowerT to precisely TC when V reaches V1 4\u20131 Insulating cylinder head replaces heat source. Adiabatic expansion to outer dead-centre (V = V1). Stirling Cycle Engines: Inner Workings and Design, First Edition. Allan J Organ. \u00a9 2014 John Wiley & Sons, Ltd. Published 2014 by John Wiley & Sons, Ltd. 248 Appendix 1 Appendix 2 Determination of V2 and V4 \u2013 polytropic processes For the polytropic process 1\u20132 of Figure A.1.1: T2V2 n\u22121 \u2248 T1V1 n\u22121 For adiabatic phase 2\u20133: T3V3 \u03b3\u22121 \u2248 T2V2 \u03b3\u22121 Dividing the latter equation by the former: V2 \u03b3\u22121\u2215V2n\u22121 = ( T3\u2215T1 ) (V3 \u03b3\u22121\u2215V1n\u22121) From the definition of compression ratio V3 = V1\u2215rv, and noting that T3\u2215T1 = NT: V2 \u03b3\u2212n = (T3\u2215T1)(V3\u03b3\u22121\u2215V1n\u22121) = NT (V1\u2215rv)\u03b3\u22121 V1n\u22121 (A.2.1) The counterpart expression for V4 which marks the end of the polytropic expansion phase is: V4 \u03b3\u2212n = 1\u2215NT V1 \u03b3\u22121 (V1\u2215rv)n\u22121 (A.2.2) Stirling Cycle Engines: Inner Workings and Design, First Edition. Allan J Organ"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002187_gt2010-22086-Figure10-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002187_gt2010-22086-Figure10-1.png",
        "caption": "Figure 10-Example of 300 hp foil bearing supported blower from Korea.",
        "texts": [
            " Industry markets machines ranging in size from 50 to 500 hp that compete very well in the commercial air-handling marketplace. These machines are electrically driven, require no maintenance and can be installed at the point where the air is needed, thus eliminating the need for elaborate and expensive air piping systems prevalent in older factories. Such direct drive machines are far more energy efficient than traditional compressors and blowers and eliminate the need for a speed increasing gearbox and the potential for oil contamination of products and factories [19]. Figure 10 shows an example of such foil bearing supported machines engineered and manufactured in Korea and sold into worldwide markets. Like ACM\u2019s, such machines utilize conventional foil bearings (35 to 75mm diameter) and polymer based foil coatings to support rotors that weigh up to about 1000 N. For larger rotors, either larger bearings or a more hybridized approach is required. Downloaded From: http://proceedings.asmedigitalcollection.asme.org/pdfaccess.ashx?url=/data/conferences/gt2010/70392/ on 02/21/2017 Terms of Use: http://www"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002656_c0cp00272k-Figure4-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002656_c0cp00272k-Figure4-1.png",
        "caption": "Fig. 4 Experimental square-wave voltammograms obtained for 2.5 10 4 M solutions of the different ionic liquid salts assayed. Solid lines correspond to the theoretical curves calculated from eqn (7). Esw = 50 mV, Es = 10 mV, t = 0.3 s and the following values of A ffiffiffiffiffiffiffiffi",
        "texts": [
            " 1, it can be observed that for any of the three electrochemical techniques used, the responses corresponding to the transfer of the ions of a salt are much less separate when a membrane system of a single polarizable interface is used, and can even overlap (the current does not fall to zero between them). For example, the separation between the voltammetric signals corresponding to the salt ions that are presented in Fig. 3a\u2013c is approximately 190 mV, in contrast with the nearly 820 mV that is observed in the system with two polarizable interfaces (see Fig. 1a\u2013c and also Table 1), i.e. the voltammetric signals of the salt ions are separated approximately four-times more in this system. This constitutes a huge advantage when ions with opposite charge are studied, as shown in Fig. 4. Fig. 4 shows the experimental and theoretical square-wave voltammograms obtained using a square-wave amplitude of 50 mV corresponding to the transfer from water to the solvent polymeric membrane of the cations and the anions of different salts, together with those of TEA+ and Pic as reference ions. Experimental data were fitted to theoretical equations for the reversible ion-transfer process (eqn (7)) using E 1=2 M;Xz and Fig. 3 Theoretical normal-pulse voltammograms obtained from eqn (1) (Fig. 3a), and theoretical LSV (Fig",
            " Under the experimental conditions used, the last two terms of the right hand side in eqn (12) and (13) are practically zero, such that, in order to determine Dw Mf0 R\u00fe and Dw Mf0 Y we have used Dw Mf0 TEA\u00fe \u00bc 21 mV and Dw Mf0 Pic \u00bc 17 mV, the former obtained from ref. 34 where a 2 : 1 (m/m) NPOE\u2013PVC membrane was used, and the latter obtained from ref. 26 where a 4 : 1 (m/m) NPOE\u2013PVC membrane was used. The average values obtained for the standard ion-transfer potentials of the different salt-ions by using SWV are shown in the first column of Table 1. As can be seen in Fig. 4 and also in the second column of Table 1, the peak separation between the voltammetric responses of the salt-ions, DEp, becomes higher as the lipophilicity of these ions decreases. This behavior can be easily explained by the following equation which has been obtained using eqn (3)\u2013(5) from the present paper and eqn (20) from ref. 30, and taking into account that in our membrane system Cl is transferred at the inner interface when the target ion is a cation, R+, and Li+ is transferred when the target ion is an anion, Y "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001440_978-1-4684-2175-0_8-Figure7-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001440_978-1-4684-2175-0_8-Figure7-1.png",
        "caption": "Fig. 7-Graphical solution of Eq. [21].",
        "texts": [
            " Since the solution is symmetric about x = 0, we will choose the + sign in the following calculations. Integration of Eq. [19] yields or ..l(lf:. + x)sinOM ~M 2 /j(x) f dO' I [ s.inO' J2' o 1 - smOM [20] where -d/2 ~ x ~ o. The solution of Eq. [20] will proceed in two steps. First OM will be determined as a function of ~M (or of H, since ~M = ~K/AX/H). Then this OM(H) can be substituted back into Eq. [20] for the solution of O(x.) as a function of H. From the definition of OM we get d . 0 2~M sm M [21] This equation can be solved graphically as in Fig. 7 by plotting, as a function of sinOM, the left- and right-hand sides on the same graph and locating the points of intersection. By expanding the integral on the right-hand side of Eq. [21], denoted here as L(sinOM), for small values of OM, we get the slope dL(SinOM)\\ 11\" d sinOM 8M _ 0 = \"2. Therefore, for d/(2~M) < 11\"/2 the only solution of Eq. [21] is OM = O. However, when d/(2~M) > 7r/2, a second solution with 8M ~ 0 is ob tained that gives lower free energy than the 8M = 0 solution. The crit ical magnetic field Hp for the transition is found by equating d/(2~M) and 7r/2"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000304_j.mechmachtheory.2007.12.009-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000304_j.mechmachtheory.2007.12.009-Figure1-1.png",
        "caption": "Fig. 1. Section of the shaft.",
        "texts": [
            " However, it is worth underlining that the approach used in the present work allows to highlight the shaft behaviour just during the transitory that leads, depending on the initial conditions, to stability or instability; such behaviour is completely ignored when an asymptotic stability methodology is applied. Consider an elastic shaft with a disk having mass m, statically and dynamically balanced, driven by a motor rotating at constant angular speed. In the deflected configuration the elastic force is directed as in Fig. 1 if x > _#, whereas it is symmetric to OG if x < _#. The equilibrium equations for the rotor are m\u20acx\u00fe k\u00bdx cos\u00f02l\u00de \u00fe y sin\u00f02l\u00de \u00bc 0; \u00f01\u00de m\u20acy k\u00bdx sin\u00f02l\u00de y cos\u00f02l\u00de \u00bc 0; \u00f02\u00de where k is the shaft stiffness and x and y are the coordinates of the centre of mass G. Let z \u00bc x\u00fe \u0131y; then it holds m\u20acz\u00fe kz\u00bdcos\u00f02l\u00de \u0131 sin\u00f02l\u00de \u00bc 0: \u00f03\u00de Since for regular steel l is small \u00f0l 10 3 rad\u00de, then Eq. (3) can be split into m\u20acz\u00fe kz\u00f01 \u01312l\u00de \u00bc 0 if x > _#; \u00f04\u00de m\u20acz\u00fe kz\u00f01\u00fe \u01312l\u00de \u00bc 0 if x < _#; \u00f05\u00de whose solution is given by z \u00bc A0 e\u0131pt \u00fe B0 e \u0131pt; \u00f06\u00de where p \u00bc ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi k m ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 1\u00fe 4l2 pr e 1 2\u0131 tan 1\u00f02l\u00de: \u00f07\u00de Let x0 \u00bc ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi k m ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 1\u00fe 4l2 pr ; \u00f08\u00de then, since 1 2 \u0131 tan 1\u00f02l\u00de \u0131l and e \u0131l 1 \u0131l; \u00f09\u00de Eq"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002973_j.1469-8986.1968.tb02804.x-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002973_j.1469-8986.1968.tb02804.x-Figure1-1.png",
        "caption": "FIG. 1. Construction of the Ag-AgCl electrode, (a) silver disc and lead wire, (b) plastic cup. (c) electrode unit, (d) soft plastic collar.",
        "texts": [],
        "surrounding_texts": [
            "An electrolytic method is presented for the preparation of a disc type Ag-AgCl electrode which has low bias potential and drift, moderate resistance and low polarization, is relatively durable, and can be easily replated if damaged or contaminated. The electrode has been used for measuring both skin potential and conductance and is compatible with a previously published device which compensates for non-isothermal electrodes in skin potential recording. Although other applications have not been explored, the general design might be useful for the recording of other peripherally assessed bioelectric activity as well. A convenient electrode paste is described which does not dry out quickly and can be stored indefinitely.\nDESCRIPTORS: Electrodes, Electrode paste, GSR, Silver chloride. (R. D. Miller)\nSome of the properties desired in a disc type, while less expensive and skin potential (SP) electrode are low easier to construct, has generally been bias potential, high stability, ease of considered inferior to other more comconstruction and maintenance, and plex and expensive designs. Venables convenient physical characteristics and Sayer (1963), how^ever, developed (e.g., size, durability, etc.). Skin con- a simple disc electrode which is funcductance (SC) measurement also re- tionally competitive with the more quires relatively low and stable elec- complicated arrangements, trode resistance and polarization. In A primary disadvantage of Ag-AgCl both cases the electrolytic conducting electrodes, in general, is their vulmedium, or electrode paste, must be nerability to contamination and physcompatible wdth both the electrode ical damage. Any such insult usually surface and the skin. While no system renders the electrode permanently unmeets these criteria completel} ,\u0302 it is usable. Although Ag-AgCl electrodes generally agreed that silver-silver made in this laboratory according to chloride (Ag-AgCl) electrodes provide the instructions of Venables and Sayer the closest approximation. Many ap- were considerably more durable than, proaches to Ag-AgCl electrode con- for example, the sponge type described struetion have been described (Lykken, by O'Gonnell and Tursky (1960), 1965; Ives & Janz, 1961; O'Gonnell, damage to the Venables-Sayer elecTursky, & Orne, 1960). The silver- trode was generally irreparable. While\nthe AgGl layer can be removed with T\u2122T^''?'^\u00b0\"', '^'fr'u'*'\"' ' ' 'u^ dilute ammonia, the bare silver re-\na U b P H b predoctoral fellowship m the . v, J J / I I - J I J . J J . I , - laboratory of David T. Lykken. ^^^^^^ ^.^^ ^^^ electrode replated, this Address requests for reprints to: Ralph Process IS tedious and does not result\nD. Miller, Department of Psychology, ^ as high a quality surface as the McMaster University, Hamilton, Ontario, original. The procedure described beCanada. low produces an equally efficient Ag-\n92",
            "July, 1968 AG-AGCL ELECTRODERMAL ELECTRODES 93\nAgCl disc electrode which is easily replated and consequently has longer hfe.\nDESCRIPTION\nDiscs 3\u0302 ^ inch in diameter and ^^g inch in width are cut from a rod of high purity silver (99.999 %)i and machined smooth. A fine plasticcoated tinned copper wire is soldered to one face of the disc (Fig. l,a), allowing no solder to run to the edge, and the unit washed in acetone. Plastic (acrylic) cups %6 inch in length and % inch in diameter are turned out to an I.D. a few thousandths under yi inch (Fig. l,b). After a ^ 6 inch hole has been drilled through one wall near the base, the cups are washed with detergent and hot water, and dried. The silver disc unit is dipped in low viscosity epoxy cement and pressed\n^ Available from American Smelt and Refining Co., South Plainfield, New Jersey.\ninto the cup so that the lead wire extends from the small hole at the base, and the unsoldered surface of the disc is flush with the open end of the plastic cup. A one inch piece of fine flexible plastic tubing is slipped down the wire and cemented into the 3\u0302 ^ 6 inch hole, thereby providing some support for the lead at the electrode edge (Fig. l,c). A tight, full circle press fit between the silver and the plastic cup is necessary for long term stability since leakage of the electrolyte into the electrode will generate intolerably high and uncorrectable bias potentials.\nAfter hardening 24 hours, the exposed silver face is sanded smooth, with fine water sandpaper and deionized water, and plated by being made the anode in a .07 Molar solution of KCl in deionized water. Batches of six such electrodes are usually plated simultaneously using a single",
            "94 RALPH D. MILLER Vol. 5, No. 1\nsilver cathode. A reasonably constant electrode area. If a single active site is current of 0.5 milliamperes per elec- to be used, it is recommended that the trode is supplied by a 45 volt battery skin under the inactive or reference (e.g., Eveready W-350) in series with electrode be drilled or sanded (Shackel, a 20K ohm potentiometer and a mil- 1959; Venables & Sayer, 1963). liameter.\" After plating for 30 min- After use, the electrodes may be disutes, the electrodes are shorted to- assembled, rinsed off, and stored dry gether for a few hours in .07M KCl or or, if they are to be used frequently, NaCl. The resulting AgCl surface is can be kept with the collars and paste generally coffee-plum in color. Neither undisturbed, shorted in a dilute salt nitric acid nor electrolytic etching solution. Prolonged soaking (e.g., prior to plating produced lower bias several months) may hasten electrode potentials or drift than did just careful leakage and shorten electrode life, sanding. The completed electrode however. Dry electrodes should probshould not be cleaned with acetone ably stand shorted in saline at least or similar solvents which may crack an hour before use in SP measurement, or soften the plastic. Deionized water If the AgCl surface should become is sufficient for most purposes. damaged or contaminated, and the\nBefore use, the finished electrode bias potentials high or unstable, the unit is inserted partway into a collar electrode may be disassembled, the of soft plastic having an I.D. of % silver face sanded smooth and reinch and an O.D. of % inch (Fig. l,d). plated, producing a new surface idenThe inside edge of the collar extends tical in quality to the original. This about }{Q inch out from the base in procedure may be repeated indefithe form of a slight lip. This component nitely, thereby extending the range of may be cut from ordinary plastic or usefulness as well as the normal funcvinyl tubing on a high speed lath, or tion life of the electrode, can be molded from silicon rubber. Since potential differences within When completely assembled, the groups of electrodes plated together flanged end of the collar extends about are smaller than those between sets, it 3-\u0302 inch beyond the AgCl surface of is recommended (for SP measurement) the electrode, thus affording it some that electrodes be prepared in groups physical protection. This unit is filled of at least six and pairs with the lowest with electrode paste and set on a felt potential selected for use. Bias potencorn pad (Dr. ScholVs No. 458) which tial and drift were measured with a has been applied to the skin (e.g.. the Tektronix type 502A oscilloscope in finger) and half filled with paste (Fig. two such sets of six electrodes (in .07 2). The whole assembly is secured with M KCl) for two hours. All possible tape and the lead wire affixed to the combinations of two electrodes within proximal phalanx of the finger to re- each of the sets were recorded, giving duce tension and strain on the elec- a total of 30 different pairs (i.e., trode. In addition to minimizing 2[6 X 5/2]). Potentials between elecmovement and pressure artifacts, the trodes, measured immediately after corn pad provides a convenient means they were unshorted, ranged from 0.00 of limiting the spread of electrode to 0.24 millivolts (mv) with a median paste and thereby defining the effective of 0.08. Readings of drift taken after\n, T,, 4. u +\u0302 \u2022 + t- one hour yielded a range of 0.01 to 2 The exact battery-resistor arrangement -\u0302 ^\nhere is not crucial since current density is 0-20 mv and a median of 0.05. Drift the critical variable. over the complete two hour period"
        ]
    },
    {
        "image_filename": "designv11_12_0002248_690114-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002248_690114-Figure2-1.png",
        "caption": "Fig. 2 - Dynamically loaded bearing with load line and shaft center fixed",
        "texts": [
            " He can discover the effect of changes in speed, power, dimensions, and oil on potential trouble spots, and he can see the difference in behavior of main bearings, crankpin bearings, and wrist pin bushings. Typical orbit diagrams, based on this procedure, are included for comparison with photographs of test bearings for which the journal paths were drawn. 548 techniques intended for day-to-day use are concerned with the unidirectional, constant load situation described in Fig. 1 . The designer must usually be content with the adaptation of one of these techniques for approximate solutions to the dynamic case illustrated in Fig. 2. This gross approach to predicted bearing performance is not consistent with motor industry trends which find sleeve bearing technology being strained to its limits and marginal bearing operation increasing. Recently, several methods have been advanced for a more direct approach to the prob lems found in reciprocating engines. They include hand, graphical, and computer solutions. For the sake of expedi ence, the hand and graphical techniques usually omit con- Table 1 - Data for Dynamometer Tested V-8 Diesel Engine Engine - - 8-cylinder, 318 BHP turbocharged diesel, 4-stroke, V-type Connecting Rod Centerline to Centerline \u2014 8",
            " An equilibrium position is reached with the line of centers making an angle 0 with the load line, the shaft center displaced a small distance e from the bearing center and the load is supported by the oil film pressure. Now imagine that the bearing as well as the shaft in Fig. 1 rotates at a constant speed with respect to the load. This will cause both the attitude angle 0 and the eccentric posi tion e to change but the system is still in static equilibrium. In a reciprocating engine, however, the magnitude of the load, its direction with respect to the bearing and the shaft speed with respect to the bearing all change with time. This presents a situation shown schematically in Fig. 2 in which all speeds are referred to the load direction as being fixed. The fact that the shaft speed with respect to the load w g^ L > the bearing speed with respect to the load w and the load B/L magnitude are continually changing at known rates means that there must be corresponding changes in 0 and e in order that the oil film pressure can support the load. The imme diate problem is therefore to find the change rates d0/dt and de/dt so that the system is in dynamic equilibrium. THE FUNDAMENTAL DIFFERENTIAL EQUATION Applying the continuity equation to incompressible oil flowing into and out of the element dx, dy, dz shown in in which u = absolute viscosity of oil in lb",
            " A-7 gives 9\u00a3 dz 6u dh 12 uc de \u2014\u2014 w - \u2014 + \u2014 cose , 3 EQ dO 3 dt h h Z + C 1 (A-8) If the origin for z is at the middle of the uninterrupted land (Fig. 3) where \u2014 = 0, the constant of integration C = 0. 5 ' DZ 5 1 Thus integrating Eq. A-8 gives the oil film pressure to be P = u dh 2pc de' ~ 3 ^EQ d6 + ~3~ dt h h \u201e 2 3z + C (A-9) or ui sin 6 > EQ 2 cos 0 de dt (A-12) This means that there are two values of 6, 180 deg apart that fix the limits between which the pressure equation (Eq. A - l l ) must be integrated to get the load capacity. Let these limits be 6 and 6 (Fig. 2) and note that from Eq. A - l l the L H film pressure will be zero if T a n 6 = T a n e 2 I H e (j (A-13) EQ Fig. 2 and Fig. 3 show 0 measured from the line of cen ters, and that the line of centers is an angle 0 from the load line. Thus if W represents the total load on the bearing the component parallel to the line of centers is Letting p - 0 at the bearing edges where z = \u00b1 \u2014 gives a value for C which substituted in Eq. A-9 gives W cos 0 = - 2 f'7 J Z = o e \u2022e = e + 7T H L p r cos 0 d 0 dz 3fi w ^ \u2014 + 2c \u2014\u2014 cos 0 EQ de dt dh 2 I I z - h 2 i psi (A-10) Let W = P 2 r 1 to get But h = C (1 + e cos 0) hence \u2014 = - C e sin 0 and Eq. A-10 becomes M \\ 1 P c o s 0 = - \u00b0\u00a3Q 6 I sin 0 cos 0 d 0 0^ (1 + e cos 0) (A-14) P = 3M e sin 0 2 cos 0 de EQ 3 3 dt (1 + e cos 0) (1 + e cos 6) ( A - l l ) psi + 2 de_ dt H 2 o cos 0 e\u0302 (1 + e cos 0) d 0 The load component perpendicular to the line of centers This is the oil film pressure according to the Ocvirk short bearing theory and will be used to get the dynamic load capacity. Refer to Fig. 2 for a typical plot of this. W sin 0 = 2 ~ ~Z /*0 J 7 J Z = o \u2022'e p r sin 0 d 0 dz 562 J . M. ROSS AND R. R. SLAYMAKER Let 7 v r J s i ^ = w E Q e ) H . 2 o sin 6 de 6 L (1 + e cos 0) H sin 6 cos 6 (A-15) d 6 6 L (1 + e cos 6) C = 2 sin 6 d e 6 L ( 1 + 6 cos 6) l 2 \u20ac sin 6.. H sin 6. 2 2 (1 + e cos 6 ) (1 + e cos 6 ) H L EVALUATION OF INTEGRALS The load component equations, Eq. A-14 and Eq. A-15 , contain definite integrals which must be evaluated in terms of 6 and 6 = 6 +TT before they can be used"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003509_j.proeng.2013.08.203-Figure3-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003509_j.proeng.2013.08.203-Figure3-1.png",
        "caption": "Fig. 3. (a) Robotic architecture including the hybrid robot with a parallel kinematic chain, (b) variables of the architecture with the 3 parameters for serial equivalence",
        "texts": [],
        "surrounding_texts": [
            "The problem of path-following with a redundant robot amounts to finding the parameters x which satisfy various constraints related to the task, the robot configuration and the capability expected. The problem can be mathematically expressed as: Let kn xfxf )(: Find *x which minimize 1..k =i )( xf i under 0)(: xhxh and under 0)(: xgxg where n is the problem size and k is the number of criteria 4. Redundancy management Actual researches present many criteria to characterize the behavior of a robotic architecture (Khalil and Dombre (2004)). A set of classical criteria (joint limit, singularity avoidance) is taken into account and specific criteria based on machining constraints are introduced in the optimization: Stiffness to limit the tool deviation Dexterity in a given direction to guarantee high speed due to HSM Mechanical performance to well orient the solicitation towards the manipulator 4.1. Stiffness The objective is to define criteria to model the stiffness behavior of the manipulator. The Cartesian stiffness Kx can be defined by the relation: (1) The results are really different from one architecture to the other because, to sweep the surface with the robot with a parallelogram closed loop, all the actuator moves but only the joints defined by the actuated motions q2 to q5 are charged. Concerning the value obtained, the stiffness is more homogeneous in its workspace. If the wrist is crooked, the force is distributed on link 4 and 5 and the IRB6660\u00ae is stiffer in this configuration (Fig. 5a). Concerning the parallel robot, the extension of the legs leads to a loss of stiffness. Moreover, it can be observed the same pattern on the edges of the swept area (Fig. 5b). Concerning the manipulator with parallel architecture, the deformation induced by the charge is relatively proportional to the median leg length. The following criterion is introduced (Robin et al. (2011)): (2) dX T K x 3 321 qqq rt Regarding serial manipulators, many studies have considered that static deformations are mainly located in the actuated joints (Dumas et al. (2011)). Nevertheless, as far as our structure is concerned, deformations are also located in the links and in the passive joints (Pashkevitch et al. (2011)). As a first approach, to simplify the problem, the elements are considered as non-shrinking elements and the study focuses in a small region of the workspace where the Jacobian is well conditioned. Simplification can be done and the simplified relation between the Cartesian stiffness Kx, the joint stiffness K and the Jacobian J can be written: (3) A stiffness analysis has been realized on the IRB6660\u00ae with the method presented in Dumas et al. (2011). Firstly, the Frobenius norm allows to understand where the Jacobian matrix of the IRB6660\u00ae has a good conditioning (Fig. 6a). Secondly, different weights and configurations allow defining the stiffness of the 2nd to 6th axis by measuring the displacement produced and the forces felt inside the force sensor (Fig. 6b). Once, the stiffness is known for the 5 last joints, tests are then realized to analyze the stiffness of the first axis. A criterion rp is introduced to focus the displacement on the stiffest joints: (4) 1-- JKJK T x )max( 6 1 6 1 i i j rp q k k j i"
        ]
    },
    {
        "image_filename": "designv11_12_0001024_09544062jmes574-Figure3-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001024_09544062jmes574-Figure3-1.png",
        "caption": "Fig. 3 Schematic model of the woodpecker",
        "texts": [
            " The predominant movement of the body is rotational, so the position of the centre of rotation in each frame was decided by taking pairs of adjacent frames of the film and assessing where the point was in the woodpecker\u2019s body that moved the least. A judgement had to be made where there was overall translational motion during the drumming cycle. Schematic model of the woodpecker was predominantly based on the kinematics observed in the woodpecker video footage. The model of the woodpecker is shown in Fig. 3 and the dimensional parameters are shown in Table 1. The following assumptions were made in producing the model: (a) the tree was modelled as a stiff spring and damper; (b) the head and body both rotate about the centre of rotation; (c) the woodpecker\u2019s vertebrae and neck tendons were modelled as a spring; (d) the force input from the woodpecker\u2019s legs is assumed to be sinusoidal and horizontal; (e) the centre of rotation is assumed to be fixed at a distance from the tree equal to the distance from the centre of mass of the head to the tip of the beak"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003713_978-1-4614-2419-2-Figure50.8-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003713_978-1-4614-2419-2-Figure50.8-1.png",
        "caption": "Fig. 50.8 Bending moment about the z axis",
        "texts": [
            "2) being b the distance between sections AA and BB (see Fig. 50.6). It must be in fact considered that: \u2022 The lateral force Fy provides a constant contribution to the bending moments about both the x and the z axis; \u2022 The vertical force Fz originates a bending moment about the x axes, which increases proportionally to the distance from the skate (triangular distribution, Fig. 50.7); \u2022 The longitudinal force Fx originates a bending moment about the z axes, which increases proportionally to the distance from the skate (triangular distribution, Fig. 50.8). Fx, Fy, Fz. Moreover the proposed measuring device also theoretically permits the estimation of the position of the application point P of the ice-skate contact force, once known the contact force itself. The z (bz) and x (bx) position of point P can be in fact determined according to: bx \u00bc MzBB Fxa Fy ; bz \u00bc Fza MxAA Fy (50.3) 490 F. Braghin et al. between the runners carrier and the skate. Both the front and the rear axles have been instrumented in order to measure the contact forces in correspondence of each of the four skates"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002481_tpas.1969.292344-Figure9-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002481_tpas.1969.292344-Figure9-1.png",
        "caption": "Fig. 9 Short-circuit torque.",
        "texts": [
            " The numerical computation of this example is actually done on a digital computer at the Computer Center of the University of Hawaii. The values of if are calculated from (69) and are plotted in Fig. 3. illd and ijug are plotted in Figs. 4 and 5, respectively. Equations (65), (70), and (72) are used for computing ili, and (71) and (73) for ill. 4b and i6, by using (67) and (68), are plotted in Figs. 6 and 7, respectively. The values of ea are calculated from (74) and (76) and are plotted in Fig. 8. By using (72), (73), and (79), T is plotted in Fig. 9. In addition to the listed parameters, Go = 7r/3, E, = 1.00 pu, and Ifo = v\"3/2 pu are assumed in all calculations. 1590 HWANG: UNBALANCED OPERATIONS OF THREE-PHASE MACHINES can also be applied gainfully to particular cases of machines with two or more additional damper circuits in each d and q axis. This method is especially useful in the analysis of a round-rotor machine that suggests the possible need for at least two damper circuits in the direct axis, one simulating induced currents near the surface of the rotor, and the second simulating eddy currents of deeper penetration into the rotor body"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003342_cjme.2012.01.081-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003342_cjme.2012.01.081-Figure2-1.png",
        "caption": "Fig. 2. Slip distribution of inner/outer contact patch",
        "texts": [
            " The tangential elastic compliance during rolling process makes the actual pattern of relative velocities within the contact patch more complex[14]. The tangential elastic compliance makes the profiles of sliding distribution in the contact ellipses fundamentally different from those depicted by HARRIS[8] and LEBLANC, et al[15]. Due to the tangential elastic compliance, the traction and slip distributions within the contact patch are no longer in a circumferential direction. The traction vectors in Fig. 1 and the sliding lines in Fig. 2 within the contact patches are apparently not circular, especially for those near to the center of locked areas where no slip occurs. Owing to the spin, the tangential elastic compliance and the close conformity, no pure rolling band actually exists within both inner and outer contact ellipses. A more careful examination reveals that the series of traction stream lines in Fig. 1 and quasi-circular sliding lines in Fig. 2 are not concentric. The smaller the radius of quasi-circle is, the closer the center to the trailing edge of the contact ellipse. To be more detailed, the stream lines are not even closed curves but internal or external spirals. The eccentricity and spirality of the stream lines are closely concerned with the tangential elastic compliance of the contact surface. In the case of Coulomb's friction, a larger coefficient of friction and a lower shear modulus of elasticity generally lead to CHINESE JOURNAL OF MECHANICAL ENGINEERING \u00b785\u00b7 more remarkable eccentricity and spirality of the vector distributions",
            " As can be seen, the gyro-like motion of the ball is seemingly unavoidable even if the gyroscopic moment induced by the orbital circulation of the ball is excluded. Thus it can be predicted that the arbitrary elimination of the yaw angle \u03b2 (i.e., \u03b2=0) is most likely to introduce an extra lateral force that cannot be counteracted. For rolling contact, the slip starts at the trailing edge of the contact patch and then spreads over the contact patch as the creepage and spin are increased. At small spin, interfacial slip is eliminated from an appreciable fraction of the contact area. It is apparently shown in Fig. 2 that two separate parts of the inner contact is void of sliding lines, which means no slip occurs there, due to the small spin i\u03d5 of the ball with respect to the inner raceway. However, even at large spin, an area of no slip may still exist somewhere within the contact patch[14]. As shown in Fig. 2, there is no slip in a small area around the center of the outer contact ellipse, probably arising from the large coefficient of friction between the ball and outer raceway. Owing to the existence of no slip area, the frictional spinning moment zM and the resultant lateral and longitudinal tractions i/oxT and i/oyT actually exerted during the rolling process are appreciably lower than those calculated by the method of JONES[7]. It is due to the fact that the traction within the area of no slip is lower than the traction bound defined by Coulomb\u2019s friction",
            " The valleys or notches of different depth on the semielliptical profiles represent the presence of locked or stick areas, i.e., areas where the resultant tangential traction is lower than the traction bound and thus no slip occurs. As shown in Fig. 3, all slices except slice 9 contain one or two valleys, which means there\u2019s a considerable portion of the inner contact ellipse without slip. But it is not true for the outer contact areas. As shown in Fig. 4 that the slip has spread almost all over the contact ellipse due to the large spin of the ball relative to the outer raceway. Referring to Fig 1 and Fig. 2, careful readers may find that the spinning center of inner contact patch is much closer to the trailing edge of the contact ellipse than that of the outer one. Apart from the neglectable difference in the sizes of inner and outer contact ellipses, it\u2019s mostly the results of the much different traction distributions of inner and outer contact patches. As shown in Fig. 3 and Fig. 4, the resultant traction within the locked area is much lower than the traction bound. To accumulate an inner lateral Part II: Results and Discussion \u00b786\u00b7 CHEN Wenhua, et al: Quasi-static Analysis of Thrust-loaded Angular Contact Ball Bearings traction iyT comparable to the outer one oyT , the spinning center of the inner contact patch needs to move a little further towards the trailing edge to compensate for the traction reduction in the locked areas",
            " The spinning center likewise moves a little further along the negative y-axis, though it is hardly notable in Fig. 3 and Fig. 4. Referring to Fig. 3, readers may also find that traction slices 1\u20134 are apparently unsymmetrical with respect to the center of the contact ellipse. The asymmetry of the traction profiles may be attributed to the unified effect of differential slip and spin of the ball relative to the inner raceway. When the ball rolls forward with a spin on the inner raceway, the counterclockwise spin as illustrated in Fig. 2 seems to alleviate the sliding trend caused by the differential slip in the upper part of the contact ellipse to some extent but to strengthen that in the lower part of it. This makes it easier for the traction in the lower part of the contact ellipse to reach the semi-ellipsoidal traction bound and then the asymmetry occurs. The results of the numerical simulation also reveal that the valleys will be a little shallower and narrower if differential slip is taken into consideration. This can be accounted for by the differential slip between the ball and raceway, due to the significant curvature of their common contact ellipse"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001001_asjc.11-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001001_asjc.11-Figure1-1.png",
        "caption": "Fig. 1. The TRMS helicopter system.",
        "texts": [
            "ey Words: TRMS, set-point stabilization, nonlinear PD control, fuzzy PID control. I. INTRODUCTION A twin rotor MIMO system (TRMS) as shown in Fig. 1 is an aerodynamic system similar to a helicopter: a beam pivoted on a base with propellers at both ends driven by DC motors. An articulated joint allows the beam to rotate in such a way that its ends move on a spherical surface. The TRMS helicopter system has main and tail rotors for generating vertical and horizontal propeller thrust, and requires only two easily-measured outputs: the pitch angle and yaw angle (or azimuth angle) of a helicopter. Feedback stabilization of the TRMS helicopter system is a favorite MIMO control problem in control system analysis and controller design"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001939_s12283-009-0033-4-Figure6-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001939_s12283-009-0033-4-Figure6-1.png",
        "caption": "Fig. 6 Moments at the ankle joint complex at push-off",
        "texts": [
            " From this, the resultant total force vector is approximately 8 \u201314 back to vertical from the V reaction force, or approximately 16 \u201322 off-center from vertical. From known skate orientation and ankle joint alignment during push-off in combination with the current study\u2019s force data, we may speculate on the mechanical interactions of the boot and foot\u2013ankle complex during push-off [5, 6]. For instance, the resultant force off-center to the calcaneous may accentuate the eversion moment at the subtalar joint. In turn, an equal and opposite moment on the talus\u2013tibia\u2013fibula complex may pivot these bones laterally (Fig. 6), thereby, pressing the medial and lateral malleoli (distal tibia and fibula) into the medial and lateral walls of the upper boot collar, respectively (stars in Fig. 6). Pressure patterns measures demonstrated by Dewan et al. [5, 6] within the skate boot while skating are congruent with this scenario. The results of this study demonstrate that the current system can be used on-ice to capture relevant kinetic data during the performance of ice hockey skating skills. Strain gauge signals produced a high linear relationship to known force values, independent of loading rate with an acceptable error. These force estimates were consistent trial-totrial. Along with rapid data collection and meaningful force\u2013time records, the use of this system permitted natural, unencumbered skating during an on-ice situation"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003520_j.euromechsol.2011.11.003-Figure6-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003520_j.euromechsol.2011.11.003-Figure6-1.png",
        "caption": "Fig. 6. T2R1-type parallel manipulator with uncoupled and bifurcated planar-spatial motion of the moving platform: constraint singularity (a), branch with planar motion (b) and branch with spatial motion (c); limb topology PkRtPtkRtRPtPttRtRSPtRkP.",
        "texts": [],
        "surrounding_texts": [
            "The term of constraint singularity (CS) has been recently coined (Zlatanov et al., 2002) to characterize the configuration of lower mobility parallel manipulators in which both the connectivity of the moving platform and the mobility of the parallel mechanism increase their instantaneous values. From a constraint singularity, the mechanism can get out with or without branching (Gogu, 2008b). When branching occurs in a constraint singularity, the mechanism can reach different configurations, called branches, and have different independent motions of the moving platform. In this case, the constraint singularity is also called branching or bifurcation singularity. A branch refers to the free-of-singularity configurations of the mechanism inwhich each structural parameter keeps its value. For this reason, this value is called global of full-cycle value for a branch. Two types of branching in constraint singularity (BCS) have been defined in Gogu (2008b). Branching of type BCS1 occurs when a parallel mechanism F ) G1 . Gj . Gk gets out from a constraint singularity in different branches characterized by the same mobility and the same connectivity of the moving platform but with different bases of the vector space of relative velocities between the moving and fixed platforms. Branching of type BCS2 occurs when a parallel mechanism F ) G1 . Gj . Gk gets out from a constraint singularity in different branches characterized by distinct values of mobility and connectivity of the moving platform. The parallel manipulators presented in this paper have constraint singularities with branching of type BCS1. Branching of type BCS2 occurs in kinematotropic mechanisms (Gogu, 2009c). The finite displacements and the velocities in the actuated joints are denoted by qi and _qi, the linear velocities of the characteristic point H of the moving platform, by v1 \u00bc _x; v2 \u00bc _y and the angular velocity of the moving platform by u \u00bc ua \u00bc _a or u \u00bc ud \u00bc _d: In both branches, the moving platform undergoes two planar translations and one rotation but the rotation axis is different in the two branches. In the first case, the rotation velocity u \u00bc ua \u00bc _a is parallel to the x-axis and the moving platform undergoes a spatial motion. In the second case, the rotation velocity u \u00bc ud \u00bc _d is parallel to the z-axis and the moving platform undergoes a planar motion. With these notations, the linear mapping between the actuated joint velocity space and the end-effector velocity space for a branch the T2R1-type PMs is defined by: 2 4 v1 v2 u 3 5 \u00bc \u00bdJ 2 4 _q1 _q2 _q3 3 5 (8) where J is the Jacobian matrix. In the constraint singularity, the T2R1-type parallel manipulators with bifurcated motions has instantaneously iM \u00bc iSF \u00bc 4 and (iRF)\u00bc(v1, v2, ua, ud). The bifurcation in this constraint singularity can be used to change motion type of the moving platform. In the two distinct branches, the parallel mechanism is characterized by M\u00bc SF \u00bc 3 and (RF)\u00bc(v1, v2,ua) or (RF)\u00bc(v1, v2,ud). In the constraint singularity, both connectivity of moving platform and mobility of parallel manipulator increase their instantaneous values. More details about the behaviour of parallel manipulators in a constraint singularity and the associated structural parameters can be found in Gogu (2009c)."
        ]
    },
    {
        "image_filename": "designv11_12_0003835_ceit.2015.7233009-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003835_ceit.2015.7233009-Figure2-1.png",
        "caption": "Fig. 2. The Coaxial-rotor structure with the associated forces and frames.",
        "texts": [
            " In Section 2 the mathematical model of a coaxial rotorcraft is presented, while in Section 3 the control algorithm is designed and applied to a coaxial rotorcraft. Section 4 contains the main results of the paper and presents the analysis of the closed-loop performance of the proposed control. In section 5, some simulations are carried out to show the behavior and stability of the closedloop system. Finally, in section 6, conclusions are presented. The mathematical model for the flying machines such as the coaxial-rotor MAV is essential for designing a control algorithms with satisfactory performance. Consider the coaxial-rotor depicted in Fig. 2 as a solid body incorporating with a force and moment generation process. Let B := (G, xb, yb, zb) the body-fixed frame attached to the center of gravity of the aerial vehicle, where xb is the longitudinal axis, yb is the lateral axis and zb is the vertical direction in hover conditions and I := (O, xI , yI , zI) be the Earth frame. The generalized coordinates describing the rotorcraft position and orientation are q = [\u03be, \u03b7]T , where \u03be = (x, y, z)T \u2208 R 3 represents the translation coordinates relative to the inertial frame and \u03b7 = (\u03c6, \u03b8, \u03c8)T \u2208 R 3 are the Euler angles representing the orientation of the rotorcraft in frame I, where \u03c6 is the roll angle around the x-axis, \u03b8 is the pitch angle around the y-axis, and \u03c8 is the yaw angle around the z-axis"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003501_2041302510394742-Figure12-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003501_2041302510394742-Figure12-1.png",
        "caption": "Fig. 12 Gearing arrangement of pitch control: (a) gearing arrangement of pitch control for the free wings (hovering mode), (b) gearing arrangement of pitch control for the free wings (mode transition), and (c) gearing arrangement of pitch control for the free wings (forward flight)",
        "texts": [
            "3 Angular velocities of tail counter-rotating propellers For the tail counter-rotating propellers, it is required that Mt1 = Mt2 (94) Therefore, one can obtain kc1\u03c9 2 t1 = kc2\u03c9 2 t2 (95) For the given Tt0, from Tt0 = kt1\u03c9 2 t1 + kt2\u03c9 2 t2 in equation (5) \u03c9t1 = \u221a kt1kc2 kt1kc2 + kc1kt2 Tt0, \u03c9t2 = \u221a kc1kt2 kt1kc2 + kc1kt2 Tt0 (96) 5.4 Gearing arrangement motion In order to realize the dynamics of \u03b81 according to the dynamics of \u03b8 and keep free wings a given angle of attack, a gearing arrangement is design shown in Fig. 12. Threaded screw rods l3 are within fixed wing Proc. IMechE Vol. 225 Part G: J. Aerospace Engineering 841 at The University of Manchester Library on April 23, 2015pig.sagepub.comDownloaded from 3 and can rotate driven by a motor. The thread collars H2 gear with threaded screw rods l3 and can glide along threaded screw rods l3. Thread collars H2 are connected with free wing through the link l2. Therefore, when threaded screw rods l3 rotate driven by a motor, thread collars H2 move along Proc. IMechE Vol",
            " The free wing should maintain a given angle of attack during mode transition and forward flight. The dynamics of \u03b8 are regulated by tail counter-rotating propellers with the pairings. We let |CH1| = l1, |H1H2| = l2, |CH2| = l3 (97) where l1 and l2 are constant, and l3 is variant. For a given angle of attack \u03b11, \u03b81 = \u03b11 + \u03b3 , with \u03b3 = tan\u22121(z2/y2). Proc. IMechE Vol. 225 Part G: J. Aerospace Engineering at The University of Manchester Library on April 23, 2015pig.sagepub.comDownloaded from From Fig. 12(a), at the hovering mode, l2 > l1. From the triangle relation in Fig. 12 l2 2 = l2 1 + l2 3 \u2212 2l1l3 cos(\u03b8 \u2212 \u03b81 + \u03b8g) (98) One can obtain l3 = l1 cos(\u03b8 \u2212 \u03b81 + \u03b8g) + \u221a l2 1 cos2(\u03b8 \u2212 \u03b81 + \u03b8g) + (l2 2 \u2212 l2 1) (99) Proc. IMechE Vol. 225 Part G: J. Aerospace Engineering at The University of Manchester Library on April 23, 2015pig.sagepub.comDownloaded from Therefore, from Fig. 12(b) sin \u03b82 = l1 cos(\u03b8 \u2212 \u03b81 + \u03b8g) l2 (100) and cos \u03b82 = \u221a 1 \u2212 ( l1 cos(\u03b8 \u2212 \u03b81 + \u03b8g) l2 )2 (101) The length of l3 should be controlled to obtain the desired angle \u03b81d. When the aircraft hovers (see Fig. 12(a)), shaft yb is upward vertically. Shaft l3 rotates driven by a motor. From Fig. 12(b) and equation (23), one can obtain the force FM generated on the free wing FM l3 l1 sin(\u03b8 \u2212 \u03b81 + \u03b8g ) l2 = M (102) Proc. IMechE Vol. 225 Part G: J. Aerospace Engineering at The University of Manchester Library on April 23, 2015pig.sagepub.comDownloaded from Xinhua Wang and Hai Lin that is FM = Ml2 l1l3 sin(\u03b8 \u2212 \u03b81 + \u03b8g) (103) It is required that \u03b8 \u2208 [\u03b81, 90\u25e6], that is to say, l1, l2, and l3 are not in the same line. Therefore, shaft l3 should be rotated and a force generated for thread collars H2 (see Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001724_tmag.2010.2072910-Figure5-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001724_tmag.2010.2072910-Figure5-1.png",
        "caption": "Fig. 5. Magnetic power loss. (a) Slip . (b) Slip . (c) Slip .",
        "texts": [
            "5 times larger than that in the transverse direction. Fig. 4(a)\u2013(c) shows the distributions of the maximum magnetic field intensity of each condition (slip 0, 0.2, and 0.5). As shown in Fig. 4(a), as well as Fig. 3(a), the magnetic field intensity was the largest because the eddy current was very small. With increasing the load and the slip, the magnetic field intensity became smaller as well. In addition, in the transverse direction (TD: vertical direction of the figures), the magnetic field intensity increased because of magnetic anisotropy. Fig. 5 shows the iron-loss distributions calculated from the vector relation between and . Usually, the iron loss is assumed to be in proportion with the square of the maximum magnetic flux density, . Actually, the iron loss cannot be determined with only . In this analysis, it was clarified the iron loss increased at parts where both and were large. Generally, the waveforms of the magnetic flux density and the magnetic field intensity in the stator yoke of the induction motors include secondly slot harmonic components"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001568_ie801291h-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001568_ie801291h-Figure2-1.png",
        "caption": "Figure 2. Schematic diagram of electrochemical cell and electrode arrangement.",
        "texts": [
            " For each experiment, two liters of solution at a desired dye concentration and supporting electrolyte was taken and constant current was applied through the D.C. power supply unit. Three different supporting electrolytes, viz., NaCl, NaNO3, and Na2SO4, were used at a uniform concentration of 0.01 M. The effect of NaCl concentration was further studied at different concentrations of 0.01, 0.05, and 0.1 M. While the current density was increased from 7.88 to 23.64 mA cm-2, the voltage was increased from 3.2 to 8.6 V. The schematic diagram of the electrochemical cell is shown in Figure 2. All the experiments were performed at ambient room temperature. The samples were collected periodically at different time intervals and analyzed for COD and TOC. Samples were centrifuged while using graphite as an anode to ensure that the samples are free from mechanically disintegrated fine carbon particles. The standard methods suggested by the American Public Health Association (APHA)16 were adopted for the analysis of COD, Cl-, SO4 2-, and total Kjeldhal nitrogen (TKN). The decolorization of the dye solution was monitored using a UV-vis spectrophotometer (Shimadzu, UV-160A)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003713_978-1-4614-2419-2-Figure17.5-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003713_978-1-4614-2419-2-Figure17.5-1.png",
        "caption": "Fig. 17.5 Dissimiliraties can be seen between mode 10 and 15 comparing relative amplitude of Fourier descriptors",
        "texts": [
            " Good approximation may reasonably be achieved by retaining only a small number of high energy terms (50 terms). We can see clearly the approximation is enhanced with number of Fourier Descriptors (Fig. 17.3). The MAC is used to analyse the correlation between the undamaged plate mode shapes with each other, the following results are obtained (Fig. 17.4). From the MAC results on a simple plate we can observe high similarities between mode 10 and 15 for example (so a coefficient close to 0.7 in the MAC matrix in Fig. 17.4). But just comparing FD of these two modes (Fig. 17.5), we can see that important dissimilarities exist. We can also conclude that reconstruction from FD can help to have \u201clocal\u201d information of dissimilarities. developed to distinguish close modes using the advantage of comparing at different levels of approximation (From Fourier to Wavelets Descriptors). Changes in modal parameters (frequency, damping, mode shapes) are commonly used in SHM to detect, localize and identify damages in structures. FRF updating process can help to localize damages [25] but modeshapes are often difficult to use as a tool for localization of damages"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003184_j.neucom.2012.01.003-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003184_j.neucom.2012.01.003-Figure1-1.png",
        "caption": "Fig. 1. Spherical inverted pendulum.",
        "texts": [
            " Since the exact solution of the complexity of the equations, the paper [21] tried a polynomial approximation of the solution of the regulator equations. In this paper, we first show that the solution of the regulator equations associated with the spherical inverted pendulum exist and then find an approximate solution to the output regulation problem of the spherical inverted pendulum via a neural network approximation approach. We also make some comparison between the method in this paper and the method in [21]. Crown Copyright & 2012 Published by Elsevier B.V. All rights reserved. Shown in Fig. 1 is the spherical inverted pendulum [17,21] where x,yAR represent the position of the base of the pendulum in the horizontal plane and X,YAR represent the x and y positions of the vertical projection of the center of the pendulum onto the horizontal plane, Fx,FyAR are the control forces being applied to the cart at the base of the pendulum, m is the mass of the uniform rod, L is the distance from the base of the pendulum to the center of mass, and g is the gravitational constant. The motion equations of the spherical inverted pendulum are as follows [17,21]: _x \u00bc f \u00f0x\u00de\u00feg\u00f0x\u00deu y\u00bc h\u00f0x\u00de \u00f01\u00de where x1 \u00bc x, x2 \u00bc _x, x3 \u00bc y, x4 \u00bc _y, z1 \u00bc X, z2 \u00bc _X , z3 \u00bc Y , z4 \u00bc _Y u1 \u00bc Fx, u2 \u00bc Fy, u\u00bc \u00bdu1 u2 T x \u00bc \u00bdx1 x2 x3 x4 z1 z2 z3 z4 T 012 Published by Elsevier B"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000161_iccas.2007.4406826-Figure9-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000161_iccas.2007.4406826-Figure9-1.png",
        "caption": "Fig. 9 Change of the joint configuration during one-step walking.",
        "texts": [
            " We assume that the joint control has been performed in Fig. 7 and accordingly joint angle and ZMP measurements are produced. Then, the remaining work is to recursively compute the CoM observer in Fig. 5. On the other hand, certain sinusoidal noises are added to real joint angles to consider the uncertainty effect owing to the flexible motion of links, sensory limit, and other disturbance inputs from the ground. As a result, real joint motions become quite different from their measurements. Finally, numerical results are shown in Figs. 9 to 11. Figure 9 denotes the walking pattern during one step motion due to the change of joint angles. The estimated CoM trajectories are compared with the reference in Fig. 10. While, the prior and posterior estimation error and the measurement error are compared one another in Fig. 11. In this paper, a Kalman filter based CoM observer was proposed for humanoid robots. To apply the discrete Kalman filter in the CoM estimation problem, the inverted pendulum equation was adopted as the CoM motion model and the CoM conversion equation as the measurement model"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000741_3.6738-FigureI-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000741_3.6738-FigureI-1.png",
        "caption": "Fig. I Structures amenable to employing initial strain procedure.",
        "texts": [],
        "surrounding_texts": [
            "1 Martin, H. C, \"Large Deflection and Stability Analysis by the Direct Stiffness Method,\" TR 32-931, Aug. 1966, Jet Propulsion Lab., Pasadena, Calif. 2 Martin, H. C., \"On the Derivation of Stiffness Matrices for the Analysis of Large Deflection and Stability Problems,\" Proceedings of the Conference on Matrix Methods in Structural Mechanics, TR-66-80, October 26-28, 1965, Air Force Flight Dynamics Lab., WrightPatterson Air Force Base, Ohio, pp. 697-715. 3 Mei, Chuh, \"Nonlinear Vibrations of Beams by Matrix Displacement Method,\" AlAA Journal, Vol. 10, No. 3, March 1972, pp. 355-357. 4 Timoshenko, S., Strength of Materials, Pt. I, 3rd ed., D. Van Nostrand, New York, 1955, pp. 267-268."
        ]
    },
    {
        "image_filename": "designv11_12_0002531_s10800-010-0153-3-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002531_s10800-010-0153-3-Figure1-1.png",
        "caption": "Fig. 1 The electrochemical cell",
        "texts": [
            "2 Electrode preparation, electrochemical cell, and procedures A mixture, constituted of 97% of graphite carbon and of 3% of Teflon used as binder, was homogenized and compacted in a mold to obtain pellets which were prepared by pressing the mixture (100 or 200 mg) under a high pressures (0.5\u20131.0 ton cm-2), using a KBr tablet mould in a Carver Laboratory press model S/N 3392-69. These pellets presented the advantage to possess a good electronic conductivity and a relatively high porosity (corresponding to the amount of electrolyte retained in the pellet). Before each electrochemical measurement, the pellet mass was weighed, and then was introduced into a specially built electrochemical cell made in Teflon (Fig. 1). This cell included a counter-electrode (stainless steel grid), a reference electrode (saturated calomel electrode, SCE), and a gold plate which was connected to a wire, used as a collector allowing the electric current to flow. The pellet had to be well fixed on the gold surface, in order to avoid any direct polymer deposit on the gold plate. An acetonitrile solution containing pyrrole (0.2 M) and NaClO4 (0.1 M) was then put into the electrochemical cell which was connected to a potentiostat. The pellets were immersed into this electrolytic solution during 8 h, in order to ensure a good impregnation before starting the PPy electrodeposition"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001884_s11431-010-3100-y-Figure8-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001884_s11431-010-3100-y-Figure8-1.png",
        "caption": "Figure 8 A combination mechanism formed by two Sarrus mechanisms.",
        "texts": [
            " Therefore, the angle of links 6 and 8 is equal to that of links 7 and 6, which also means that  is not an independent parameter. dX III=dX III( 0 0, 0 y z)=dX III(0 0 0, 0 y z)=2. FIII= PIII i=1fi dX III=22=0. The mobility of this mechanism is F=FI +FII +FIII =4+(1)+0=3. In this loop, since dIII =3 and dX III =2, they are not equal. This loop has a virtual constraint. It can be seen that the PR(Pa)R chain, which is generally considered as a mixed chain with parallel and serial chains, is actually a composite chain with an independent loop. It can be decomposed. Example 5. Figure 8 shows a combination mechanism formed by two Sarrus mechanisms. Loop I contains links 1-2-3-4-5-6, loop II includes links 1-2-3-7-8-6, and loop III consists of links 4-5-3-7-8-6. Any two loops can be chosen as independent loops. If the K-G criterion is used to calculate the mobility, since all links cannot rotate about z-axis, this mechanism has one common constraint, i.e., m=1, d=5, F= d(ng1)+ g i=1fi =5(891)+9= 1. If the mobility is calculated by the modified K-G criterion F= d(ng1)+g i=1fi +v, the redundant constraint v (v=2) is needed to calculate according to screw theory, and then the correct mobility F=1 will be obtained"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003497_amr.498.127-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003497_amr.498.127-Figure1-1.png",
        "caption": "Fig. 1. Material usage in a GP7000 engine of the Airbus A380 (Source: MTU Aero Engines)",
        "texts": [
            " In combination with the necessity to save resources, this leads to a great demand for energy saving short- and midrange air planes in the next years. The highest potential to reduce fuel consumption of an air plane lies within its engines. Future engine components will be of a more complex geometry, like Blisks, and made of advanced materials. Under these conditions in the future the production technology has to deal with the demand of increased productivity during machining of highly heat resisting nickelbased alloys and titanium alloys with an elevated mechanical strength. Materials used in aircraft engines Fig. 1 shows the Material usage of a GP7000 engine which is used for the Airbus A380. For the colder parts of the turbine, as the fan, the low and high pressure compressor and the low pressure turbine, titanium alloys and high alloyed steels can be used. In the hot sections, like the combustion chamber and the high pressure turbine, highly heat resisting materials are necessary. The main proportion of materials to be used are nickel-based and titanium alloys. Most common are the nickelbased alloy Inconel 718 and the titanium alloy TiAl6V4 which are in focus of this paper"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001336_amc.2008.4516091-Figure8-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001336_amc.2008.4516091-Figure8-1.png",
        "caption": "Fig. 8. Mobile manipulator",
        "texts": [],
        "surrounding_texts": [
            "Experiment was carried out to check the validity of proposed method using two wheeled mobile manipulator shown in Fig.(1). At the start of our experiment, constant velocity tests were carried out and tested motion result was presented in earlier section of this paper. The purpose of this work is to deal the robot with in the rough terrain. Therefore, running environment of the robot was prepared as shown in Fig.(8). Then, robot was run over the rough terrain. Robot trajectory was set to straignt line so that it allows to go over the obstacle that was kept on the floor. First experiment was conducted without proposed method and then it was repeated with proposed method. System parameter of robot is shown TABLE I PARAMETER USED FOR ENVIRONMENTAL MODEL greac=30.0 Kenv=20.0 Denv=32.0 Menv=8.5 Aenv=0.001 in table (II) and (I) show the value of environmental model. Input torque of the wheel motor is shown in Fig.(9). Fig.(9) (a) shows the input torque with proposed method and Fig.(9) (b) shows the wheel motor torque variation in both cases. In Fig.(9) (b), green line is the case where without proposed method. It is clear that input torque is increasing rapidly and it means system is unstable. In the case of Fig.(9) (b) red line, system is stable. Fig.(10) depicts the reaction torque from the environment and Fig.(11) is the position response of the robots. Fig.(12) and Fig.(13) represent the inclination angle variation of robot\u2019s body and the centre of gravity position of manipulator part respectively."
        ]
    },
    {
        "image_filename": "designv11_12_0002949_etep.1642-Figure4-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002949_etep.1642-Figure4-1.png",
        "caption": "Figure 4. The search coil positions.",
        "texts": [
            " A dedicated experimental test rig has been set up, which consists of one induction motor and three identical rotors, as shown in Figure 3. The three rotors include a healthy one used as the reference and two broken ones by deliberately drilling holes in their bars on all the depth. In the course of the experiment, each drill hole is checked carefully to make sure the bars were totally broken, and we can see clearly the complete cross section of the rotor bar from the drill hole place. The air-gap flux density is measured by the search coil technology [23,34]. The search coil was inserted around the stator tooth tip (Figure 4). Copyright \u00a9 2012 John Wiley & Sons, Ltd. Euro. Trans. Electr. Power (2012) DOI: 10.1002/etep 4.1. Magnetic field distributions and typical air-gap flux density waveform at standstill Under the locked-rotor conditions, the frequency of rotor e.m.f. equals the supply frequency so that rotor-related skin effects and the saturation of the upper rotor tooth can be more severe than the rated load. Furthermore, the rotor magnetic field distribution is no longer symmetrical upon a broken bar fault, as illustrated in Figures 5 and 6"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003761_9781118818428-Figure5.2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003761_9781118818428-Figure5.2-1.png",
        "caption": "Figure 5.2 Drive mechanism of the Vari-Engine. Reproduced by permission of Ian Larque",
        "texts": [
            " Intuition which has served himwell in related endeavours (a thermal lag engine achieving 3100 no-load rpm) suggested offsetting the effect of decreasing temperature ratio by displacing an increasing fraction of gas between the temperature extremes. In the beta (coaxial) and gamma configurations this is achieved by a decrease in \u03bb (= Sp\u2215Sd). The optimum versus optimization 35 Vaizey collaborates closely with Larque whose \u2018Vari-Engine\u2019 was under construction at the time. The Vari-Engine was introduced in Section 1.2.6 as a small but \u2018serious\u2019 coaxial engine having slotted heat exchangers (as per the PhilipsMP1002CA) and annular regenerator formed by winding a bandage of dimpled foil. Figure 5.2 shows the drive mechanism, consisting of twin parallel crankshafts rotating in a common sense and synchronized by a central idler gear. Piston and displacer are each driven off a radius arm pivoted remote from the cylinder axis. This arrangement minimizes side loads and associated friction. The hollow, cylindrical piston rod runs in a guide in the crank-case below the attachment points of the radius arms. The displacer rod is guided within the piston rod at a point also below the pin attachment point"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003072_s11012-012-9630-6-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003072_s11012-012-9630-6-Figure1-1.png",
        "caption": "Fig. 1 The analyzed hydrodynamic lubricated tilted pad thrust slider bearing",
        "texts": [
            "com Ff,a,Ff,b friction forces per unit contact length at the upper and lower contact surfaces in the bearing respectively F\u0304f,a, F\u0304f,b dimensionless friction forces per unit contact length at the upper and lower contact surfaces respectively fa,conv, fb,conv friction coefficients at the upper and lower contact surfaces in the bearing respectively calculated from conventional hydrodynamic lubrication theory fa,slip, fb,slip friction coefficients at the upper and lower contact surfaces in the bearing respectively when the boundary slippage is introduced h fluid film thickness h0 fluid film thickness at the outlet hi fluid film thickness at the inlet h1 fluid film thickness at x = x1 H h/h0 Hi hi/ho H1 h1/ho Iw relative increase of the carried load of the present bearing compared with the conventional bearing (without boundary slippage) k tan(\u03b8) l1, l2 widths of the no slippage zone and slippage zone of the bearing respectively, Fig. 1(a) lp distance of the pivot of the bearing from the outlet l\u0304p dimensionless distance of the pivot of the bearing from the outlet, lp/(l1 + l2) p fluid film pressure p1 fluid film pressure at x = x1 pconv pressure in the bearing calculated from conventional hydrodynamic lubrication theory P dimensionless pressure, ph0/(u\u03b7a) qv fluid volume flow rate per unit contact length Qv dimensionless volume flow rate of the fluid through the bearing, qv/(uh0) u moving speed of the lower contact surface of the bearing, Fig. 1(a) wconv carried load per unit contact length of the bearing calculated from conventional hydrodynamic lubrication theory wslip carried load per unit contact length of the bearing with boundary slippage W dimensionless load, w/(u\u03b7a) x coordinate x1 x coordinate at the boundary between the \u201cA\u201d and \u201cB\u201d sub zones x2 x coordinate at the inlet X dimensionless coordinate, x/[(l1 + l2) cos \u03b8 ] X1 x1/[(l1 + l2) cos \u03b8 ] \u03c4sa fluid-contact interfacial shear strength at the stationary contact surface in the inlet zone \u03c4\u0304sa dimensionless fluid-contact interfacial shear strength at the stationary contact surface in the inlet zone, \u03c4sah0/[u\u03b7a cos(2\u03b8)] \u03c4sa,B fluid-contact interfacial shear strength at the upper contact surface in the \u201cB\u201d sub zone \u03c4\u0304sa,B \u03c4sa,Bh0/(u\u03b7a) \u03b8 tilted angle of the bearing \u03c4a,A shear stress of the fluid film in the x coordinate direction at the upper contact surface in the \u201cA\u201d sub zone, Fig. 1(a) \u03c4sb fluid-contact interfacial shear strength at the lower contact surface \u03c4\u0304sb \u03c4sbh0/(u\u03b7a) \u03c4s0 u\u03b7a cos(2\u03b8)/h0 \u03b1 h0/(l1 + l2) \u03b7a fluid viscosity at ambient pressure \u03c4 shear stress \u03c4\u0304 dimensionless shear stress, \u03c4h0/(u\u03b7a) ua,x fluid film slipping velocity at the upper contact surface in the \u201cA\u201d sub zone which is in the x coordinate direction Subscript a at the upper contact surface b at the lower contact surface A in the \u201cA\u201d sub zone B in the \u201cB\u201d sub zone slip for the boundary slippage case conv calculated from conventional hydrodynamic lubrication theory, which neglects the boundary slippage Conventional hydrodynamic lubricated titled pad thrust slider bearings have been developed into a mature body",
            " As a further work, the present paper presents a new type of hydrodynamic lubricated tilted pad thrust slider bearings which are artificially augmented with the boundary slippage at the stationary contact surface in the fluid inlet zone. The design is aimed to reduce the friction coefficient but increase the load-carrying capacity of such bearings. A theoretical analysis is presented for this kind of bearing to explore out the bearing performance. The computation is made for typical cases and design guides are given for such bearings. Figure 1(a) demonstrates the configuration of the proposed new type of hydrodynamic lubricated tilted pad thrust slider bearing in the present study. In the figure, the upper contact surface of the bearing is stationary but the lower contact surface of the bearing is moving with the speed u. The lubricated area of the bearing is divided into two sub zones, i.e. the \u201cA\u201d sub zone and the \u201cB\u201d sub zone. The \u201cA\u201d sub zone refers to the bearing inlet zone, while the \u201cB\u201d sub zone refers to the bearing outlet zone",
            " The boundary slippage is augmented at the stationary contact surface in the \u201cA\u201d sub zone where the fluid-contact interfacial shear strength \u03c4sa is significantly low to ensure the occurrence of the boundary slippage. At the other bearing contact surfaces no boundary slippage occurs with sufficiently high fluid-contact interfacial shear strengths. The tilted angle of this bearing is \u03b8 . The used coordinates are also shown. A theoretical analysis is presented in the following for the bearing shown in Fig. 1(a). The analysis is based on the following conditions: (a) The fluid is isoviscous; (b) The fluid is incompressible; (c) The fluid inertia is negligible; (d) The fluid is isothermal; (e) The fluid is in laminar flow. 3.1 Shear stress analysis at the boundary slippage interface Figure 1(b) gives an infinitesimal fluid film element nearby the boundary slippage interface in Fig. 1(a). In this figure, \u03c4sa is the shear strength of the fluid-contact interface where boundary slippage occurs, \u03c4a,A is the shear stress of the fluid film in the x coordinate direction at the upper contact surface in the \u201cA\u201d sub zone, and p is the pressure acting on the element. According to the momentum equilibrium of this element, it is obtained that: \u03c4a,A = \u03c4sa cos(2\u03b8) (1) 3.2 Analysis for the \u201cA\u201d sub zone The Reynolds equation for the \u201cA\u201d sub zone is: dp dx = 3\u03c4sa 2h cos(2\u03b8) \u2212 3u\u03b7a h2 \u2212 3qv\u03b7a h3 (2) where p is fluid film pressure, h is fluid film thickness, \u03b7a is the fluid viscosity at ambient pressure, and qv is the fluid volume flow rate per unit contact length",
            " The friction force per unit contact length at the upper contact surface in the bearing is: Ff,a = \u222b x2 x1 \u03c4a,Adx + \u222b x1 0 \u03c4a,Bdx = \u03c4sa(x2 \u2212 x1) cos(2\u03b8) \u2212 2u\u03b7a k lnH1 + 6qv\u03b7a k ( 1 h0 + x1k \u2212 1 h0 ) (18) The friction force per unit contact length at the lower contact surface in the bearing is: Ff,b = \u222b x2 x1 \u03c4b,Adx + \u222b x1 0 \u03c4b,Bdx = 3u\u03b7a k ln Hi H1 \u2212 \u03c4sa(x2 \u2212 x1) 2 cos(2\u03b8) \u2212 3qv\u03b7a k ( 1 h0 + x2k \u2212 1 h0 + x1k ) + 4u\u03b7a k lnH1 \u2212 6qv\u03b7a k ( 1 h0 + x1k \u2212 1 h0 ) (19) The friction coefficients at the upper and lower contact surfaces in the bearing are respectively: fa,slip = Ff,a wslip and fb,slip = Ff,b wslip (20) The fluid film slipping velocity at the upper contact surface in the \u201cA\u201d sub zone which is in the x coordinate direction is finally: ua,x = h\u03c4sa 4\u03b7a cos(2\u03b8) + u 2 + 3qv 2h (21) 3.7 The pivot position In Fig. 1(a), the distance lp of the pivot of the bearing from the outlet satisfies: lpwslip cos \u03b8 = \u222b x2 0 xpdx = 6u\u03b7a k2 ( x1 \u2212 h0 k lnH1 ) \u2212 3u\u03b7ax 2 1 kh0 + 6qv\u03b7a k3 ( lnH1 + h0 h0 + x1k \u2212 1 ) \u2212 3qv\u03b7ax 2 1 kh2 0 + 3\u03c4sa 2k cos(2\u03b8) [ h0(x2 \u2212 x1) 2k \u2212 x2 1 2 ln H1 Hi \u2212 x2 2 \u2212 x2 1 4 \u2212 h2 0 2k2 ln Hi H1 ] + 3u\u03b7a k2 ( x2 \u2212 x1 \u2212 h0 k ln Hi H1 ) \u2212 3u\u03b7a(x 2 2 \u2212 x2 1) 2k(h0 + x2k) + 3qv\u03b7a 2k3 ( ln Hi H1 + h0 h0 + x2k \u2212 h0 h0 + x1k ) \u2212 3qv\u03b7a(x 2 2 \u2212 x2 1) 4k(h0 + x2k)2 (22) For comparison, this section gives the analytical results for the bearing in Fig. 1(a) when no boundary slippage is considered, according to conventional hydrodynamic lubrication theory [1]. For this case, the carried load per unit contact length of the bearing is: wconv = 6u\u03b7a [ (l1 + l2) cos \u03b8 (Hi \u2212 1)h0 ]2[ lnHi \u2212 2(Hi \u2212 1) Hi + 1 ] (23) where l1 and l2 are respectively the widths of the no slippage zone and slippage zone of the bearing as shown in Fig. 1(a). The friction coefficient at the upper contact surface in the bearing is: fa,conv = \u03b1 lnHi(Hi \u2212 1) 6 cos \u03b8 [lnHi \u2212 2(Hi\u22121) Hi+1 ] \u2212 k 2 (24) where \u03b1 = h0/(l1 + l2). The friction coefficient at the lower contact surface in the bearing is: fb,conv = \u03b1 lnHi(Hi \u2212 1) 6 cos \u03b8 [lnHi \u2212 2(Hi\u22121) Hi+1 ] + k 2 (25) The pressure in the bearing is: pconv = 6u\u03b7a(l1 + l2) cos \u03b8 (Hi \u2212 1)h2 0 \u00d7 [ 1 H \u2212 Hi (1 + Hi)H 2 \u2212 1 1 + Hi ] (26) where H = h/h0. The obtained analytical results are here normalized. The normalized parameters are as follows: W = w u\u03b7a , P = ph0 u\u03b7a , Qv = qv uh0 \u03c4\u0304sa = \u03c4sah0 u\u03b7a cos(2\u03b8) , \u03c4\u0304 = \u03c4h0 u\u03b7a , F\u0304f,a = Ff,a u\u03b7a F\u0304f,b = Ff,b u\u03b7a , DU = ua,x u , l\u0304p = lp l1 + l2 X = x (l1 + l2) cos \u03b8 5.1 Results for the present bearing with boundary slippage For the bearing in Fig. 1(a), the dimensionless volume flow rate of the lubricant is: Qv = 3 2 \u03c4\u0304sa ln H1 Hi \u2212 3( 1 H1 + 1 Hi \u2212 2) 9 2H 2 1 + 3 2H 2 i \u2212 6 (27) The dimensionless pressure in the \u201cA\u201d sub zone is: Pslip = 3\u03c4\u0304sa 2k ln H Hi + 3 k ( 1 H \u2212 1 Hi ) + 3Qv 2k ( 1 H 2 \u2212 1 H 2 i ) (28) The dimensionless pressure in the \u201cB\u201d sub zone is: Pslip = 6 k ( 1 H \u2212 1 ) + 6Qv k ( 1 H 2 \u2212 1 ) (29) The dimensionless carried load of the bearing is: Wslip = 6 k2 lnH1 \u2212 6 k2 X1(Hi \u2212 1) \u2212 6Qv k2 ( 1 H1 \u2212 1 ) \u2212 6 k2 QvX1(Hi \u2212 1) + 3\u03c4\u0304sa 2k2 [ ln Hi H1 \u2212 Hi + 1 \u2212 X1(Hi \u2212 1) ( ln H1 Hi \u2212 1 )] + 3 k2 [ ln Hi H1 \u2212 (1 \u2212 X1)(Hi \u2212 1) Hi ] \u2212 3Qv 2k2 [ 1 Hi \u2212 1 H1 + (1 \u2212 X1)(Hi \u2212 1) H 2 i ] (30) where X1 = x1/[(l1 + l2) cos \u03b8 ]",
            " The dimensionless fluid film slipping velocity at the upper contact surface in the \u201cA\u201d sub zone which is in the x coordinate direction is: DU = \u03c4\u0304saH 4 + 3Qv 2H + 1 2 (37) The condition for hydrodynamic lubrication in the bearing is: \u03c4\u0304sa < cos(2\u03b8) ln H1 Hi [ 2 ( 1 H1 + 1 Hi \u2212 2 ) \u2212 H1 1 + H1 ( 3 H 2 1 + 1 H 2 i \u2212 4 )] (38) The dimensionless distance l\u0304p of the pivot of the bearing from the outlet satisfies: l\u0304pWslip = 6X1 k2 [ 1 \u2212 lnH1 X1(Hi \u2212 1) ] \u2212 3 ( X1 k )2 (Hi \u2212 1)(1 + Qv) + 6Qv k2(Hi \u2212 1) ( lnH1 + 1 H1 \u2212 1 ) + 3\u03c4\u0304sa 2k [ 1 \u2212 X1 2k \u2212 ln Hi H1 2k(Hi \u2212 1) \u2212 X2 1 2k (Hi \u2212 1) ln H1 Hi \u2212 X2 1 4k (Hi \u2212 1) ( 1 X2 1 \u2212 1 )] + 3 k2 [ 1 \u2212 X1 \u2212 ln Hi H1 Hi \u2212 1 ] \u2212 3 2Hi ( X1 k )2 (Hi \u2212 1) ( 1 X2 1 \u2212 1 ) + 3Qv 2k2(Hi \u2212 1) ( ln Hi H1 + 1 Hi \u2212 1 H1 ) \u2212 3Qv 4 ( X1 kHi )2 (Hi \u2212 1) ( 1 X2 1 \u2212 1 ) (39) 5.2 Results for no boundary slippage When no boundary slippage is considered, the normalized results for the bearing in Fig. 1(a) is presented as follows. The dimensionless carried load per unit contact length of the bearing is: Wconv = 6 k2 [ lnHi \u2212 2(Hi \u2212 1) Hi + 1 ] (40) The dimensionless pressure in the bearing is: Pconv = 6 k [ 1 H \u2212 Hi (1 + Hi)H 2 \u2212 1 1 + Hi ] (41) The calculations are made for the bearing in Fig. 1(a) for typical cases for wide operational parameter values. In these calculations, the following dimensionless parameter values are chosen: Hi = 2.2, \u03b1 = 2.5E\u20134. The results are obtained and discussed as follows. 6.1 Hydrodynamic pressure Figures 2(a) and (b) respectively plot the hydrodynamic pressure distributions (Pslip) in the bearing for different values of X1 when the values of the dimensionless interfacial shear strength \u03c4\u0304sa are 0.01 and 0.1. They are compared with the hydrodynamic pressure distribution (Pconv) calculated from conventional hydrodynamic lubrication theory",
            " It is shown that for a given X1 the carried load of the present bearing (or the value of Iw) is linearly increased with reducing \u03c4\u0304sa . For X1 = 0.2, the increase of the carried load of the present bearing by the boundary slippage reaches about 30 % when \u03c4\u0304sa is very low. Figure 3(b) plots the value of Iw against X1 for different \u03c4\u0304sa values. For a given \u03c4\u0304sa , the value of Iw reaches the maximum when X1 is around 0.2. For the maximum load-carrying capacity, the optimum value of X1 of the present bearing appears to be around 0.2. X1 = 0.2 and Hi = 2.2 The optimum condition for the load-carrying capacity of the bearing in Fig. 1(a) with no boundary slippage found from conventional hydrodynamic lubrication theory is Hi = 2.2 [1]. For X1 = 0.2 and Hi = 2.2, the dimensionless pressure in the \u201cA\u201d sub zone in the present bearing is: Pslip = 1.25 [ \u03c4\u0304sa \u03b1 ln H 2.2 + 2 H \u2212 0.91 + (0.3\u03c4\u0304sa \u2212 0.8) ( 1 H 2 \u2212 0.207 )] (42) The dimensionless pressure in the \u201cB\u201d sub zone in the present bearing is: Pslip = 5 \u03b1 [ 1 H \u2212 1 + (0.3\u03c4\u0304sa \u2212 0.8) ( 1 H 2 \u2212 1 )] (43) The dimensionless carried load per unit contact length of the bearing is: Wslip = 1 \u03b12 (0",
            " For comparison, the corresponding analytical results obtained from conventional hydrodynamic lubrication theory are also presented for this type of bearing when no boundary slippage is considered. Typical computational results and design guides for the bearing are given. It is found that the load-carrying capacity of the bearing is increased with both the reductions of the interfacial shear strength \u03c4\u0304sa at the boundary slippage interface and the coordinate X1 between the \u201cA\u201d and \u201cB\u201d sub zones (in Fig. 1(a)) when X1 \u2265 0.2. It is recommended to employ a low value of \u03c4\u0304sa and X1 = 0.2 in design of the bearing for achieving the lowest friction coefficient and the maximum load-carrying capacity. In this condition, the reduction of the friction coefficient of the bearing by the boundary slippage can be more than 40 %, while the increase of the loadcarrying capacity of the bearing by the boundary slippage can be about 30 %. This design recommendation also gives the moderate values of the dimensionless interfacial shear strengths for preventing the interfacial slippage at the remaining contact surfaces of the bearing, which is of importance in engineering"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001284_s10015-006-0401-0-Figure7-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001284_s10015-006-0401-0-Figure7-1.png",
        "caption": "Fig. 7. Inertia force about a hand-type tray",
        "texts": [
            " A characteristic of the proposed method is that some kinds of pattern are derived corresponding to the working time, T. Figure 6 shows the relation between the energy consumed and the release point, under the condition that the distance from the origin to the point of arrival is x = \u22122.0m, the release angle is \u03c6 = 3\u03c0/4, and the working time is T = 0.6s. It is clear that the energy consumed is small at the position of release (\u03b81 = 3\u03c0/8, \u03b82 = \u03c0/8). 4.2 Simulation for holding the object The case when using a hand-type tray is as follows. Figure 7 shows the hand shape of the sector (\u03bb is the angle in the vertical plane). F is the resultant force about gravity and the inertia force, and the components of the force are Fx and Fy. In order to prevent the object falling from the tray, the angle of the tray (\u03b83) is controlled by motor 3 on link 3, as follows: \u03b8 \u03c0 \u03b83 1 22 =     + \u2212\u2212tan F F y x (8) Figure 8a, which looks like a stairway, shows the angular acceleration calculated from the angular velocity (as shown in Fig. 4). The response of the resultant force which acts on the object is influenced by the angular acceleration, as shown in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001326_icalt.2009.85-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001326_icalt.2009.85-Figure1-1.png",
        "caption": "Figure 1 Curriculum Maintenance System",
        "texts": [
            " The knowledgebase system is required to manage and maintain the complex interrelationships between curriculum content, cases for enquiry based learning and their intended learning outcomes (ILOs), and assessment blueprinting based on the ILOs. The system is generic for any structured curriculum but is being tested on the Manchester MBChB curriculum, using Index Clinical Situations, Module ILOs, Formative and Summative Assessment and Problem Based Learning (PBL) cases [4]. The principal users of the system, illustrated in Figure 1, will be tutors in the academic curriculum development teams and students who will use appropriate search facilities to plan their learning. In addition, administrators will use the system for planning programme delivery. 978-0-7695-3711-5/09 $25.00 \u00a9 2009 IEEE DOI 10.1109/ICALT.2009.85 136 The project test case is the curriculum for undergraduate medical students at the Manchester Medical School which is centred on a Problem Based Learning (PBL) approach to learning. The curriculum consists of a number of themed modules each having its own ILOs and set of clinical skills"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002335_detc2010-28926-Figure7-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002335_detc2010-28926-Figure7-1.png",
        "caption": "Figure 7. Analysis process of a mechanism through the decomposition into Assur Graphs.",
        "texts": [
            " a) The linkage. b) The structural scheme. c) The decomposition graph. In this example, the first Assur Graph to be analyzed is the tetrad \u2013 (B,C,D,J) where the two outer vertices, A and p4, are ground vertices thus the inner velocities \u2013 B,C,D,E,J can be calculated. The second AG that can be analyzed is (G,H,I) or the dyad F. In this example, the second AG chosen, arbitrarily, to be analyzed is the dyad F, where now the velocity of the outer vertex \u2013 J is known from the previous AG, as shown in Figure 7b,b1. The last AG is the triad (G,H,I) whose velocity of the outer vertex E is known from the first AG \u2013 (B,C,D,J). 4 Copyright \u00a9 2010 by ASME Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 08/12/2018 Terms of Use: http://www.asme.org/about-asme/terms-of-use 5 Copyright \u00a9 2010 by ASME The analysis process is done somewhat in a similar way related to the analysis of velocities in kinematics, but this time in a reversed order, i.e., in the decomposition order The process is as follows: 1"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002026_6.2009-1802-Figure4-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002026_6.2009-1802-Figure4-1.png",
        "caption": "Fig. 4 - Left Wing Damaged Generic Transport Model",
        "texts": [
            " (12) yields P = [ K\u22121 i Kp + K\u22121 p (Ki + I) K\u22121 i K\u22121 i K\u22121 p ( I + K\u22121 i ) ] > 0 (88) A\u22121 m is computed to be A\u22121 m = [ \u2212K\u22121 i Kp \u2212K\u22121 i I 0 ] (89) Evaluating the term B>PA\u22121 m B yields B>PA\u22121 m B = \u2212K\u22122 i < 0 (90) Applying the adaptive optimal control modification (11), the weight update law is then given by \u0398\u03071 = \u2212\u0393\u03a6 ( e>PB + \u03bd\u03a6>\u03981K\u22122 i ) (91) IV. Simulation Results To evaluate the adaptive optimal control modification, a simulation was conducted using a generic transport model (GTM) which represents a notational twin-engine transport aircraft as shown in Fig. 4.24 An aerodynamic model of the damaged aircraft is created using a vortex lattice method to estimate aerodynamic coefficients, and stability and control derivatives. For the simulation, a damage configuration is modeled corresponding to a 28% loss of the left wing. The damage causes an estimated C.G. shift mostly along the pitch axis with \u2206y = 0.0388c\u0304 and an estimated mass loss of 1.2%. The principal moment of inertia about the roll axis is reduced by 12%, while changes in the inertia values in the other two axes are not as significant"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001739_icc.2009.5198621-Figure3-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001739_icc.2009.5198621-Figure3-1.png",
        "caption": "Fig. 3. Estimation of Selection Error",
        "texts": [
            " Sensors that receive this message stop their timers and use this message among other messages they receive from other skeleton sensors to determine their hexagon. Although collisions are unlikely to happen, we are still able to break ties by allowing colliding sensors to compete in another countdown round starting from a randomly selecting value. We show how a sensor can estimate ei, the distance between itself and the center of the target hexagon identified by the tuple \u3008s, r, c\u3009. We distinguish between two different cases. Case I: this case is applicable only for the six skeleton sensors in the first row of each sector. Figure 3(a) shows ei when sensor S is selected to represent the hexagon \u3008si, 1, 1\u3009. From the figure, ei can be evaluated as, ei = \u221a tx 2 + d2 \u2212 2 \u00b7 tx \u00b7 d \u00b7 cos(\u03c6i \u2212 \u03b8i) (5) Where d is the distance between the sink and the sensor and is estimated using RSSI. Case II: this case is applicable for all skeleton sensors other than those handled by Case I. As shown in figure 3(b), we assume the existence of another skeleton sensor S0 that had been previously selected by the protocol to represent the hexagon S\u3008s0,r0,c0\u3009. The selection error of sensor S0 is e0 and represents the distance between S0 and the center of the hexagon S\u3008s0,r0,c0\u3009. Based on the selection rules we described earlier, it is S0 turn to select sensor S1 to represent the hexagon S\u3008s1,r1,c1\u3009. S1 is selected such that the selection error e1 is minimum (e1 is the distance between S1 and the center of the hexagon S\u3008s1,r1,c1\u3009)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001521_s0022-0728(72)80104-9-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001521_s0022-0728(72)80104-9-Figure1-1.png",
        "caption": "Fig. 1. Cell for constant-potential electrolysis.",
        "texts": [
            " Measurements on thefirst drop were carried out by applying a potential step which was manually synchronized with the fall of the drop. Cell for polaroyraphic measurements. A Metrohm water jacketed cell EA 876 was used for polarographic measurements. A commercial calomel electrode was used as J. Electroanal. Chem., 38 (1972) reference. A platinum wire dipping directly in the solution was used as a counter electrode. A polarographic capillary with enforced fall of the drop at -c = 2.65 s was used; hug = 60 cm, m = 1.026 mg s-1. Cell for constant potential electrolysis (Fig. 1). The cell insures uniform current distribution on the mercury pool electrode. Thus, undesirable electrochemical processes are prevented and an accurate measurement of n is possible. The cylindrical working compartment is isolated by an ion exchange membrane or a thirsty glass separator (Corning, type 7930 glass) from the anode compartment. The working and auxiliary electrodes have the same area and are parallel to each other. The reference electrode (Radiometer K 4018) is a calomel electrode which is connected to the cathode compartment via a Tygon tube filled with saturated KC1",
            " It is applicable also when the current is kinetically controlled provided that il is proportional to the concentration of reactant in solution, as is the case here. It is very important to achieve a uniform current distribution on the working electrode in this experiment, otherwise the metal/solution potential difference MASq~ will vary on different parts of the electrode, giving rise to side reactions (in particular, hydrogen evolution) which lead to erroneous values of n. For this purpose the special cell described above (cf Fig. 1) was devised. The average of several experiments using different concentrations of GA gave n = 2.1 _+ 0.1. 1.2 Followin 9 the change in concentration in solution durin9 constant-potential electrolysis. In these experiments the same cell was used and electrolysis was carried out at the same potential, to the extent of 10-20~ conversion of the reactants. The change of concentration in solution was followed polarographically and the charge was determined by integrating the current/time plot observed with the mercury pool working electrode"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001884_s11431-010-3100-y-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001884_s11431-010-3100-y-Figure2-1.png",
        "caption": "Figure 2 ED\uff1dFG=HK, ED//FG//HK.",
        "texts": [
            " So we define some new concepts: link-group rank, generalized-pair rank, virtual pair, virtual loop, virtual-loop rank. We then propose a general formula expressed by the virtual-loop rank to look for a general mobility formula. To simplify analysis, a link group is regarded as the combination of non-coincident links between any independent loop and its adjacent loop. The mobility of a link group can be equal to 0, or less than 0, or more than 0. The links of the link group can be driven or motive. The link group defined here is a generalized group, not the Assur group. For the planar 9-bar linkage shown in Figure 2, it contains four independent loops, and has four link groups, ABCD, EFG, HK and PRM, as shown in Figure 3. To describe the motion transmission manner of two adjacent loops by a terminology, we define the concept of virtual kinematic pair. Assume that links A and B are responsible for the motion transmission between two adjacent loops. No matter whether they are adjacent, it is supposed that they are connected by a kinematic pair, which is regarded as a virtual kinematic pair, short for virtual pair"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003509_j.proeng.2013.08.203-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003509_j.proeng.2013.08.203-Figure2-1.png",
        "caption": "Fig. 2. (a) Robotic architecture including the hybrid robot with a parallelogram closed loop, (b) angular variable of the architecture, (c) equation to obtain the serial equivalence",
        "texts": [
            " So, the performances of two industrial robots embedded in a kinematically redundant robotic cell dedicated to machining tasks are evaluated. The different models are then detailed and a new procedure for managing kinematic redundancy whilst integrating various criteria is proposed. Simulation and first results are finally presented to assess the performances of the two different architectures (Fig. 1). The modeling of the hybrid robot with a closed parallelogram loop has been largely presented in Subrin et al. (2011) and the one with parallel architecture in Robin et al. (2011) (Fig. 2a, 3a). These architectures are modeled like a serial architecture. The model includes movement reversing relative to the movement of the rotary table. This is equivalent to positioning the observer on the rotary table instead of on the base as usual. The solution of the Direct Geometric Model (DGM) is obtained by multiplying the homogeneous operators associated with TCS method (Gogu et al. (1997)). Concerning the robot with parallel architecture, the kinematic chain includes 3 prismatic joints. To model the architecture, an equivalent structure (a serial one) which includes 2 revolute joints ( , ) and 1 prismatic joint (r) is taken into account (Robin et al"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002425_vppc.2012.6422784-Figure4-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002425_vppc.2012.6422784-Figure4-1.png",
        "caption": "Fig. 4. (a) FEM simulation model and generated mesh of the PMSM with a phase winding fault, (b) external circuit for the time based FEM simulation.",
        "texts": [
            "t\u00b7 = COSOJt+ 2 2 2 SITIOJt a2OJ2 +b2 a OJ +b (20) The coupled flux linkage from if to the healthy winding [Adf\\ Aqf'y is derived as From (12), (13), (22), and (23), we obtain the full dynamic model of the PMSM under the a-phase winding turn short such that If the negative sequence current controller regulates the negative sequence current I \ufffdq, to zero, the negative sequence voltage is calculated as V- - OJ- +--I'!. [0] I'!.mOJ [ \u00b0 dqe 3 Iff III I 3 -1 L I'!. m + L III 1 ( (L )J L OJ [ bd+aeOJ ] 3\" f- 2 I b2+a2OJ2 -be+adOJ (26) Assuming that the three phase current is balanced and d-axis current in the positive sequence SRF is zero, i.e., I;e = I\ufffde = I;e = 0 , the torque equation is obtained as follows III. SIMULATION RESULT (27) To validate the model, FEM simulations are performed. System parameters are shown in Table 1. Fig. 4 shows a FEM mesh model and an external circuit model to make the a phase tum short fault state. In Fig. 4(b), the a-phase is composed of two winding. The left one is healthy and the right one is faulted when the switch is turned on. Balanced three phase current sources are connected to the windings. At the beginning, the motor is healthy. The FEM simulation is conducted for the healthy state. After one rotor rotation ( () > 2Jf), the switch is turned on and the FEM simulation is conducted for the fault state. Fig. S shows the comparison of the FEM simulation results and dynamic model simulation results when the motor speed is 3000rpm and the peak value of the balance phase current is SA"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001978_s10846-010-9408-9-Figure15-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001978_s10846-010-9408-9-Figure15-1.png",
        "caption": "Fig. 15 Model of the tracked",
        "texts": [
            " The WSN system involves a host computer, a location dongle connected to the host computer through serial port (RS232), eight reference nodes, a blind node installed on the robot. During the experiment, on the one hand the robot communicates with the WSN system and receives the controlling instructions from the host computer through the blind node; on the other hand the robot gets the environment information surrounding the robot through WSN system and realizes the self-location. 6.2 Model of Tracked Mobile Robot [17] In the experiment, a tracked mobile robot is adopted. The model of the robot is shown in Fig. 15, where x\u2013o\u2013y and x0\u2013o0\u2013y0 are the global and body coordinate systems respectively. vl and vr are the velocities of the left and right tracks. \u03b8 l and \u03b8 r are the rotating angles of left and right driving wheels. v is the velocity at the center of tracked robot. b is the distance between the left and right wheels. l is the distance between the centers of wheel and robot. R is the radius of driving wheel. \u03b8 is the orientation of the robot with respect to the global coordinate system. Supposing that the sideslips of inner and outer tracks are synchronous when the robot turns and there is no any displacement between the wheel and track of crawler mechanism"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003690_10255842.2012.713646-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003690_10255842.2012.713646-Figure1-1.png",
        "caption": "Figure 1. JCS and SCS axes and origins.",
        "texts": [
            " 2002), aH is the FE axis and cH is the IER axis (along with the longitudinal direction of the femur). For the knee JCS (Grood and Suntay 1983), aK is the FE axis (e.g. computed functionally) and cK is the IER axis (along with the longitudinal direction of the tibia\u2013 fibula complex). For the ankle JCS (Cole et al. 1993), aA is the FE axis (e.g. computed functionally) and cA is the AA axis (along with the longitudinal direction of the foot). Finally, bH, bK and bA are the floating axes. The axes cH and aK are both embedded in the thigh but are not orthogonal (Figure 1). A first SCS (denoted i \u00bc 3) can be classically defined (Cappozzo et al. 1995; Wu et al. 2002) with Y3 \u00bc cH, X3 normal to the frontal plane, Z3 \u00bc X3 \u00a3 Y3 and the origin P3 fixed at the hip joint centre. But another SCS (i \u00bc 3*) is required for the knee kinematics, with Z3* \u00bc aK and the origin D3 fixed at the knee joint centre (KJC). The other axes may be deduced as Y3* \u00bc aK \u00a3 X3 and X3* \u00bc Y3* \u00a3 Z3*. The transformation T3!3* from the first to the second thigh SCS is constant (e.g. obtained in a calibration procedure)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003607_j.mechatronics.2013.05.005-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003607_j.mechatronics.2013.05.005-Figure1-1.png",
        "caption": "Fig. 1. Motion control view during wafer scanning.",
        "texts": [
            " The discussion is centered around the wafer stage dynamics and the problem to be addressed: the loss of throughput during the wafer stage (chuck) exchange. In Section 3, the Levenberg\u2013Marquardt algorithm along with the derivation of the sampled-data gradient error signals will be given. In Section 4, the results of the optimization scheme applied to the controlled wafer stage will be presented. In Section 5, the main conclusions will be summarized. Integral sliding mode control (SMC) will be studied as a potential solution for the chuck exchange problem. To explain this, consider Fig. 1 which shows a motion control view during wafer scanning, i.e. the process of making chips. To create an image of a chip, (extreme) ultraviolet light passes a mask and an optical column before exposing the photo-sensitive layers of the wafer Please cite this article in press as: Heertjes M, Verstappen R. Self-tuning in inte tronics (2013), http://dx.doi.org/10.1016/j.mechatronics.2013.05.005 surface. During wafer exposure, a second wafer is being measured and prepared for so-called chuck exchange. The chuck exchange problem involves swapping the wafer stage modules from exposure side to measurement side (and vice versa) in minimal time with the aim to achieve high wafer throughput. During the chuck exchange the wafer stage modules perform large scale motion commanded by various trajectories. In counter clockwise direction, for example, the upper trajectory of the exposure module (see Fig. 1) differs from the lower trajectory of the measurement module, this to prevent module collision. Other trajectories involve material handling operations like wafer load and unload. The lack of recurrence between these trajectories is one of the reasons to not consider feedforward control techniques like iterative learning control (ILC). Contrarily, if recurrence exists, ILC can be very effective in wafer scanning control, see for example Mishra et al. [36]. Another reason to not consider feedforward control techniques is given by the position-dependent dynamics during the chuck exchange for which an accurate plant model is often lacking"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001968_09507110902836879-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001968_09507110902836879-Figure1-1.png",
        "caption": "Figure 1. Schematic set-up of laser beam and wire.",
        "texts": [
            " The MIG arc welder used was an invertercontrolled DC digital pulsed MIG arc welder. The fibre that carried the YAG laser was a 0.6-mm diameter SI type. The focal length of both the collimating and condensing lens was 200 mm, and the beam diameter at the focal point was 0.6 mm. The fibre that carried the fibre laser was a 0.1-mm diameter SI type. The focal lengths of the collimating and condensing lens were, respectively, 125 and 250 mm, and the beam diameter at the focal point was 0.2 mm. The arrangement of the laser beam and the wire are shown in Figure 1 ((a) the present method: distance between the laser irradiation point and the wire aim point, DL , 0 and (b) previous method: DL ^ 0). Bead-on-plate welding was carried out using both methods, the wire melting characteristics during laser irradiation were examined and the butt weld permissible gap was investigated. The welding conditions used are shown in In order to investigate the effects of the beam diameter, examination was made using a fibre laser with the beam diameter varing from a diameter of 0",
            " The reason for the generation of this dispersal, is that, since the A5356 wire used has 5% magnesium content, when the droplet at a high temperature is irradiated by the laser beam and heated until it reaches its boiling point, it vaporizes violently. The reason why the droplet tends to vaporize more readily at defocusing Df . 0 than at Df , 0 is explained below with reference to Figure 9. Figure 9(a) shows the measured values (1/e 2) for the beam radius of the fibre laser used in these experiments and an approximation found by assuming that propagation follows a hyperbolic curve. Figure 9(b), which is calculated from the positional relationship between the laser beam and the wire shown in Figure 1 and the approximation equation for the laser beam propagation shown in Figure 9(a), shows the positional relationship between the laser beam and wire on the base metal surface. It is clear that the beam diameter of the laser beam directly irradiated onto the surface of the wire at defocusing at DF . 0 is smaller than when this is DF , 0. This means that the power density of the laser beam at the wire surface with defocusing at Df . 0 became higher than when Df , 0. Thus, even when the beam diameter at the base metal surface is the same, for defocusing Df "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000880_tmag.2007.891399-Figure3-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000880_tmag.2007.891399-Figure3-1.png",
        "caption": "Fig. 3. Analyzed model of EMI drive mechanism. (a) Overview; (b) conductor; (c) ring magnet.",
        "texts": [
            " It is assumed the oscillation is expressed as follows: (5) where is the amplitude of oscillation, and is the phase difference. The magnification factor which is the relationship between the dynamic and the static amplitude is derived from the (4) and (5) as follows: (6) where is the natural angular frequency, and is the damping ratio. The magnification factor becomes the maximum at the natural angular frequency. The EMI torque can be obtained from the transient torque amplified by the resonance. Fig. 3 shows the analyzed model of EMI drive mechanism. It mainly consists of the stator part and the rotation ring magnet. The stator part has 8 thin conductors, and inner and outer stator cores. The conductors have the width of 0.8 mm, the conductivity of , and are fixed on the inner and outer stator cores in steps of 45 , respectively. The ring magnet has 8 poles, and is inserted in the gap between the inner and outer conductors. The air gaps of both sides between the conductors and the ring magnet are 0"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002643_j100858a084-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002643_j100858a084-Figure1-1.png",
        "caption": "Figure 1. The rotating-ring-disk electrode, the electrochemical cell, and the auxiliary optical system used in the study of electrogenerated chemiluminescence.",
        "texts": [
            "0 G Figure 1. the dark: bleaching with a 300-W tungsten lamp using a Toshiba IR-DlB filter. with a microwave power of about 0.03 mW (The wavelength ranged from 1 t o 3~.) Esr spectra of n-decane y irradiated a t 77\u00b0K in (a) immediately after irradiation; (b) after Measurements were made at 77\u00b0K in the dark was thermally stable a t 77\u201dK, and the appreciable decrease of the intensity was not found during 0.5 hr. The yield of the trapped electrons was of comparable order with that for the 3-methylpentane or 3-methylhexane glasses, which were purified by the same procedures and were irradiated at the same condition, but the sensitivity to the infrared light illumination seems higher in the polycrystalline case of n-decane, although the quantitative comparison is not made yet",
            " The Journal of Physical Chemistry deactivate the excited state of the chemiluminescent product), Because there is no alternation of potential a t either electrode, charging currents are negligible once the steady state is attained. As an added advantage, the typically larger electrode area of the rotating ringdisk results in greater quantities of chemiluminescent emission. The purpose of this communication, then, is to report the use of a rotating-ring-disk electrode for the study of ECL phenomem in an attempt to overcome the difficulties encountered in the single-electrode experiment. The cell shown in Figure 1 has features which include an exterior compartment in which either a platinum or a mercury-pool auxiliary electrode is separated from an interior compartment with fritted disks, a Luggin compartment extending into the interior compartment and fitted with an aqueous saturated calomel reference electrode, and a side arm for the attachment of the sealed cell to a vacuum-helium line for the purposes of solution deaeration.8 The ring-disk electrode assembly shown consists of a platinum ring and disk separated by a 0"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001868_elan.200800014-Figure8-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001868_elan.200800014-Figure8-1.png",
        "caption": "Fig. 8. Reaction Schemes at the anode and cathode of the biofuel cell and a picture of a miniature biofuel cell on a U.S. penny. Reprinted from [71]; copyright 2006 Wiley InterScience",
        "texts": [
            " Although the current obtained dropped very fast with operation time, these biofuel fuel cells reached power outputs in the order of 300 mW/cm2. [71] Tests with different types of buffers were performed and their importance in influencing the performance and stability of biofuel cells was demonstrated. PEMbased biofuel cells offer an interesting way of improving power outputs and new approaches to extending the cell lifetime, which are key-points for the development of Electroanalysis 2010, 22, No. 7-8, 732 \u2013 743 2010 Wiley-VCH Verlag GmbH & Co. KGaA, Weinheim www.electroanalysis.wiley-vch.de 739 implantable enzymatic power devices. Figure 8 is a schematic of the reactions that take place in the PEM biofuel cell and shows this cell on a United States penny coin. An interesting variation in biofuel cell design can be found in a device where the output power can be tuneable and switched between the ON state and OFF state with a shortcircuit current of 550 mA/cm2 [72]. The biocatalysts were integrated in a copper-poly (acrylic acid) hybrid matrix. The formation of the Cu0 state and Cu2\u00fe state controls the reversible activation and deactivation yielding the ON biofuel cell and OFF biofuel cell, respectively"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000802_1.2908921-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000802_1.2908921-Figure1-1.png",
        "caption": "Fig. 1 Ball-ring elliptical contact",
        "texts": [
            " The total stiffness coefficient involves both the inner and outer race contacts and can thus be expressed as follows: Kc tot = 1 Kc in 2/3 + 1 Kc out 2/3 \u22123/2 5 The inner and outer race contact stiffness coefficients Kc in and Kc out, for the elliptical contact conjunction between two solids can be calculated using the generalized expressions for the elliptic integrals and ellipticity parameter as follows: Kc = k\u0304eE R\u0304 4.5\u03043 6 where the ellipticity parameter k\u0304e and the effective modulus of elasticity E can be defined as follows 19 : k\u0304e = 1.0339 Rz Rx 0.6360 7 1 E = 1 2 1 \u2212 a 2 Ea + 1 \u2212 b 2 Eb 8 where is Poisson\u2019s ratio of the bearing material and subscripts a and b refer to solids a and b, respectively see Fig. 1 . Curvature sum R presented in Eq. 6 can be defined as follows 18 : 1 R = 1 Rx + 1 Rz = 1 rax + 1 rbx + 1 raz + 1 rbz 9 where Rx and Rz represent the effective radii of the curvature in the Y-Z and Y-X planes. In Eq. 9 , rax, raz, rbx, and rbz are radii of curvature of two solids a and b in two directions, as shown in Fig. Transactions of the ASME ata/journals/jotre9/28757/ on 02/06/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use 1 d d s l B p c X i w f w t R b J Downloaded Fr . It is important to note that parameters rbx and rbz are negative ue to concave surfaces in the ball-ring elliptical contact"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002811_1754337112441112-Figure3-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002811_1754337112441112-Figure3-1.png",
        "caption": "Figure 3. Instrumentation diagram of the apparatus used to measure dynamic impact response of softball foam samples.",
        "texts": [
            " The initial elastic region resists compression with the inherent stiffness of the foam structure. At a critical strain, the foam structure begins to collapse, causing a plateau region in the stress\u2013strain response. As the cells become fully compressed, a densification region begins and the material stiffens, behaving much like the matrix material. A single curve representing the compressive loading response of the PU foam used in softballs was needed to numerically model the foam material. Thus, a method was developed to obtain a master compressive loading curve. The apparatus shown in Figure 3 was constructed to impact foam samples, and can be used to test other polymers or similarly compliant materials. The device used a small, horizontally-mounted air cannon that achieved a larger range in impact speed than is possible with standard drop towers.18 When the cannon was fired, a pneumatic valve released compressed air from an accumulator tank to a 610mm long barrel. An at Virginia Tech on November 11, 2014pip.sagepub.comDownloaded from 18.3 g aluminum striker bar, 13mm in diameter and 51mm long, was projected out of the barrel and speed was measured as it passed in front of infrared sensors spaced 38mm apart"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003821_ls.1184-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003821_ls.1184-Figure1-1.png",
        "caption": "Figure 1. Schematic view of the aerostatic porous bearing.1",
        "texts": [
            " Monte Carlo method is used to introduce variability in the solutions. The purpose of this work was to check the effects of such variability on the predictions of bearing loading capacity. Copyright \u00a9 2012 John Wiley & Sons, Ltd. Lubrication Science (2012) DOI: 10.1002/ls Aerostatic porous bearings are air-lubricated bearings, where the supporting mechanism of the shaft is pressurised air. Pressurised air is injected through the porous matrix (bearing casing) and defines the fluid pressure distribution in the rotor-bearing interface (Figure 1). The modified Reynolds equation that represents the fluid behaviour in the rotor-bearing interface of an aerostatic porous bearing is given by1: 2Lt L 2 @ @ x p h3 @ p @ x \u00fe @ @ y p h3 @ p @ y \u00bc \u039b @ p h\u00f0 \u00de @ y \u00fe\u03a8 @ p h\u00f0 \u00de @t \u00fe\u0393 p p\u00fe 1 \u03a6 ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi \u03a62 \u00fe 2\u03a6 1 p2\u00f0 \u00de q \u00de (1) where L and Lt are the bearing width and the inner perimeter, respectively; p is the non-dimensional pressure (p/Ps); h is the non-dimensional clearance (h/c); c is the assembled clearance; Ps is the air supply pressure; t is the non-dimensional time (ot); and x and y are non-dimensional coordinates in axial (2x/L) and tangential (y/Lt) directions, respectively"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001546_j.jmatprotec.2010.12.002-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001546_j.jmatprotec.2010.12.002-Figure1-1.png",
        "caption": "Fig. 1. View of laser welding process.",
        "texts": [
            " The moving heat source results in ocalized heat generation and large thermal gradients. In addition, he thermo-physical properties of the material depend on the temerature. .1. Thermal analysis Consequently, for thermal analysis, the transient temperature eld T of the plate is a function of time t and the spatial coordiates (x, y, z), and is determined by the non-linear heat diffusion quation: (T)Cp(T) \u2202T \u2202t = \u2202 \u2202x ( k(T) \u2202T \u2202x ) + \u2202 \u2202y ( k(T) \u2202T \u2202y ) + \u2202 \u2202z ( k(T) \u2202T \u2202z ) + q (1) here x, y and z are the axes as shown in Fig. 1, T is the temperature, t s time, and q is the heat generation rate per unit volume. (T), Cp(T) nd k(T) are the temperature dependent density, specific heat and onductivity of the plate material, respectively. Heat losses from all surfaces of the tube by convection qc and radiation qr are introduced as boundary conditions, using the following equation: qc = h(T \u2212 T\u221e) (2) qr = \u00d7\u02d9(T4 \u2212 T4 \u221e) (3) where h is the convection coefficient, T\u221e is the room temperature, is the emissivity of the body\u2019s surface ( = 0"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003942_j.isatra.2015.09.003-Figure17-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003942_j.isatra.2015.09.003-Figure17-1.png",
        "caption": "Fig. 17. Sketch of TQ MA3000 5-DOF robotic manipulator.",
        "texts": [
            " The tracking control achieved in this example portrays the efficient performance of the proposed controller in the presence of the measurement noise. In this section, we consider a real-time application of the proposed LNT controller. A decentralized scheme for position tracking Please cite this article as: Aftab MS, Shafiq M. Neural networks for tra ISA Transactions (2015), http://dx.doi.org/10.1016/j.isatra.2015.09.003 of the major joints of TQ MA3000 robotic manipulator has been implemented. TQ MA3000 is a 5-DOF industrial robotic system, developed by TecQuipment Limited. As shown in Fig. 17, this robot has three major and two minor axes of rotation. The angular positions associated with these axes and their corresponding characteristics are given in Table 1. The DC motors connected at the joints of this robotic system are manufactured by Bodine Electric Company. These motors are characterized by low speed and high torque throughput, with very high gear-ratio. Brush type analog PWM servo amplifiers (Model 30A8C manufactured by Advanced Motion Controls) are used to drive the joint DC motors"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003713_978-1-4614-2419-2-Figure54.1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003713_978-1-4614-2419-2-Figure54.1-1.png",
        "caption": "Fig. 54.1 Geometric model of the containment",
        "texts": [
            " The springing point was found to experience lesser damage due to higher stiffness in this region. The numerical simulation of Boeing 707\u2013320 aircraft crash on the nuclear containment was carried out using ABAQUS/ Explicit-6.9.3 finite element code. A three-dimensional model of the containment was made using preprocessing module of the code. Its geometry and reinforcement detailing was considered identical to that of the containment used by Abbas et al. [4, 5]. The containment structure had a cylindrical wall and a spherical dome with 1.2 m uniform thickness, Fig. 54.1. The structure was doubly reinforced with \u00d840 mm bars placed at 80 mm c/c both ways at the inner and outer faces of the cylindrical wall as well as the spherical dome. The effective cover to the concrete was assumed to be 100 mm. The reinforcement modeled as 3D wire was placed in the structure using linear/radial pattern option available in ABAQUS/CAE. The contact between the steel and concrete was modeled using embedded element technique available in the code. The containment was considered to be fixed at its base",
            " Figure 54.7 compares the deformation profile obtained in the present study with that of the Abbas et al. [5]. A close correlation between the predicted profiles of deformation as well as the peak deflection was found. The peak deflection occurred at 0.247 s as per the study carried out by Abbas et al., [5] and at 0.25 s according to the present investigation. The impact locations i.e., the junction of dome and cylinder and the midpoint of cylindrical portion were designated as location \u201cA\u201d and \u201cB\u201d, Fig. 54.1. The deformation of the containment was found to be confined to the impact zone at both the locations. However, it was more localised at location \u201cB\u201d. The intensity of deformation was also found to be higher at location \u201cB\u201d. This is due to the fact that the structural stiffness is higher at location \u201cA\u201d being junction point of dome and cylinder. It should be noticed that the dome and cylinder were modeled monolithic to each other in this study. Abbas et al., [4] also modeled the dome monolithic to that of the cylindrical portion of the containment"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002241_978-1-4471-4141-9_70-Figure70.7-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002241_978-1-4471-4141-9_70-Figure70.7-1.png",
        "caption": "Fig. 70.7 Parameters of the tripod mechanism. a The triangular prism profile, b lateral view of the prism",
        "texts": [
            " Then from D\u2013H loop equation, we have the closure equation: cos h2 cos h4 cos h3 sin h2 sin h4 \u00bc cos h6 \u00f070:11\u00de In Eqs. (70.9) and (70.10), there are three equations what include four variables, that can also prove that the general line-symmetric 6R mechanism has only one DOF. In this research, it is assumed that the lateral profile of the triangular prism is isosceles trapezoid, i.e., all the Ls s in one module are identical. 1z 2z 3z 4z 5z 6z D 1 H F E IG P O Q M Fig. 70.6 Geometry of general line-symmetric 6R mechanism From Fig. 70.7, one can see that Lu \u00bc 2L2 sin a=2\u00f0 \u00de \u00fe 2L1 cos c cos b \u00f070:12\u00de P is the concurrent point of the three axes z1; z3; z5; PM ! and PN ! are the two vectors along z3 and z1 respectively, a \u00bc \\GPH; a\u00fe h2 \u00bc 2p \u00f070:13\u00de L2 contains two parts, link length of Bricard Lc1; second part is length of PG; PH; PI Lc2 \u00bc di tan h2=2\u00f0 \u00de \u00f070:14\u00de L1 is the width of the lateral link, c is the obliquity of the lateral link, b is the angle of GH with respect to the radius of circumcircle of MGHI: The edges of MGHI are 2L0 sin h2=2\u00f0 \u00deL; 2L0 sin h4=2\u00f0 \u00de and 2L0 sin h6=2\u00f0 \u00de respectively. L0 is length of PG. h2; h4; h6 can be derived from Eq. (70.10). Given edges l1; l2; l3 for one triangle, the radius of its circumcircle can be given as r2 \u00bc l1l2l3\u00f0 \u00de2= 2 l21l22 \u00fe l2 2l2 3 \u00fe l21l23 l4 1 \u00fe l4 2 \u00fe l43 \u00f070:15\u00de Then b can be calculated by cos b \u00bc li=2r \u00f070:16\u00de where li is edge of triangle, using Eqs. (70.12\u201370.16), one can obtain the relation of b and h2: From Fig. 70.7b, we can obtain the relation: cos c \u00bc ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 1 cos /=cos b\u00f0 \u00de2 q \u00f070:17\u00de where / is the angle of Lu and Ls, then we can derive the relation of Lu and h2: The relation for the other base and the whole tripod mechanism can be derived using similar approach. Based on the conceptual trussed surface structure given in Fig. 70.4, one can see that in order to construct a modular trussed structure, every two adjacent triangular prisms are sharing one common lateral profiles, in order to build a \u2018\u2018continuous and smooth\u2019\u2019 surface, the lateral profile of one module that is used to connected to the adjacent one must be identical, therefore, as soon as the parameters of one module, i"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002400_1.4006324-Figure4-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002400_1.4006324-Figure4-1.png",
        "caption": "Fig. 4 Contour line graphs of tooth contact stress distributions of left web gears (unit: MPa)",
        "texts": [
            " LTCA is conducted for the three types of thin-rimmed gears shown in Fig. 2, when these gears have both straight webs and inclined webs, and they are engaged with the solid mating gear at the highest point of the single pair tooth contact. The tooth contact stresses are calculated under a torque load of 294 N m for all of the cases in the paper. 4.1 Tooth Contact Stress Distributions of the Thin-Rimmed Left Web Gears. At the time when the webs are located at the left side of the tooth, the tooth contact stresses are calculated. Figure 4 shows the results of the tooth contact stresses of the left web gears. In Fig. 4, the abscissa is the tooth longitudinal dimension and the ordinate is the contact width WH as shown in Fig. 1. The web position and the geometrical contact line are also shown in Fig. 4(a). Figures 4(a)\u20134(d); are the contour line graphs of the contact stresses of the gears when the web angles are 0, 15, 30, and 45 deg, respectively. The maximum contact stresses of the teeth are indicated in Fig. 4, and the resulting relationship between the maximum contact stress and the web angle is summarized in Table 1 and is shown in Fig. 5. In Table 1 and Fig. 5, the contact stress when web angle is 0 deg is used as the standard value, and the rate of change is defined as the change in the contact stress when the web is inclined relative to this standard value. From Fig. 4, it is apparent that the maximum contact length of the teeth becomes shorter as the web angle increases. This finding also indicates that the partial tooth contact becomes more significant as the web angle becomes larger. From Table 1 and Fig. 5, it is apparent that the maximum contact stress increases by 8% when the web is inclined from a straight state to a 15-deg web angle. It is also observed that the maximum contact stress increases with increasing increments in the web angle. The relationship between the maximum contact stress and the web angle is nonlinear for the left web gears",
            "org/about-asme/terms-of-use Figures 8(a)\u20138(d) are the results for the cases when the web angles are 0, 15, 30, and 45 deg, respectively. The relationship between the maximum contact stress and the web angle is summarized in Table 3 and is shown in Fig. 9. From Fig. 9, it is apparent that the relationship between the maximum contact stress and the web angle takes the shape of a parabolic curve. From Fig. 8(a), it can be seen that a partial tooth contact occurred at the right side of the tooth. The contact side of the tooth is shown in the opposite side of Fig. 4(a). From Fig. 8(b), it is apparent that the partial tooth contact improves when the web is inclined from a straight state to a 15-deg web angle. Clearly, the maximum contact stress is also reduced at the same time. From Fig. 8(c), we find that the partial tooth contact is greatly improved, and it is nearly a uniform distribution of the contact stress when the web angle is changed to 30 deg. The contact pattern of Fig. 8(c) is the pattern that gear designers expect. From Fig. 8(d), it can be seen that the uniform distribution pattern becomes a partial tooth contact pattern again at the left side of the tooth when the web angle is increased to 45 deg"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000731_j.matdes.2007.03.015-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000731_j.matdes.2007.03.015-Figure1-1.png",
        "caption": "Fig. 1. (a) Laterally extruded four teethed gear form and half of the tooth and (b) assumed deformation zones.",
        "texts": [
            " Then a series of fatigue and hardness tests were carried out to tooth of the gear forms to evaluate influence of the forming process. The upper bound analysis is used for modeling of lateral extrusion of the gear forms with tapered teeth. The upper bound method is characterized by its less computation time and memory requirement and reasonable accuracy among the widely known methods. The kinematically admissible velocity fields required for the upper bound calculation must be chosen. The process is divided into 2N zones, each zone is then subdivided into four different regions, as shown in Fig. 1. It is assumed that metal cannot cross or shear along the plane of symmetry. Therefore, material in this unit is displaced to form the half of a tooth. Region 1 does not deform and moves as a rigid body and the velocity field is assumed as, U r1 \u00bc 0; U h1 \u00bc 0; Uz2 \u00bc V 0 \u00f01\u00de Plastic deformation does not take place in the radial direction in region 2 because of the constraint caused by the die wall and axial velocity Uz2 changes with z linearly, thus the velocity field in region 2 in cylindrical coordinates is assumed as U r2 \u00bc 0; U h2 \u00bc V 0 r h \u00f0a h\u00de; Uz2 \u00bc V 0 z h \u00f02\u00de The velocity boundary condition along the OA surface can be checked by substituting h = a \u00f0U h2\u00deh\u00bca \u00bc 0 \u00f03\u00de The velocity field in region 3 is assumed in Cartesian coordinates as U x3 \u00bc V 0 x 2h 1\u00fe 2\u00f0a b\u00de sin 2b ; U y3 \u00bc V 0 y 2h 1 2\u00f0a b\u00de sin 2b ; U z3 \u00bc V 0 z h \u00f04\u00de The resultant normal velocity components along the OB surface in zone 3 must be equal to Uh2 when h = b, can be checked by the following equation: U y3 cos b Ux3 sin b \u00bc r0 h a b\u00f0 \u00de \u00f05\u00de Material in region 4 does not flow in the axial direction and therefore the velocity components are independent of z and velocity in the y direction, Uy3, exists due to straight tapered tooth profile in this region, thus the velocity field in region 4 is assumed as: U x4 \u00bc 1 2h x\u00fe 2x1\u00f0a b\u00de sin 2b ; Uy4 \u00bc y 2h tan c h x\u00fe x1\u00f0a b\u00de sin 2b x3 2 ; U z4 \u00bc 0 \u00f06\u00de The resultant normal velocity vector must be zero along the side surface of the zone 4, which leads to Uy3 Ux3 y\u00bcy\u00f0x\u00de \u00bc tan c \u00f07\u00de The other boundary condition is that velocity in the x direction, Ux3 = Ux4 at x = x1"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000501_s00227-007-0891-x-Figure5-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000501_s00227-007-0891-x-Figure5-1.png",
        "caption": "Fig. 5 A larva in a linear shear flow. The gray circle denotes a larva-particle, and the thick gray arrow indicates the vector of its self-thrust, which rotates due to the larva\u2019s rotation",
        "texts": [
            " Equations of larval motion in two-dimensional linear shear and Poiseuille flows Consider the motion of a larva in the plane O1XZ. For such two-dimensional flow the vector of translational velocity has two non-zero components VX and VY whereas the vector of angular velocity x has only one component xy : x. The fluid velocity vector U is directed along the longitudinal axis and varies linearly with the coordinate Z: U \u00bc 2Ua Z h jX : \u00f029\u00de Here h is the distance above the bottom where the fluid velocity reaches the value 2Ua (Fig. 5). Relation 29 can be interpreted in several ways. It can represent a flow unbounded from above (linear shear flow), a flow where the fluid velocity varies linearly within a layer of depth h and is constant and equal to 2Ua outside the layer (linear boundary layer, Schlichting 1979) and a flow between two plates, where one of them is fixed and the other one moves with velocity 2Ua (Couette flow). In all three cases the characteristic geometric scale of the flow h and the V aU 3 DF V C 3 ( )DF U V A A O BF GF B D p a T p s M SM Substrate Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000301_gamm.200890001-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000301_gamm.200890001-Figure1-1.png",
        "caption": "Fig. 1 Revolute joints: link frames of neighboring links with velocity, force and torque vectors.",
        "texts": [
            " First, link velocities and accelerations are computed from link 1 to link n (outward recursion) applying Newton-Euler formulae to each link: Initial values: \u30080\u3009\u03c90, \u30080\u3009\u03c9\u03070, \u30080\u3009v0, \u30080\u3009v\u03070 prescribed, i : 1 \u2192 n : \u3008i\u3009\u03c9i = \u3008i\u3009R\u3008i\u22121\u3009 \u3008i\u22121\u3009\u03c9i\u22121 + \u03b8\u0307i ez|i (1) \u3008i\u3009\u03c9\u0307i = \u3008i\u3009R\u3008i\u22121\u3009 \u3008i\u22121\u3009\u03c9\u0307i\u22121 + ( \u3008i\u3009R\u3008i\u22121\u3009 \u3008i\u22121\u3009\u03c9i\u22121 ) \u00d7 \u03b8\u0307i ez|i + \u03b8\u0308i ez|i (2) \u3008i\u3009v\u0307i = \u3008i\u3009R\u3008i\u22121\u3009 [ \u3008i\u22121\u3009\u03c9\u0307i\u22121 \u00d7 \u3008i\u22121\u3009 pi +\u3008i\u22121\u3009\u03c9i\u22121 \u00d7 ( \u3008i\u22121\u3009\u03c9i\u22121 \u00d7 \u3008i\u22121\u3009 pi ) +\u3008i\u22121\u3009 v\u0307i\u22121 ] (3) \u3008i\u3009V\u0307i = \u3008i\u3009\u03c9\u0307i \u00d7 \u3008i\u3009 Pi +\u3008i\u3009 \u03c9i \u00d7 ( \u3008i\u3009\u03c9i \u00d7 \u3008i\u3009 Pi ) +\u3008i\u3009 v\u0307i (4) \u3008i\u3009Fi = mi \u3008i\u3009V\u0307i (5) \u3008i\u3009Ni = Ii \u3008i\u3009\u03c9\u0307i +\u3008i\u3009 \u03c9i \u00d7 Ii \u3008i\u3009\u03c9i (6) In a second step, joint forces and torques are computed from link n to link 1 (inward recursion) by Initial values: \u3008n+1\u3009fn+1 = 0, \u3008n+1\u3009n\u0303n+1 = 0, i : n \u2192 1 : \u3008i\u3009fi = \u3008i\u3009R\u3008i+1\u3009 \u3008i+1\u3009fi+1 +\u3008i\u3009 Fi (7) \u3008i\u3009n\u0303i = \u3008i\u3009Ni +\u3008i\u3009 R\u3008i+1\u3009 \u3008i+1\u3009n\u0303i+1 +\u3008i\u3009 Pi \u00d7 \u3008i\u3009 Fi +\u3008i\u3009pi+1 \u00d7 \u3008i\u3009 R\u3008i+1\u3009 \u3008i+1\u3009fi+1 (8) Ti = eT z|i \u3008i\u3009n\u0303i (9) Time t \u2208 [t0, tf ] is the independent variable. Dots indicate time derivatives. \u03b8\u0307iez|i = \u3008i\u3009(0, 0, \u03b8\u0307i) T denotes the rotational velocity at the revolute joint i; ez|i = (0, 0, 1)T is the unit vector in z-direction of frame {i} expressed in its own frame (ex|i, ey|i are defined analogously). vi is the linear velocity and \u03c9i is the angular velocity of frame {i}. pi represents the position of the origin of frame {i} with respect to frame {i \u2212 1} (see Fig. 1). c\u00a9 2008 WILEY-VCH Verlag GmbH & Co. KGaA, Weinheim joint i-1 link i-1 joint i Link i has a total mass of mi and its center of mass (CM i) is moving with the velocity Vi. Ii is the inertia tensor of link i written in a frame, which has it\u2019s origin at CM i and has the same orientation as frame {i}. Pi is the position vector of CM i with respect to frame {i}. \u3008j\u3009fi and \u3008j\u3009n\u0303i denote the force and torque exerted on link i by link i \u2212 1, \u3008j\u3009Fi and \u3008j\u3009Ni represent the force and the torque acting on CM i"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000246_978-3-540-88518-4_120-Figure4-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000246_978-3-540-88518-4_120-Figure4-1.png",
        "caption": "Fig. 4. The schematic diagram of the marsupial robot",
        "texts": [
            " But the marsupial robot can make up for many deficiencies of a single robot in walking, climbing over obstacles, functions, power, energy and communication. Here, the marsupial robot for coal mine rescue adopts a sharing control operation mode based on that the tele-operation is primary and the local independence is secondary. It mainly consists of a mother robot, a baby robot, a remote control center, a remote communication and power system, a sensor system and a lighting system. Its schematic diagram is shown in Fig. 4. The mother robot is an independent and whole unit, with the ability of walking, climbing over obstacles, detecting and measuring. And it can carry food and pharmacy to supply necessary rescue to the wounded. There will be a lot of falling coal blocks after explosion. And according to the analysis of coal quality in China, a coal block doesn\u2019t exceed 500 mm. So to adapt to the utmost unstructured environment composed of all kinds of bodies damaged by accidents, the mother robot must be able to climb over the obstacles of 500 mm, walk on all kinds of roads and debris with different sizes, span over the entrenchments under 600 mm consisting of debris, and pass the alleyway over 600 mm long"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001209_ssp.147-149.542-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001209_ssp.147-149.542-Figure2-1.png",
        "caption": "Fig. 2. The view of parabolic micro-bearing journal surfaces: a) parabolic surface with circumferential grooves, b) cross sections of the parabolic journal with circumferential and longitudinal grooves, c) parabolic journal in magnetic induction field with longitudinal grooves",
        "texts": [
            " In this application, herringbone grooves have the advantage of self-sealing which causes the lubricant to be pumped inward, and therefore, reduces side leakage. They also prevent whirl instability that is observed in the plain journal bearings at concentric operating conditions. Groove location causes the dynamic performances of micro-bearing. Fig. 1 show the system of a HDD and a coupled journal and thrust hydrodynamic bearing used in the HDD spindle [1], [3]. The groove and ridge geometry located on the parabolic surface are presented here. Fig. 2 shows that the grooves on the parabolic journal and sleeve surfaces can be situated in circumferential or longitudinal directions [3]. Random conditions are taking into account. Micro-bearing has application in medical drill bits and hard disc driver HDD spindle medical systems [5]. The HDD spindle samples have a shaft diameter 3.0 mm; rotational speed 20 000 rpm; viscosity of 18 cP (0.018 Pas), radial clearance 3 micrometers, mass 27 g, mass moment of inertia 0.000167 kgm 2 . The width of upper and lower journal bearing changes from 1",
            " Yoon [3], b) classical ridges and grooves on the HDD journal surface, c) height of ridge and groove on the journal and sleeve after Wierzcholski Pressure distributions in parabolic micro-bearings gaps For the parabolic micro-bearing we assume following parabolic co-ordinates: a1=j, a2=yp, a3=p and the non-monotonic generating line of the journal in length direction is taken into account. We have: a  the largest radius of the parabolic journal, a1  the smallest radius of the parabolic journal, 2bp  the bearing length (see Fig. 2). From the system of conservation of momentum and continuity equation [2], [7], [8], [9] after thin boundary layer simplifications and boundary conditions in the parabolic coordinates (j,yp,p) we obtain the dimensional pressure function p(j,p,t) satisfying the modified Reynolds equations in the following form [3]: t )(E h)(EdyAEhdyAE )p(E h h h h dyAE )p(E T2 1T T 0 ps 2 1 T 0 p p3 1 p3 1 T 0 p                      j  w                                        j  j      , (1) where: E  denotes expectancy function, T(j,p,t)  gap height. The equation (1) describes the pressure function p(j,p,t) in parabolic micro-bearing if oil viscosity changes in gap height direction are taken into account. Parabolic micro-bearing are illustrated in Fig. 1 and Fig. 2. The lubricant flow in bearing gap is generated by rotation of a parabolic journal. Bearing sleeve is motionless. Lubricant velocity components v1,v2,v3 in a1=j, a2=yp, a3=p directions, respectively, have the following form[7]:   ,h)A1(A p h 1 t,,y,v 1s 1 pp1 w j  j  (2)   ,A p h 1 t,,y,v p3 pp3    j (3)     ,dy vh hh 1 dy v h 1 t,,y,v pp y 0 p p 31 31 y 0 p 1 1 pp2     j  j (4) and ,dy y ),y,(Ady y ),y,(A, dy 1 dy 1 ),y,(A p T 0 p ppsp py 0 p pp T 0 p py 0 p pps   j  j     j   (5) for =(j,yp,p,B) , t  time, 0  yp  T, 0  j < 2p1, 0  1 < 1, bp  p  bp, Bmagnetic induction. If =(j,p), then we have: As=yp/T and 2A=yp(ypT). For the parabolic shapes of micro-bearing journals we have following coordinates: a1=j, a2=yp, a3=p, and Lame coefficients are as follows [7]:        1p1p1p1p 22 1p1p31p1p 2 1 cossinL/41h,cosah  , (6) , a b L, a aa , a a arccos 1 p 1p 1 1p 1 1p 1p      where a denotes the largest radius of the parabolic shaft, a1  the smallest radius of the parabolic journal, 2bp  the bearing length (see Fig. 2a). Friction forces in parabolic micro-bearing gap This section presents the friction forces calculation in parabolic micro-bearing gaps. The components of friction forces in parabolic j, p directions occurring in micro-bearing gaps have the following forms [7]:  j            j               j F p31 Tpyp 3 pR3R F p31 Tpyp 1 R1R ddhh y v FF,ddhh y v FF , (7) where: =(j,yp,p) \u2013 fluid dynamic viscosity, t  time, 0  yp  T, 0  j < 2p1, 0  1 < 1, bp  p  bp, F  lubrication surface, v1,v3  fluid velocity components in a1=j, a3=p directions, respectively, h1, h3  Lame coefficients (6) in a1=j, a3=p directions"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002595_0022-2569(69)90050-0-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002595_0022-2569(69)90050-0-Figure2-1.png",
        "caption": "Figure 2. Zero-position.",
        "texts": [
            ") If a point of s describes a circle on S and a great circle of s passes permanently through a fixed point on S the motion is called a spherical oscillating slider crank motion. We denote the spherical center and the spherical radius of the circle C on S by M and p~ respectively (Fig. I). Point A of S describes the circle C. The great circle/of s through A and D passes permanently through D. The motion of s fixed to 1 with respect to S is investigated and especially the instantaneous position when the crank A M and I coincides i.e. the crank angle :~=0. This position is called the zero-position (Fig. 2). In this position the instantaneous pole-axis coincides with the Z-axis. The Cartesian co-ordinate systems defined in section 1 are used and are situated as in Fig. 1 and Fig. 2. The diametrical point of d (being the point of s on l coinciding with D) d t also passes through the Z-axis for l is a great circle. Therefore the plane x = 0 of s through l passes permanently through the Z-axis. The line y = 0 , z = 0 is perpendicular to this plane and thus perpendicular to the Z-axis during the motion. The line y = O. z = 0 is a Ball-axis with excess infinity. We consider a motion with the crank-angular velocity d~/dt constant. Moreover we take the magnitude of ~ to be equal to 1 in the zero position"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000488_j.ultras.2008.01.008-Figure10-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000488_j.ultras.2008.01.008-Figure10-1.png",
        "caption": "Fig. 10. Frame sequence for test data. The test sequence consists of evenly spaced frames with an elevational centre offset (along the dashed line) of de \u00bc 0:2 mm and a fixed tilt angle, h, between each adjacent frame pair. We produced sequences with h set to 0.30 and 0.45 .",
        "texts": [
            " For the same reason that the sensor cannot be used to determine the correct offset for each frame pair, it also cannot reliably determine the correct final reconstruction from the two possibilities. This must be determined manually, either from the scanning protocol or from features in the images. In order to demonstrate the improvement in reconstruction accuracy achieved by correcting the decorrelation curve, we recorded several test sequences of frames. The scanning subject was the same phantom used for the decorrelation calibration. Four sequences of 15 frames were recorded with known elevational separation and tilt, as shown in Fig. 10. The elevational offset at the frame centre was 0.2 mm between each frame, and the tilt angle was set to either 0.30 or 0.45 . We also produced one additional test data set and calibration using simulated data (generated using Field II [22]) modelling the 5\u201310 MHz probe used for the real experiments. For the initial experiments, we avoided the distance ambiguity issue by making use of the known frame positions. With the correct distance known for each patch pair, we can determine which side of the decorrelation curve should be used and therefore avoid having multiple distance results"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002335_detc2010-28926-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002335_detc2010-28926-Figure2-1.png",
        "caption": "Figure 2. Example of a determinate truss that is Assur Graph.",
        "texts": [
            "org/ on 08/12/2018 Terms of Use: http://www.asme.org/about-asme/terms-of-use 3 Copyright \u00a9 2010 by ASME between the two graphs in Figure 1. Now, we shall define Assur graphs and outline what distinguishes them from other rigid graphs. Assur Graph \u2013 is a minimally rigid graph with e(G)=2*v(G) where e(G) and v(G) stand for the number of edges and inner vertices of graph G, respectively. The main property of the graph is that removal of any vertex with its incident edges makes the graph non-rigid. The graph, appearing in Figure 2(a) is an Assur Graph since the number of the edges is twice the number of the inner vertices, it is rigid and all its sub-graphs are not rigid. For example, the graph in Figure 2(b) is obtained from the graph Figure 2(a) by deleting vertex C and all its incident edges, resulting in a linkage. The system in Figure 2(c) is obtained by deleting vertex D and is also a linkage. In contrast, the structure in Figure 3(a) is not an Assur Graph since deleting vertex C results in an Assur Graph, known as the Triad, shown in Figure 3(b). A B A B A B a) Assur Graph. b,c) The graphs after deleting vertices C and D, respectively. In each Assur Graph there are two types of vertices: ground vertices, called also pinned vertices, and inner vertices. For example, in a triad type Assur Graph (Figure 3b) there are three inner and ground vertices while in the dyad type Assur graph there are two ground vertices and one inner vertex"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001487_j.rehab.2009.12.001-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001487_j.rehab.2009.12.001-Figure1-1.png",
        "caption": "Fig. 1. A photo of the prototype wheelchair (shown without the braking system).",
        "texts": [
            " The objective of the present article was to evaluate user satisfaction of a new wheelchair prototype with a lever-based manual propulsion technique. Handling ability (related to the prototype-user interface) and other parameters were evaluated with a user self-questionnaire. Data was collected after 2 days of practice on the new prototype. Our prototype is an adjusted non-folding, lightweight Quickie GT1, integrating a system of drive chains and freewheels linked to a two-handed, push bars placed in front of the user. The push-bars are linked to the wheelchair\u2019s rear wheel axles on each side (Fig. 1 and 2). Brake calipers are fitted on the chassis behind the back wheel and can be activated by brake levers fitted onto the push-bars (Fig. 3). None of the other mechanical components of the conventional wheelchair were modified. The prototype was a subject of a publication elsewhere [23]. The objectives are described as follows: to identify the prototype\u2019s advantages and disadvantages in different conditions; help us in improving the prototype\u2019s design; to evaluate the user\u2019s overall satisfaction with the prototype\u2019s push-bar propulsion system",
            " L\u2019e\u0301valuation de la maniabilite\u0301, en relation avec l\u2019interface prototype-utilisateur, est subjectivement e\u0301value\u0301e par un questionnaire de satisfaction des diffe\u0301rents parame\u0300tres. Un recueil des donne\u0301es a e\u0301te\u0301 fait apre\u0300s une utilisation du prototype base\u0301 sur la propulsion manuelle nonconventionnelle. Notre prototype a e\u0301te\u0301 conc\u0327u a\u0300 l\u2019aide d\u2019un syste\u0300me de cha\u0131\u0302nes de bicyclette relie\u0301 aux deux poigne\u0301es de pousse\u0301e. Ces deux poigne\u0301es de pousse\u0301e servent comme des leviers a\u0300 la porte\u0301e de l\u2019utilisateur qui les prend par ses deux mains. Ces leviers ont e\u0301te\u0301 fixe\u0301s et relie\u0301s aux deux axes des roues-arrie\u0300re principales du fauteuil roulant (Fig. 1 et 2). Des e\u0301triers de frein sont monte\u0301s sur le cha\u0302ssis derrie\u0300re la roue arrie\u0300re et active\u0301s par des poigne\u0301es de frein monte\u0301es sur les poigne\u0301es de pousse\u0301e (Fig. 3). Les accessoires et les composantes me\u0301caniques du fauteuil, type Quickie GT1 : ultrale\u0301ger et non-pliable, sont reste\u0301s les me\u0302mes. Notre prototype a e\u0301te\u0301 l\u2019objet d\u2019une e\u0301tude expe\u0301rimentale publie\u0301e [23]. Les objectifs sont de\u0301crits comme suit : identifier les points forts et les points faibles du prototype pour les diffe\u0301rents terrains et dans les diffe\u0301rentes conditions d\u2019utilisation"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000549_tmag.2008.2001316-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000549_tmag.2008.2001316-Figure1-1.png",
        "caption": "Fig. 1. Basic experimental apparatus.",
        "texts": [
            " In this paper, we investigate the loss reduction effect of the divided magnets from both results of the simple experimental apparatus and the synchronous motor, in order to clarify the effects due to the insulation resistance, frequency, and number of magnet divisions. First, the measured results of the simple experiment are compared with the 3-D finite element analysis to verify the validity. In addition, the appropriate modeling of the divided magnet is also investigated. Next, the analysis of a permanent magnet synchronous motor is carried out. It is clarified that the effect of the eddy current loss reduction by the magnet division is weakened in the case of high frequency harmonic Digital Object Identifier 10.1109/TMAG.2008.2001316 Fig. 1 shows the basic experimental apparatus to measure the eddy current loss of the divided magnets. The Nd\u2013Fe\u2013B magnets with a thermal sensor are placed at the center of an exciting coil. Table I shows the specifications of the magnet. After surrounding the magnet by heat insulation material, the coil is excited by ac current whose frequency is 1 to 30 kHz. Then the eddy current loss is estimated by the following expression [2] with the increase of the temperature during 5 min: (1) where is the magnet eddy current loss, is the mass, is the specific heat, and is the temperature",
            "25 mm while the resistivity is modified as 50 , in order to set the is constant. As a result, the number of the finite element mesh is reduced less than 1/100. This mesh is also applied to Model (c), which is a full 3-D model without the magnetomotive force compensation. In this case, the analyzed region is reduced as 1/8 due to the symmetry. Fig. 3 shows the experimental and calculated eddy current losses due to the insulation resistance, when the exciting current is 11.15 A, frequency is 1 kHz, and the magnet is divided into 4 pieces as shown in Fig. 1. The results calculated by model (a)\u2013(c) are almost identical. The validity of the modeling is verified. Although the calculated results slightly underestimate the loss, the tendencies agree well with the experiment. Both results show that the loss increases more when the insulation resistance is less than m, while it is almost constant over this region. The critical insulation resistance is clarified. Fig. 4 shows the calculated eddy current losses in the magnet and insulated layer region. It implies that considerable eddy current loss is generated at the insulated layer, which corresponds to the contact surface of the divided magnet"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003540_14763141.2012.684698-Figure5-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003540_14763141.2012.684698-Figure5-1.png",
        "caption": "Figure 5. Location of control point and COP (a) before adjustment and (b) after adjustment and the separation between adjusted COP and control point (ys).",
        "texts": [
            " Representation of the five paths (1\u20135) that the loaded trolley wheel took over the force plates; the centre of each plate is included (x). Five trials were collected for each path. D ow nl oa de d by [ N ov a So ut he as te rn U ni ve rs ity ] at 0 2: 11 3 1 D ec em be r 20 14 error arising from transverse rotation of the trolley about one wheel, x and y marker position and COP data were combined for further analysis, as follows. Global COP location and position of the control point were adjusted by translating the vectors connecting each point to the global origin so that they lay on the y-axis (Figure 5). Adjusted COP and control point locations were calculated using trigonometric equations. The mean distance between the adjusted control point and adjusted COP location was calculated whilst the trolley was stationary and resting on one plate; the distance measure was used as a reference value with which to compare the dynamic trials. The distance between the adjusted control point location and the adjusted COP location was calculated throughout each trial and compared to the reference value for the section of each trial ^0",
            " Inertia data presented by Dempster (1955) and de Leva (1996) were used for the inverse dynamics analyses as recommended by Hunter et al. (2004). Mean and maximum COP error values calculated from the trolley trials were used to D ow nl oa de d by [ N ov a So ut he as te rn U ni ve rs ity ] at 0 2: 11 3 1 D ec em be r 20 14 determine the effects of COP errors on joint kinetic results during sprint running. The inverse dynamics analysis was repeated for all trials (n: eight athletes) having altered the COP input data by the mean and maximum differences reported between the reference ys value and the mean ys value (Figure 5) as the trolley crossed the plate boundaries. The root mean squared difference (RMSD) between the originally measured and altered joint moment and power time profiles was quantified for each trial. RMSD values were normalised as a percentage of the range of the originally measured values for each variable (%RMSD). The group mean and standard deviation of the %RMSD values were calculated to determine the sensitivity of the COP errors on the joint kinetics determined for dynamic sprint running trials using inverse dynamics analyses"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000165_iros.2007.4399391-Figure4-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000165_iros.2007.4399391-Figure4-1.png",
        "caption": "Fig. 4. Nonmagnetic piezo-electric ultrasonic vibration motor (Shinsei Corporation Inc., Japan).",
        "texts": [
            " The limited purpose of the present study, however, is the feasibility evaluation of synchronous control of MR system and an actuator without electromagnetic interference. Thus, we used an actuator as the controllable object and as a noise source instead of prototyping a surgical manipulator. We adopted a nonmagnetic piezoelectric ultrasonic motor with rotary encoder and its driver unit (USR60-E3N and D6060E, Shinsei Corporation Inc., Japan), which is widely used in MRI-compatible surgical manipulators[5], [7](Fig. 4). The motor was controlled using the above-mentioned digital input/output board (PCI-2726C, Interface Corporation, Japan) for motion command, a DA converter board (PCI-3338) for speed control, and an encoder counter board (PCI-6201) for rotational angle sensing. 4) Control computer: A computer (CPU: Pentium III, 500 MHz, RAM: 384 MB) was implemented to integrate above subsystems. The operating system was originally built using Red Hat Linux 9 (kernel version 2.4). The task of the computer was to manage subsystems"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001247_0954406jmes321-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001247_0954406jmes321-Figure2-1.png",
        "caption": "Fig. 2 Generation of a helicoid",
        "texts": [
            " The constant \u03b8os determines half the space width on the base cylinder, and the transverse tooth profile of a standard involute helical gear is represented by equations (1) to (3) [10] \u03b8os = \u03c0 2Ns \u2212 inv\u03b1s inv\u03b1s = tan \u03b1s \u2212 \u03b1s (1) \u03b1s = tan\u22121 ( tan \u03b1n cos \u03b2 ) (2) where \u03b2 is the helical angle, and parameters \u03b1s and \u03b1n are the transverse pressure angle andthe normal pressure angle, respectively. xs = \u00b1rbs[sin(\u03b8ks + \u03b8os) \u2212 \u03b8ks cos(\u03b8ks + \u03b8os)] ys = \u2212rbs[cos(\u03b8ks + \u03b8os) + \u03b8ks sin(\u03b8ks + \u03b8os)] zs = 0 (3) The upper and lower signs in equation (3) correspond to the right- and left-side profiles, respectively, while rbs is the radius of the base circle. The tooth surface of the helical gear can be obtained by transferring the transverse plane curve with a screw motion as shown in Fig. 2 and equation (4) [10] R1(\u03b8ks, \u03c8) = M1s(\u03c8)Rs(\u03b8ks) (4) where R1 = \u23a7\u23aa\u23aa\u23aa\u23a8 \u23aa\u23aa\u23aa\u23a9 x1 y1 z1 1 \u23ab\u23aa\u23aa\u23aa\u23ac \u23aa\u23aa\u23aa\u23ad M1s = \u23a1 \u23a2\u23a2\u23a2\u23a3 cos \u03c8 \u2212 sin \u03c8 0 0 sin \u03c8 cos \u03c8 0 0 0 0 1 p\u03c8 0 0 0 1 \u23a4 \u23a5\u23a5\u23a5\u23a6 Rs = \u23a7\u23aa\u23aa\u23aa\u23a8 \u23aa\u23aa\u23aa\u23a9 xs ys zs 1 \u23ab\u23aa\u23aa\u23aa\u23ac \u23aa\u23aa\u23aa\u23ad (5) The screw rotation angle is denoted as \u03c8 , while parameter p is the parameter of the screw motion [10] p = rps cot \u03b2 (6) where rps is the radius of the pitch circle. The surfaces of the helical gear in coordinate system S1 are shown in equation (7), in which the upper and lower signs correspond to right- and left-side profiles, Proc"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001069_1.3070580-Figure4-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001069_1.3070580-Figure4-1.png",
        "caption": "Fig. 4 Comparison of friction variable with illustrations in Szeri \u202015\u2021",
        "texts": [],
        "surrounding_texts": [
            "T f e\nA l m e\nT s B b n r t p\n3\nc\nC c t a\nF fl\n0\nDownloaded Fr\n1\n\u0304k\nh\u0304 3 p\u0304 + 1 \u0304kz R2 L2\nz\u0304 h\u0304 3 p\u0304 z\u0304 = 1 2 h\u0304 2\nhe lubricant film profile is determined by superimposing the efect of normalized tilt and radial adjustments on the conventional xpression for the film thickness 10 ,\nh\u0304 = 1 + cos \u2212 + Radj + Tadj 3\ns suggested by Taylor and Dowson 11 , the linearized turbuence model of Ng and Pan 12 is used to analyze the perforance in the turbulent regime. The turbulent coefficients are valuated from\nk = 12 + kx ReL nx 4\nkz = 12 + kzz ReL nz 5\nhe parameters kx, nx, kzz, and nz are taken from Taylor and Dowon 11 . A simplified adiabatic model proposed by Pinkus and upara 13 is employed to study the effect of temperature on the earing performance. This adiabatic model 13 assumes an expoential relationship between viscosity and temperature and further elates the temperature within the film at any angular position to he film thickness in the circumferential direction. The required arameters are taken from Martin 6 ,\n\u0304 = e\u2212 T\u2212Ti 6a\nT \u2212 Ti = 1 ln 1 + E\n1\n2 d\nh\u0304 2 6b\nPressure Boundary Condition Pressure along the bearing edges was set to zero. Mathemati-\nally it is given as\np\u0304 = pa = 0 at z\u0304 = 0 and z\u0304 = 1.0\np\u0304 = pa = 0 at = 1 and = 3 7\navitation is allowed to occur at ambient pressure by setting all alculated negative pressures equal to zero throughout the iteraive solution scheme. This implies that the lubricant film ruptures nd reforms when\np\u0304 = p\u0304 = 0 8\nig. 1 Schematic diagram of single pad externally adjustable uid film bearing\n21701-2 / Vol. 131, APRIL 2009\nom: http://tribology.asmedigitalcollection.asme.org/ on 01/29/2016 Terms\n4 Steady State Performance Characteristics The load carrying capacity along the line of centers and in its perpendicular directions is evaluated in the nondimensional form,\nW\u0304r = \u2212 0 1 1 2 p\u0304 cos d dz\u0304 9a\nW\u0304t = 0 1 1 2 p\u0304 sin d dz\u0304 9b\nThe resultant force acting on the journal is\nW\u0304 = Wr 2 + Wt 2 10\n0 = tan\u22121 Wt\nWr 11\n5 Oil Flow\nThe total volume flow rate Q\u0304z is evaluated as\nQ\u0304z = Q\u03041 + Q\u03042 = 1\n2 h\u03043\nkz\u0304 p\u0304 z\u0304 c d 12\nThe slope of pressure curves at the sides of the bearing is given as, p\u0304 / z\u0304 c,\nQ\u03041 = \u2212 1\n2 h\u0304 3\nkz\u0304 p\u0304 z\u0304 z\u0304=0 d 13a\nQ\u03042 = \u2212 1\n2 h\u0304 3\nkz\u0304 p\u0304 z\u0304 z\u0304=1 d 13b\n6 Friction Parameter Friction force, as per Pinkus and Sternlicht 14 , is given as\nF = dA\nFor the dominant Couette flow, Taylor and Dowson 11 gave the shearing stress acting on the surfaces as\n= \u2212 U\nh \u0304c\nh\n2\np x 14\nwhere\nTransactions of the ASME\nof Use: http://www.asme.org/about-asme/terms-of-use",
            "7 c a e o\nt\nF \u2020\nJ\nDownloaded Fr\n\u0304c = 1 + 0.00232 ReL 0.86 for ReL 10,000 15a\n\u0304c = 1 + 0.00099 ReL 0.96 for ReL 10,000 15b\nF\u0304 = 0 1 1 2 h\u0304 2 p\u0304 d dz\u0304 + 0 1 1 2 \u0304 h\u0304 \u0304cd dz\u0304 16\nF\u0304 W\u0304 = f R C 17\nComputational Technique A computer program is developed to determine the performance haracteristics such as load capacity, attitude angle, friction varible, lubricant flow rate, and Sommerfeld number. The governing quation is solved using the Gauss\u2013Seidel method with successive ver-relaxation scheme.\nLoad carrying capacity and attitude angle are calculated from he pressure distribution. In a conventional bearing analysis, the\nig. 3 Comparison of attitude angle with illustrations in Szeri 15\u2021\n\u00af\nournal of Tribology\nom: http://tribology.asmedigitalcollection.asme.org/ on 01/29/2016 Terms\ncoordinate i.e., circumferential direction , is taken from the position of maximum film thickness. The present analysis requires that the position of maximum film thickness be found beforehand. This is done by assuming an arbitrary value of attitude angle . After each calculation, the attitude angle computed from Eq. 11 is compared with the assumed value of the attitude angle . The value of is modified with a small increment, and Eq. 2 is solved using this modified value until is equal to 0. Figure 2 shows the schematic diagram of various angles and adjustments considered in the present analysis.\n8 Validation Externally adjustable fluid film bearing will perform as a conventional partial arc bearing when Radj and Tadj are set to zero. Steady state performance curves in Figs. 3 and 4 compare the results of the present analysis with the results illustrated in Szeri 15 for a centrally loaded 160 deg fixed pad partial bearing with L /D=1.0, operating with a Reynolds number 5000 and 10,000. This comparison validates the implementation of the turbulence model.\nearing midplane with \u03b5=0.4\nAPRIL 2009, Vol. 131 / 021701-3\nof Use: http://www.asme.org/about-asme/terms-of-use",
            "F u\nRadj for E=0.0\nand Radj for E=0.0\n021701-4 / Vol. 131, APRIL 2009\nDownloaded From: http://tribology.asmedigitalcollection.asme.org/ on 01/29/2016 Terms\n9 Results and Discussion\nFigures 5\u201314 illustrate the results of the present analysis for a bearing having Radj of 0.25C and of 0.0061 deg. Figure 5 shows that for adiabatic parameter E=0.0, the maximum value of the pressure p\u0304max developed in the fluid film with negative Tadj and negative Radj configuration is nearly 18 times higher when compared with positive adjustments.\nMoreover, p\u0304max for zero adjustments is about three to four times higher than that for positive adjustments. This rise in p\u0304max is obvious because negative adjustment configurations result in a narrow convergent zone in the film. However, an increase in parameter E reduces p\u0304max.\nFigure 6 depicts that in laminar as well as turbulent flow conditions, the load carrying capacity is higher with negative Tadj and negative Radj. Reynolds number Re has a positive influence on load carrying capacity. However, at lower values of eccentricity ratio, negative Radj and negative Tadj with a laminar flow condition have lesser load carrying capacity than zero Radj and zero Tadj with Re=16,000. Similarly, zero Radj and zero Tadj with a laminar flow condition have lesser load carrying capacity than positive Tadj and\ny ratio for various values of Tadj and\nity ratio with various values of Tadj\nig. 7 Load capacity versus eccentricity ratio for various vales of Re and E\nFig. 8 Attitude angle versus eccentricit\nTransactions of the ASME\nof Use: http://www.asme.org/about-asme/terms-of-use"
        ]
    },
    {
        "image_filename": "designv11_12_0001827_2010-01-1198-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001827_2010-01-1198-Figure1-1.png",
        "caption": "Figure 1. Typical valve train configurations: (a) type I - bucket tappet, (b) type II - finger follower, (c) type III - rocker arm, (d) type IV, (d) type V - pushrod.",
        "texts": [
            " First, the cam design and optimization methodologies are discussed. Then, the benefit and limitation of current valve train dynamic simulation methods are presented. The mathematical modeling considerations involving the valve spring, hydraulic lash adjuster, oil aeration, bulk modulus, and dry contact are detailed. The characteristics of the elastohydrodynamic cam to follower contact and implications of various theories on the accuracy of the lubricant film predictions are pointed out. Finally, friction and wear of main contact interfaces are presented. Figure 1 shows typical conventional valve train configurations. A discussion of the main advantages and disadvantages of each configuration is found in [12]. Variations of these standard configurations can be found in production engines (bridge, dual lifter, etc). A typical kinematic model of the valve train uses lumped masses which move under geometrical constraints. A dynamic model includes elasticity of the bodies and component separation. In both cases, motion is represented by an imposed angular speed at cam",
            " On the other hand, an aggressive cam profile leads to high accelerations, high forces, and undesirable dynamic behavior of the valve train. In addition, the cam design must satisfy manufacturing limitations with respect to maximum concave radius and durability requirements. Therefore, the cam shape generation is an integrated process which requires simultaneous evaluation of valve train performances rather than a standalone process of curve generation. Design considerations and performance parameters used in the cam design process are summarized in Table 1. <figure 1 here> For a given valve train layout and thermodynamically optimized valve lift duration, the cam shape and, therefore, effective valve lift is normally restricted by the mechanical requirements and manufacturing capabilities. Ideally, the effective valve lift will match perfectly the thermodynamic optimized curve. However, the effective valve lift can differ from the thermodynamic curve without a significant loss of gas exchange efficiency depending on the valve port flow coefficients. The flow coefficient can increase significantly in the first section of the valve event as shown in Figure 2 after that being almost constant"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001135_s11661-009-9870-9-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001135_s11661-009-9870-9-Figure2-1.png",
        "caption": "Fig. 2\u2014Schematic of LPD process, leading half and trailing half are shown in the molten region.",
        "texts": [
            " In order to reduce the order of nonlinearity of the problem, the local surface heat transfer at the molten pool is estimated according to Vinokurov:[20] q \u00bc h\u00fe re T3 \u00fe T0T 2 \u00fe T2 0T\u00fe T3 0 T T0\u00f0 \u00de \u00bc Hlump T T0\u00f0 \u00de \u00bd7 where Hlump can be estimated as Hlump \u00bc 2:4 10 3eT1:61 m \u00bd8 The associated loss in accuracy using this relationship is estimated to be less than 5 pct.[2] The emissivity coefficient around the melting temperature is considered to be 0.6.[16] 3. Assumptions The assumptions during the thermal modeling of the LPD process are as follows. (1) The substrate is initially at room temperature (25 C). The boundary condition around the substrate is the convection heat transfer with a constant coefficient. (2) The heat flux on the leading half (Figure 2) has uniform distribution, based on Eq. [5]. (3) The latent heat is considered in the temperaturedependent definition of specific heat. (4) The activated elements of the molten pool are at the melting temperature. The convective redistribution of heat in the molten pool is ignored. Based on the results from the thermal model, the temperature history of the nodes is used to predict the phase transformations during the heating and cooling cycles. The microstructure of the material after solidification consists of austenite",
            " This effect, however, for Ms and Mf is negligible, and the assumed values for the martensite start and end temperatures are 350 C and 150 C, respectively. The mentioned coupled models of thermal and thermokinetic analysis are used to investigate the effect of path planning on the microstructure of the deposited material in the LPD process. For this purpose, four cases with the same deposition area are studied for different deposition patterns, as shown in Figure 1. In the LPD process, as shown in Figure 2, the deposition region can be divided into two regions: the \u2018\u2018leading half,\u2019\u2019 which moves in front of the laser beam and melts the substrate; and the \u2018\u2018trailing half,\u2019\u2019 which follows the leading half and contains the molten material. This concept is used for modeling the deposition process by the finite element method, as shown in Figure 3. In this figure, four regions are distinguished: the white elements represent the substrate or the layer underneath, the light-gray elements represent the deposited elements that were activated in the previous timesteps, the dark-gray elements are the activated elements Table II"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001369_j.jsv.2008.09.045-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001369_j.jsv.2008.09.045-Figure1-1.png",
        "caption": "Fig. 1. Chain CVT configuration and chain link structure.",
        "texts": [
            " In spite of the several advantages proposed by a CVT system, its complete potential, in terms of the mass-production and market penetration of CVT-equipped vehicles, has not been realized so far. In order to achieve lower emissions and better performance, it is necessary to understand the various complex dynamic interactions occurring in a CVT system in detail so that efficient controllers could be designed to overcome the existing losses and enhance vehicle fuel economy. A chain CVT consists of two variable diameter pulleys kept at a fixed distance apart and connected to each other by a chain. Fig. 1 depicts the chain structure and the basic configuration of a chain CVT drive. A rockerpin chain consists of plates and rocker pins, as depicted in the figure. All plates and pins transmit tractive power. One of the sheaves on each pulley is movable. So, the chain is capable of exhibiting both radial and tangential motions depending on the torque loading conditions and the axial forces applied to the pulley sheaves. Moreover, the contact forces between the rocker pins and the pulleys are discretely distributed",
            " It is important to note that although an exact knowledge of the friction characteristic in a CVT system can only be obtained by conducting experiments on a real production CVT, these mathematical models give profound insight into the probable behavior of a CVT under different operating conditions. This knowledge could be further exploited to design more efficient controllers and identify various loss mechanisms in a CVT system. The modeling analysis and the results corresponding to the chain CVT model are discussed in detail in the subsequent sections. As depicted in Fig. 1, a chain CVT consists of two variable diameter pulleys connected to each other by a chain. The system is subjected to an input torque on the driver pulley and a resisting load torque on the driven pulley. The model development and analysis includes the following assumptions: (1) The pulleys do not have any misalignment between them. (2) The chain links are rigid. (3) The interactions between the rocker pins of neighboring links and between a rocker pin and a plate can all be accounted for by modeling the chain link as a planar rigid body"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002043_pime_proc_1967_182_025_02-Figure6-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002043_pime_proc_1967_182_025_02-Figure6-1.png",
        "caption": "Fig. 6. Synchronous response of rotor bearing system",
        "texts": [
            " As the running speed rises, the separation between the synchronous oil film impedances reduces, that is higher critical speed pairs are nearer to each other in frequency than lower pairs. Furthermore, the separation increases whh oil viscosity as one might expect since increasing viscosity and raising speed both operate in the same manner on the journal attitude, that is they reduce the eccentricity ratio. A typical plot of rotor response in the vicinity of the first critical speeds is shown in Fig. 6. Q factors deduced from similar plots at four shaft locations for three grades of oil Proc Instn Mech Engn 1967-68 Sync n r on a u s 0 16 3 x !C!-?reyns a; 76\u00b0F beu rili g i in p e d a n ces A 75.5 x 10-7reyns a t 76\u00b0F --- Zero bear ing darn p i n q i I T ~ p e d c ri c e si 1 - 7 r x i 84 x 10-7revni a t 76\u00b0F O 2 0-1 I0 L-J--- h-&=+Jk=k.- i L ~ L - d show an increase as oil viscesity decreases (that is as eccentricity ratio increases). This result is confirmed by the plots of bearing damping against eccentricity ratio for circular bearings, presented in reference (6)",
            " It appears difficult to extrapolate the results to the analysis of a rotor which is non-symmetric about its centre of span. A few comments and questions appear to be pertinent to the \u2018Experimental work\u2019 section of the paper. Approximate calculations indicate the bearings of the rig may be operating at a Sommerfeld number of the order of 1 to 4, which indicates light loading. One wonders how the apparently poor correlation between theory and experiment as shown in Fig. 4b affects the validity of the author\u2019s conclusions. While Fig. 6 is an interesting presentation of response data, might not more traditional plots of calculated journal vibrational amplitude versus shaft speed, both with and without the bearing cross-coupling coefficients in the analysis, provide a clearer and more immediate answer to the question at hand ? The ultimate result of the analysis is, after all, a prediction of how badly the system will vibrate under given conditions. In addition, a sample response calculation showing the magnitudes and indicating the source of the bearing coefficients would be very useful to the design engineer who will be evaluating this information and applying it to his own particular problem",
            " With regard to the dynamic characteristics of the bearings, it would have been of added value if the author had obtained rough estimates of the displacement coefficients from the static locus and load curves (Fig. 4a and b). These and their associated velocity coefficients could then have been utilized alongside his other coefficients to assess the relative effects on whirl onset (Fig. 7). In this connection a little explanation of the derivation of \u2018equivalent\u2019 bearing impedances would have been useful. The author concludes that the relative axial slope of journal to bearing accounts for the fact that experimental amplitudes (Fig. 6) are much less than theoretically predicted; presumably oil-film non-linearity effects and stiffness of the drive could have had a similar effect, especially at high vibration amplitudes. It would also be interesting to hear what criteria the author adopted to define the experimental critical speed and whether he would expect its insensitivity to oil-film stiffness to be associated with a sensitivity of whirl-onset speed to film stiffness. One might expect that if a rotor oil-whirled with large amplitude it would exhibit a hysteresis effect on running down due to modification of the oil-film pressure in the whirling condition from that in the initial stable condition"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000880_tmag.2007.891399-Figure6-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000880_tmag.2007.891399-Figure6-1.png",
        "caption": "Fig. 6. Distributions of Lorenz force density vectors: (a) t = 0:125 ms; (b) t = 4:25 ms.",
        "texts": [
            " Therfore, the initial potentials of all edge elements are set to be zero in order to simulate under the conditions of that the ring magnet is rotated at constant speed and suddenly inserted into the gap between the conductors, and the ring magnet is rotated from the rotation angle of 45 . Figs. 5 and 6 show the distributions of eddy current and Lorentz force density vectors, respectively, when the ring magnet is rotated at the rotation speed of 2000 rpm. From Fig. 5(a) and (b), the distributions of eddy current density vectors are amazingly different because the flux density vectors are suddenly occurred at the initial position. For that reason, the distributions of Lorentz force density vectors are also different as shown in Fig. 6(a) and (b). Fig. 7 shows the calculated time variations of transient torque when the ring magnet is rotated at the rotation speed of 2000 rpm. The peak value of torque is about 0.8 at the position where the boundary of ring magnet passes through the center of conductors. From above Figs. 5 and 6, the eddy current and Lorentz force density vectors of initial position are larger than that of . However, it is found that the torque of initial position is nearly zero because of the directions of both vectors"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001680_biorob.2010.5626817-Figure13-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001680_biorob.2010.5626817-Figure13-1.png",
        "caption": "Fig. 13. The initial and final configurations of an intentional goal walking (a) and the animated locomotion generated by our controller (b)",
        "texts": [
            " The person A goes first along the main corridor, then avoiding the other person B by sideward steps and finally, continue to go ahead. The other walker B goes from the left side of the main corridor on a straight line and turn on his right when arriving in the main corridor. He will meet the person A and will stop to let A pass first, then he will continue. B. Intentional locomotions An intentional trajectory from initial to final configurations is produced by using the inverse optimal control approach in [11]. Such a locomotion is presented in figure 13 (a) and it is close to locomotions acquired from data of human locomotion experiment in the neighboring space [15]. Like in [15], the mannequin adapts the body orientation to the goal during locomotion. We chose desired velocities and then generated the locomotion. A postures sequence of the result during the animation was shown in figure 13 (b). Nonholonomic and holonomic locomotions have been combined in our locomotion controller. We produced some feasible locomotions (look like human walking). The ongoing work allows us to animate locomotions following predefined paths in some interesting scenarios. Moreover, with the optimizer used in [11], an optimal (human-like) trajectory would be planned and tracked by locomotions generated by our controller. From a top-down approach, we defined a computational model explaining trajectory shape of human locomotion between different configurations"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003713_978-1-4614-2419-2-Figure51.1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003713_978-1-4614-2419-2-Figure51.1-1.png",
        "caption": "Fig. 51.1 Re-design of the handlebar (in the red ellipse the strain gage load cell) (color figure online)",
        "texts": [
            " Thus, handlebar, footrest and seat were re-designed to accommodate load cells. Moreover, due to the fact that on existing rowing machine the inclination of the footrest cannot be modified and one of the main targets of the research was to assess the influence of such inclination on the athlete\u2019s performance, it was decided to re-design the footrest to allow inclination changes. Since these changes can easily be done with the footrest that is mounted on rowing boats, it was decided to adopt the same footrest as on boats. Figure 51.1 shows the re-designed handlebar with the interposed load cell. Since the forces exerted on the handlebar are mainly axial, i.e. are mainly directed as the chain that connects the handlebar to the fan system, an unidirectional strain gauge load cell was adopted. The measuring range is 1,000 N and its dimensions (external diameter equal to 54.5 mm) allow to fit the load cell directly into the handlebar. The only drawback, as stated by the tester, is the mass of the load cell (250 g) and of its aluminium assembly (350 g)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000881_jmes_jour_1973_015_005_02-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000881_jmes_jour_1973_015_005_02-Figure1-1.png",
        "caption": "Fig. 1. How hydrodynamic firm pressure tends to reduce relative face nutation",
        "texts": [
            " INY Before developing the theory of seal face lubrication it would be instructive to show that whilst a load-carrying film can be generated purely by a relative nutation between the rotating and stationary faces (Nahavandi and Osterle (a), Whiteman (7)), a wavy surface is necessary for stable operation with hydrodynamic lubrication. This may be shown to be true by considering the dynamics of the rotor-stator combination, The relative nutation of flat rotating faces generates in the fluid film between the faces a pressure wave whose resultant thrust combines with the axial load to impose on both the rotating and stationary components a couple which, if either component is free to move, tends to reduce the relative nutation and hence to reduce the film pressure (see Fig. 1). The single pressure wave must provide both the axial force to balance the applied face loading and the tilting moment required to move the components in their mounting; these forces are therefore interdependent and the load carrying capacity is determined by the stiffness of the mountings. If one of the components is flexibly mounted, as in all mechanical seals, the load capacity becomes very small and wear is inevitable except under the lightest loads. If one of the faces has a wavy surface comprising at least two approximately symmetrically spaced humps the axial force and tilting moment are no longer interdependent"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001362_6.2008-7254-Figure10-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001362_6.2008-7254-Figure10-1.png",
        "caption": "Figure 10. Lateral/directional geometry for pursuit guidance",
        "texts": [
            " 1sin A T long i h h r \u03bb \u2212 \u239b \u239e\u2212\u2032 = \u239c \u239f \u239d \u23a0 (23) where Ah is an altitude of the aircraft, and Th is a height from the ground to the center of the recovery net. Note that the radius ir determining the imaginary landing points is constant in Fig. 9. This radius can be changed adaptively proportional to the remaining range to improve the performance of the terminal guidance. In the lateral/directional guidance law, pure pursuit guidance commands to aim the target directly to make the error between a heading angle and a lateral line-of-sight angle zero. Figure 10 shows the aircraft and recovery net in the X - Y plane. In Fig. 10, \u03c8 is a heading angle of the aircraft, and lat\u03bb is a lateral line-of-sight angle provided by the image processing module. American Institute of Aeronautics and Astronautics 11 In the same way as the longitudinal landing guidance, a desired heading angle d\u03c8 is defined to make the error between the lateral line-of-sight angle lat\u03bb and heading angle \u03c8 zero. Lead angle scheme is also applied to steer ahead and align perpendicular to the recovery net. ( )d latK\u03c8\u03c8 \u03bb \u03c8= \u2212& (24) [ ]d lat leadK\u03c8\u03c8 \u03bb \u03c8 \u03c3= \u2212 +& (25) where lead\u03c3 is a lead angle, and d\u03bb is a desired line-of-sight angle at a point of impact"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000224_1-4020-2933-0_13-Figure20-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000224_1-4020-2933-0_13-Figure20-1.png",
        "caption": "Figure 20. An example for constructing the velocity transformation matrix.",
        "texts": [
            " A useful definition is vector jid , , shown in Figure 19, connecting the origin of body i to a point on joint j. This vector is computed as: P jiji, rrd \u2212= . (46) By deriving the velocity relationships, it can be shown that there is a transformation between body velocities and joint velocities as: Bv = . (47) Matrix B is called the velocity transformation matrix and it is orthogonal to the system Jacobian: 0DB = . (48) The structure of matrix B can be demonstrated through a simple example. Consider the open-chain single-branch system shown in Figure 20(a) containing 5 bodies, 1 floating joint, 2 revolute joints, 1 prismatic joint, and 1 spherical joint. The velocity transformation equation for this system, in symbolic form, is found to be: \u23aa \u23aa \u23aa \u23ad \u23aa\u23aa \u23aa \u23ac \u23ab \u23aa \u23aa \u23aa \u23a9 \u23aa\u23aa \u23aa \u23a8 \u23a7 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a6 \u23a4 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a3 \u23a1 = \u23aa \u23aa \u23aa \u23ad \u23aa\u23aa \u23aa \u23ac \u23ab \u23aa \u23aa \u23aa \u23a9 \u23aa\u23aa \u23aa \u23a8 \u23a7 5 4,3 3 2 1 5,55,45,35,25,1 4,44,34,24,1 3,33,23,1 2,22,1 1,1 5 4 3 2 1 v RSPRF 0SPRF 00PRF 000RF 0000F v v v v v . The elements in the velocity transformation matrix are called block matrices. Block matrices for four fundamental joints are described in Table 3. The description of the di j , vectors that appear in the block-matrices for our example can be found in Figure 20b). We note that d1,1 = 0. The main conclusion we draw from this example is that the velocity transformation matrix for any open-chain system can be constructed systematically from the topology of the tree structure and the block-matrices. The time derivative of Eq. (47) yields the acceleration transformation equation as: BBv += . (49) Substituting Eqs. (47) and (49) in the constraints of Eqs. (16) and (17), and then using Eq. (48) shows that all of the constraints will disappear when we transform the constraints to the joint coordinates"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002821_ichr.2010.5686331-Figure3-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002821_ichr.2010.5686331-Figure3-1.png",
        "caption": "Fig. 3 Movement of passive element during discrete change of object behavior",
        "texts": [
            " Although the above-mentioned related works have needed and developed several kinds of sensors for these measurements, it is difficult to implement all these sensors in the robotic hand. In this study, instead, a combination of a mechanical passive element and its deflection sensor was used for measuring object characteristics. A special feature of the combination of a mechanical passive element and its deflection sensor is that it can measure not only the applied force but also the position of the output link of the mechanical passive element. It is assumed the robot finger pushes the object constrained only by the friction force at the contact with the table (Fig. 3). When the robot finger grows gradually applies higher force, the object will not move (because of the constraint due to the friction force) while the applied force is lower than the maximum friction force. At the moment that the applied force equals the maximum friction force, the object slips discretely, because the friction force changes from static to dynamical. The feature of the sensing method proposed here is that it identifies this discrete change of object behavior because the discrete change appears clearly in the position information concerning the object"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000698_j.ijmecsci.2008.02.003-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000698_j.ijmecsci.2008.02.003-Figure2-1.png",
        "caption": "Fig. 2. Cross-section of a fluid film journal bearing.",
        "texts": [
            " Chen / International Journal of Mechanical Sciences 50 (2008) 1090\u20131113 1095 C.-W. Chang-Jian, C.-K. Chen / International Journal of Mechanical Sciences 50 (2008) 1090\u201311131096 Rx \u00bc \u00f0e d\u00dekc e \u00bd\u00f0f \u00fe bv\u00deX \u00fe Y (12) 2.3. Dynamic equations Fig. 1 shows a flexible rotor supported horizontally by two identical and aligned turbulent journal bearings with non-linear springs. Om is the center of rotor gravity, O1 is the geometric center of the bearing, O2 is the geometric center of the rotor, O3 is the geometric center of the journal. Fig. 2 shows the cross-section of the fluid film journal bearing, where (X,Y) is the fixed coordinate and (e, j) is the rotated coordinate, e being the offset of the journal center and j being the attitude angle of the X-coordinate. From the equilibrium of force, the forces applied to the journal center O3 are Fx \u00bc f e cos j\u00fe f j sin j \u00bc ks\u00f0X 2 X 3\u00de=2 (13) Fy \u00bc f e sin j f j cos j \u00bc ks\u00f0Y 2 Y 3\u00de=2 (14) where fe and fj are the resulting viscous damping forces in the radial and tangential directions. The equations of motion of O2 in the Cartesian coordinates and the equations of motion of the bearing C"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003424_s10035-011-0272-5-Figure3-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003424_s10035-011-0272-5-Figure3-1.png",
        "caption": "Fig. 3 Definition of contact angle \u03b8 , the deviation of contact direction from the force chain direction",
        "texts": [
            " Only those particles experiencing average stress greater than a threshold value, as indicated by average image intensity per particle, were retained to be considered as parts of force chains. A particle was defined as being part of a chain if it continued in approximately the same direction as the previous part of the chain, as defined by an angle \u03b8 between the chain propagation direction and a line drawn between the center of the previous particle assigned to the chain and the center of the current particle (Fig. 3). If \u03b8 is greater than the threshold angle \u03b80 the chain terminates, the length is recorded, and the process moves to the beginning of the next chain. If \u03b8 is less than the threshold angle the particle is considered to be a member of the chain and the next particles in contact are identified and tested in the same way to see if the chain continues. The dependence of the force chain lengths was studied thoroughly as a function of the intensity and angle thresholds. Chain length distributions were calculated for all 38 data sets as a function of intensity thresholds in increments of one percent of maximum intensity"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003713_978-1-4614-2419-2-Figure31.9-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003713_978-1-4614-2419-2-Figure31.9-1.png",
        "caption": "Fig. 31.9 Laser tachometer installation orientation variables",
        "texts": [
            " the effect of sensor orientation on the sensor\u2019s output pulse train, the raw output signal of the 140 pulse per revolution laser tachometers were examined on an oscilloscope while their installation orientations were adjusted slightly. The results of this preliminary analysis suggested that the laser tachometer\u2019s output signal quality was strongly affected by all three installation orientation degrees of freedom. Changes in the laser tachometer\u2019s installation skew angle, offset distance, and orientation angle relative to the encoder hub are illustrated in Fig. 31.9. A formal study into the effects of laser tachometer installation orientation suggested that the sensor\u2019s output signal quality and pulse train duty cycle were dependent on all three installation degrees off freedom. The sensor\u2019s pulse train quality was optimized with the following orientation: (1) a 40\u2013200 mm offset distance; (2) the laser tachometer\u2019s emitter and receiver orthogonal to encoder stripes 45 , and; (3) the axis of the sensor passing through the shaft\u2019s axis and normal to the encoder hub\u2019s surface"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002151_s11434-009-0089-3-Figure5-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002151_s11434-009-0089-3-Figure5-1.png",
        "caption": "Figure 5 Assembling process of the biomimetic undulating fin. (a) Detailed connection of the fin ray to the motor; (b) three-step assembling procedure.",
        "texts": [
            " The undulation kinematics model reveals that propulsive waveforms are related with the inclined angle of fin rays, that is, it can generate a symmetric propulsive wave with \u03b2 = 90\u00b0 and an asymmetric mode with \u03b2<90\u00b0. Thus, the fin-ray inclined angle, \u03b2, should be regulable in the undulating fin. Given that the fin-ray inclined angle can hardly be regulated automatically in the preexisting prototype, several sets of fin rays with different inclined angles are provided for experiments. As shown in Figure 5, both symmetric and asymmetric waveforms are kept in the same mechanical structure, whereas the assembling process is a little different by only replacing upright fin rays with inclined fin rays. The modular mode enables us to easily shift one specified angle to another. As for the present mechanism, the space between fin rays is limited by the geometrical dimension of the single joint driven mechanism. Two sets of fin rays are designed and used in this experiment for bionic asymmetry 566 HU TianJiang et al"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003501_2041302510394742-Figure11-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003501_2041302510394742-Figure11-1.png",
        "caption": "Fig. 11 Longitudinal model of rotor-fixed wing hybrid aircraft",
        "texts": [
            " The authors select f3 = 1 Jzbh (kp1\u03c8 + kp2\u03c8\u0307) (27) where kp1 and kp2 are positive constants. The system is exponentially stable. 4.3 Control of longitudinal dynamics The result of controlling the pitch angle \u03b8 is the longitudinal dynamics (\u03c6 = 0, \u03c8 = 0), which is described by mX\u0308 = (Tf + Tt \u2212 f1)c\u03b8 + f2s\u03b8 + L1(c\u03b8 s\u03b1 + s\u03b8 c\u03b1) + D1(\u2212c\u03b8 c\u03b1 + s\u03b8 s\u03b1) mZ\u0308 = \u2212(Tf + Tt \u2212 f1)s\u03b8 + f2c\u03b8 \u2212 mg + L1(\u2212s\u03b8 s\u03b1 + c\u03b8 c\u03b1) + D1(s\u03b8 c\u03b1 + c\u03b8 s\u03b1) Jyb \u03b8\u0308 = \u2212f2h + M + Mfw J1\u03b8\u03081 = \u2212M (28) Thus, the mathematical model of the longitudinal dynamics is shown in Fig. 11, where \u03b81 is the pitch angle of the free wing, \u03b11 is the attack of angle of the free wing, and \u03b3 is the track angle. Other parameters have been introduced in Fig. 8. Because \u03b3 = \u03b8 \u2212 \u03b1, one has mX\u0308 = (Tf + Tt \u2212 f1) cos \u03b8 + f2 sin \u03b8 \u2212 L1 sin \u03b3 \u2212 D1 cos \u03b3 mZ\u0308 = \u2212(Tf + Tt \u2212 f1) sin \u03b8 + f2 cos \u03b8 \u2212 mg + L1 cos \u03b3 \u2212 D1 sin \u03b3 Jyb \u03b8\u0308 = \u2212f2h + M + Mfw J1\u03b8\u03081 = \u2212M (29) Set u1 = (Tf 0 + Tt0)/mg, u2 = f20h/J \u03b5 = J /(hmg), \u03b51 = J1/J , u3 = M /J1 (30) x = X g , z = Z g , L = L1 mg , D = D1 mg (31) 1 = ( f + t \u2212 f1) cos \u03b8 + 2r sin \u03b8 mg 2 = \u2212( f + t \u2212 f1) sin \u03b8 + 2r cos \u03b8 mg 3 = Mfw J (32) The equations of motion finally read x\u0308 = u1 cos \u03b8 + \u03b5u2 sin \u03b8 \u2212 L sin \u03b3 \u2212 D cos \u03b3 + 1 z\u0308 = \u2212u1 sin \u03b8 + \u03b5u2 cos \u03b8 + L cos \u03b3 \u2212 D sin \u03b3 \u2212 1 + 2 \u03b8\u0308 = \u2212u2 + \u03b51u3 + 3 \u03b8\u03081 = \u2212u3 (33) The dimensionless parameter \u03b5 is constant",
            " 12(b) and equation (23), one can obtain the force FM generated on the free wing FM l3 l1 sin(\u03b8 \u2212 \u03b81 + \u03b8g ) l2 = M (102) Proc. IMechE Vol. 225 Part G: J. Aerospace Engineering at The University of Manchester Library on April 23, 2015pig.sagepub.comDownloaded from Xinhua Wang and Hai Lin that is FM = Ml2 l1l3 sin(\u03b8 \u2212 \u03b81 + \u03b8g) (103) It is required that \u03b8 \u2208 [\u03b81, 90\u25e6], that is to say, l1, l2, and l3 are not in the same line. Therefore, shaft l3 should be rotated and a force generated for thread collars H2 (see Fig. 11(c)). Ignoring the mass of shaft l2, force Fcollar on thread collars H2 is Fcollar = FM cos \u03b82 = Ml2 l1l3 sin(\u03b8 \u2212 \u03b81 + \u03b8g) cos \u03b82 (104) From equation (104) and u3 = M /J1 in equation (30), the force on thread collars is Fcollar = J1l2u3 l1l3 sin(\u03b8 \u2212 \u03b81 + \u03b8g) cos \u03b82 (105) 6 SIMULATIONS Data for the rotor-fixed wing hybrid aircraft: m = 110 kg, g = 10 m/s2, h = 1.58 m, l = 0.5 m, J1 = 200 kg m2, Jxb = 300 kg m2, Jyb = 850 kg m2, Jzb = 600 kg m2, \u03b5 = 850/(1.58 \u00d7 110 \u00d7 9.8) = 0.5, \u03b51 = 200/ 850 = 0.235, aD = 0"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001642_j.ijsolstr.2010.05.004-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001642_j.ijsolstr.2010.05.004-Figure2-1.png",
        "caption": "Fig. 2. 8-Noded Cosserat point with deformable directors.",
        "texts": [
            " Let X0 \u00bc X0 [ oX0 R3 be a closed, simply-connected body with Lipschitz-continuous boundary and P0i 2 X0 be a point surrounded by a region P 0i \u00bc P 0i [ oP 0i X0 such that X0 \u00bc S i2IP 0i . Let X \u00bc X [ oX R3 be the region occupied by the body in its current configuration. Let there be a 1-1 map x\u0302 : X0 ! X with P 2 X and P \u00bc P [ oP X being the unique images of P0 and P 0 on X, respectively. Let d0 be the position vector of P (Cosserat point) with respect to some inertially-fixed reference frame (Fig. 2). For an 8-noded Cosserat point element, the position vector of any point in P may be represented as: x\u00f0h1; h2; h3\u00de \u00bc d0 \u00fe h1d1 \u00fe h2d2 \u00fe h3d3 \u00fe h1h2d4 \u00fe h2h3d5 \u00fe h1h3d6 \u00fe h1h2h3d7: \u00f03\u00de Here hi (i = 1,2,3) denote the convected coordinate system in P with origin at P and di (i = 1,2,3) represent three linearly-independent directors attached to P. di (i = 4\u20137) are the additional directors, required to describe the inhomogeneous deformation. In the stressfree reference configuration, position vectors are represented as: X\u00f0h1; h2; h3\u00de \u00bc D0 \u00fe h1D1 \u00fe h2D2 \u00fe h3D3 \u00fe h1h2D4 \u00fe h2h3D5 \u00fe h1h3D6 \u00fe h1h2h3D7; \u00f04\u00de b led L B 2n 1n led membrane part (dotted lines) with length L < L0 , (b) the (fictive) non-wrinkled where Di (i = 1,2,3) are three linearly-independent directors associated with P0"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003713_978-1-4614-2419-2-Figure61.13-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003713_978-1-4614-2419-2-Figure61.13-1.png",
        "caption": "Fig. 61.13 Base-upright FEM with addition of steel rib perturbation primarily affecting mode shape",
        "texts": [
            " The full 40,000 DOF BU model is then reduced down to a master set of 52 DOF, which consists of selected directions at each point, denoted by the arrows in Fig. 61.11. This set of 52 DOF will be considered as the a-DOF set for the cases discussed in this paper. The a-DOF set is also the measurement directions that were used to acquire experimental data in the cases that use VIKING to expand test data. The typical mode shapes for the BU model are shown in Fig. 61.12 for reference. 1 in. wide by 2 in. thick steel rib, as shown in Fig. 61.13. The rib is considered to be a realistic, but significant modeling error, especially in the case of a helicopter wing model. Typical FEMs used in industry contain a large number of ribs and stiffeners which result in several potential points for error to be introduced into the model. The rib was included to intentionally create significant differences in the BU mode shapes, to mimic the potential modeling errors that are commonly experienced. For the analytical cases discussed in this paper, mode shapes from the perturbed model are used to expand mode shapes from the unperturbed model"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001324_inds.2009.5227977-Figure4-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001324_inds.2009.5227977-Figure4-1.png",
        "caption": "Fig. 4. Phase control mechanism. (a) Each oscillator has its own phase and firing frequency. (b) Oscillator i receives positive stimulus and promote firing frequency, oscillator k receives negative stimulus and repress the firing frequency. (c) After iterations, the phase offset between each oscillator becomes equal and anti-phase synchronization is realized.",
        "texts": [
            " [16] suggested the following phase shift function: g(\u0394) = \u03b1 sin \u0394 (4) where \u03b1 > 0 is the coupling coefficient of a pulse-coupled oscillator model. When \u0394ji < \u03c0, then g(\u0394ji) > 0 and oscillator j advances the firing frequency to extend the phase offset with respect to oscillator i. On the contrary, when \u0394ji > \u03c0, then g(\u0394ji) < 0 and oscillator j slows down the firing frequency in order to spread the phase offset with respect to oscillator i. After these interactions, the oscillators are assumed to be in a stable anti-phase synchronized state when the following conditions of Eqs. (5) and (6) are fulfilled (Figure 4). \u0394ij = \u0394ji (5) g(\u0394ij) = g(\u0394ji) = 0 (6) We then consider the group N , in which n oscillators are coupled with each other. When oscillator j fires at time tj (t1 < t2 < \u00b7 \u00b7 \u00b7 < tn), it changes the firing frequency \u03c9j as follows: \u0394ji = \u03c6j(ti) \u2212 \u03c6j(ti) (7) \u03c9+ j = \u03c90 + \u2211 k\u2208N g(\u0394jk) (8) When the phase offsets between oscillators which fire consistently are all equal and the repulsive force of all oscillators is negated, the group is assumed to be in a stable anti-phase synchronized state. These conditions are described below together with the case of two oscillators"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001369_j.jsv.2008.09.045-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001369_j.jsv.2008.09.045-Figure2-1.png",
        "caption": "Fig. 2. Free body diagram of a chain link.",
        "texts": [
            " (xc, yc, y)): two translations of the center of mass of the link and one rotation about an axis passing through the link\u2019s center of mass. The force elements take into account the elasticity and damping of the links and joints. In addition to these interconnecting forces, the links also experience contact forces in normal and tangential directions whenever they come in contact with the pulley sheaves. The chain in the CVT is modeled link by link to account for its discrete structure. Figs. 2 and 3 illustrate the free-body diagram of a chain link. It is to be noted that in Fig. 2, the dotted arrows represent the forces (fx and fy), which only arise when a chain link comes in contact with the pulley sheave. Using Newton\u2013Euler formulation [8\u201312] and the Theory of Unilateral Contacts [2,21], the equation of motion for all links under unilateral contact conditions with the pulleys can be written as M\u20acq h \u00f0WN \u00feWSl\u0302SjWT\u00de kN kT ! \u00bc 0, where l\u0302S \u00bc f mi\u00f0 _gi\u00de sign\u00f0 _gi\u00deg. (1) kN, kT are the normal and the sticking constraint forces of the links that are in contact with the pulley and _gi is the relative velocity between a link, i, and the pulley in the sliding plane (refer to Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001724_tmag.2010.2072910-Figure4-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001724_tmag.2010.2072910-Figure4-1.png",
        "caption": "Fig. 4. Magnetic field strength. (a) Slip . (b) Slip . (c) Slip .",
        "texts": [
            " 3(c), the magnetic flux density became smaller at the back core part in comparison to ones of Fig. 3(a) and (b). However, in such larger slip condition, the magnetic flux concentrated at the teeth edge part due to the eddy current. Also, it is evident that the magnetic flux density is easily increased in the rolling direction (RD: horizontal direction of the figures), because of higher magnetic permeability. The magnetic permeability in the rolling direction is about 1.5 times larger than that in the transverse direction. Fig. 4(a)\u2013(c) shows the distributions of the maximum magnetic field intensity of each condition (slip 0, 0.2, and 0.5). As shown in Fig. 4(a), as well as Fig. 3(a), the magnetic field intensity was the largest because the eddy current was very small. With increasing the load and the slip, the magnetic field intensity became smaller as well. In addition, in the transverse direction (TD: vertical direction of the figures), the magnetic field intensity increased because of magnetic anisotropy. Fig. 5 shows the iron-loss distributions calculated from the vector relation between and . Usually, the iron loss is assumed to be in proportion with the square of the maximum magnetic flux density, "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000686_s10409-008-0184-8-Figure3-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000686_s10409-008-0184-8-Figure3-1.png",
        "caption": "Fig. 3 Sketches of the orientation of the wing section and stroke plane. a Hovering; b Forward flight",
        "texts": [
            " However, at forward flight, especially at medium to high flight speeds (ue = 2.5, 3.5 and 4.5 m/s; J = 0.31, 0.44 and 0.57), when varying any one of the control variables ( , \u03b11, \u03c6\u0304 and \u03b12), generally, changes in all the vertical and horizontal forces and pitching moment are produced. The reason for this is explained as following. Let us take the case of varying as an example for the explanation. The orientation of the stroke plane and the attitude of a representative wing section (section at r2) at hovering and at forward flight (ue = 2.5 m/s) are shown in Fig. 3. Keep in mind that for an insect wing, the total aerodynamic force of the wing is approximately normal to the wing surface [10,11]. At hovering, the stroke plane is approximately horizontal (Fig. 3a), and the lift of the wing is approximately in vertical direction and the drag of the wing is approximately in horizontal direction. When is changed, e.g. increased, the translation velocity of the wing, ut ,increases (ut \u223c nr2), hence the lift and drag of the wing increase in both the down- and upstrokes, and moreover, the changes in the lift and drag in the downstroke are the same as their counterparts in the upstroke because \u03b1d and \u03b1u are the almost same. The increases in lift in the down- and upstrokes result in the change in the mean vertical force",
            " This explains why only the mean vertical force is changed when varying . Since the line of action of the mean vertical force passes the position of the mean stroke angle, the line of action of the mean vertical force does not change when varying ; thus the mean pitching moment does not change. At forward flight, especially at medium to high flight speeds, the stroke plane tilts forward and the wing surface is approximately horizontal during the downstroke and approximately vertical during the upstroke (Fig. 3b). Since the total aerodynamic force of the wing is approximately normal to the wing surface, when , hence ut , is changed, e.g. increased, a change in vertical force will be produced in the downstroke and a change in horizontal force will be produced in the upstroke. This explains the changes in the mean vertical and horizontal forces produced by varying . Since the wing is well above the center of mass of the insect, the change in the mean horizontal force will produce a change in pitching moment (the change in the vertical force will also produce a change in pitching moment, but it is relatively small because the line of action of the change in the mean vertical force is generally close to the center of mass of the insect); this explains the change in the mean pitching moment produced by varying "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003447_acc.2013.6580292-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003447_acc.2013.6580292-Figure2-1.png",
        "caption": "Fig. 2. A sketch of the satellite control system",
        "texts": [
            " Remark 5: From Theorems 1 and 2, for a given threshold \u03b5 > 0, the controller parameters and the event-triggering parameter \u2126 > 0 can be co-designed. Similarly, for given \u2126>0, the controller parameters and the threshold can also be co-designed. However, although impulsive system approach [16] and passive method [17] are employed to discuss the event-based dynamic output feedback control, the controller parameters are not easy to be designed, which certainly limits the applications of the proposed results. Suppose that the plant in Fig. 1 is a satellite control system [21]. A sketch of the satellite control system is shown in Fig. 2, where it is assumed that two masses are connected by a spring with torque constant k and viscous damping constant d. The equations of motion from Fig. 2 are given by { J1\u03b8\u03081 + d(\u03b8\u03071 \u2212 \u03b8\u03072) + k(\u03b81 \u2212 \u03b82) = Tc J2\u03b8\u03082 + d(\u03b8\u03072 \u2212 \u03b8\u03071) + k(\u03b82 \u2212 \u03b81) = 0 (35) where Tc is the control torque and J1 and J2 are inertias. Set the state vector to be x = col{\u03b82, \u03b8\u03072, \u03b81, \u03b8\u03071} and set the control input to be u = Tc. Then the state-space representation of (35) can be given by (1) with A =     0 1 0 0 \u2212 k J2 \u2212 d J2 k J2 d J2 0 0 0 1 k J1 d J1 \u2212 k J1 \u2212 d J1     , B =     0 0 0 1 J1     C = [ 0 0 1 0 1 0 1 0 ] In this paper, we suppose J1 = J2 = 1, k = 0.09 and d = 0"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002043_pime_proc_1967_182_025_02-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002043_pime_proc_1967_182_025_02-Figure1-1.png",
        "caption": "Fig. 1. Characteristic frequencies of general system",
        "texts": [
            " -AflflS22+E,, Vol I82 Pt 1 No 13 Proc Instn Mech Engrs 1967-68 at NANYANG TECH UNIV LIBRARY on June 9, 2016pme.sagepub.comDownloaded from INFLUENCE OF COUPLED ASYMMETRIC BEARINGS ON THE MOTION OF A MASSIVE FLEXIBLE ROTOR 259 Expressions (6) for Z,, Z , remain unchanged but more complicated expressions of the type shown in equation (12) must be evaluated in order to find the rotor impedance. Equation (12) is applicable to only the symmetric modes. A similar expression must be separately evaluated for the asymmetric rotor modes, A,,, A,,, etc. now becoming $Z/12 , f 2 (m74,+mh24\u2019,), respectively. Fig. 1 illustrates how the equivalent impedances Z , and Z, influence the juxtaposition of corresponding undamped frequencies for a rotor system with many degrees of freedom. In this respect this system does not differ from the simple system previously described. . . * (13) where A,, and Aa, are defined in equation (12) and A , is the first minor of the determinant found by interchanging the 1st and rth rows of A,, or Ayv. The expressions for the forced vibrations at a bearing for the general system are shown in equations (13) above",
            " Assessment of the results of such a study would be greatly facilitated by an appreciation of the physical significance of these coefficients. The author\u2019s \u2018Single rotor mode\u2019 section is well suited to a discussion of this significance, in that it deals with a relatively simple rotor-bearing system which is amenable to analytical treatment. The results obtained are thus not complicated by an excessive number of parameters. The author\u2019s more general \u2018Unlimited rotor modes\u2019 section appears to be based upon the assumption of rotor symmetry about the centre of span. Validity of the results obtained therein and pictured in Fig. 1 is claimed for both symmetric and asymmetric rotor vibration modes. While these results appear adequate to allow the designer to make predictions about a similarly shaped rotor, it seems natural to ask how results may be obtained which are pertinent to a rotor having more arbitrary geometry. Perhaps such an analysis is contemplated for the future. Proc Instn Mech Engi-s 1967-68 The results obtained in the \u2018Self-excited\u2019 section and pictured in Fig. 2 are useful and instructive, but the comments of the previous paragraph again apply",
            " Here the term \u2018damping\u2019 refers to the quadrature component of the equivalent impedance and this is very much influenced by the bearing cross stiffness. I agree that it would have been useful to give a sample response calculation and to list bearing impedances for a range of eccentricity ratios and bearing types. The necessity for keeping the paper within the length prescribed by the Institution prevented this. The first point raised by Professor Downham and Mr Woods concerns the modal representation of the rotor. If the requirement is merely to find the undamped characteristic frequencies of the system (Fig. 1) or to establish the instability boundary (Fig. 2), then all that is required is the impedance of the rotor \u2018looking in\u2019 from the bearing positions. This may be found using pinned-pinned modes (equation (12)) or free-free modes (reciprocal imProc Instn Mech Engrs 196768 m pedance = 2 + z ( u ) / ~ r ( ~ ~ - ~ z ) defined in (5)) . Alterna- tively one could employ say the Myklestad technique to evaluate the natural frequencies for a series of rotor-spring support combinations and derive the rotor impedance diagram directly"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001005_ramech.2008.4681401-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001005_ramech.2008.4681401-Figure2-1.png",
        "caption": "Fig. 2 a 5-DOF manipulator for experiment in its reference configuration",
        "texts": [
            " Obviously the equation (1) has a same form as that of the subproblem 1, so we can solve the equation (1) to find O1 and 02 02 = a tan 2(42'(u'xz2),uTz ) 01l = a tan 2(coT(zlx v'), z\"'V) (4 Where u' and Z' are the projections of u and z2 on C0222 and z', v are the projections of z1, v on c4. If there are multiple solutions for c, each of these solutions gives a value for 0, an 0.Two solutions exist in the case where the circles in Figure 1 intersect at two points, one solution when the circles are tangential, and none when the circles fail to intersect. III. EXPERIMENTS To verify the correctness of the solution the extended sub-problem, we configure a 5-DOF manipulator for which the extended sub-problem must be employed, as shown in Figure2. Z2 =acwl+/ko2 +)(CO x 02 ) (6) Z, = acol +o2+ y(co1xco2) + d = acol +/c2+(r+ jjdjj)(c1xco2) (7) (2) Solve for 03, 04 and 05 Since 01 and 02 are known, so e J3 3e J404eJ55 =e-e2 2 e- IAg1 Let e J2eC2e lg1 = g2 ' we get e J3 3eJ;4 4eJ;55 = 92 (17) In figure 2, The axes ofjoint 1 and joint 2 are perpendicular to each other and not intersecting, and the axes ofjoint4,5,6 intersect atapoint qwi Where, l1 =355, 12= 245, 13= 90, 14= 300, 15 = 180. If 0 = 0 , the configuration of the fixed coordinate relative to the end-effector coordinator is given by 1 0 0 /3 0 10 '2 gst(0) 0 0 1 1I+14+15 _0 0 0 1 and we can choose axis points [fl0 2 [/] qw ['2/4\u00b01 01o 0l CtdC03 5= 0 C02= 0 C04 1 With the provided q, and Ci (i = 1, 2, ... 5), we can get the twists of the manipulator"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000399_s10999-008-9077-z-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000399_s10999-008-9077-z-Figure2-1.png",
        "caption": "Fig. 2 Variation of contact stiffness with relative nodal displacement: (a) normal stiffness KN and (b) tangential stiffness KT",
        "texts": [
            " The normal stiffness KN is used to penalize interpenetration between the two bodies, while the tangential stiffness KT is used to approximate the sudden jump in the tangential force, as represented by Coulomb\u2019s friction law when sliding is detected between the two contacting nodes. Advantages of contact elements Because of the simplicity of their formulation, contact elements enjoy the following advantages: (i) they are simple to formulate, (ii) they are easily accommodated into existing FE codes, and (iii) they are easy to use. However, experience with contact elements indicates that they suffer from several difficulties as indicated below. Disadvantages of contact elements Figure 2a, b show the variation of the normal and tangential stiffness, KN and KT, with displacement. A major problem in the implementation of contact element is the assignment of values to KN and KT. The values of KN and KT are required to be very large. However, the use of excessively high values of KN and KT results in ill-conditioned global stiffness matrices, leading to numerical errors and divergence. On the other hand, the use of smaller values of KN and KT, results in convergence to the wrong solution allowing for interpenetration and wrong estimates of the stick and slip regions"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000597_978-3-540-73129-0-Figure6.22-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000597_978-3-540-73129-0-Figure6.22-1.png",
        "caption": "Fig. 6.22. Elliptec motor. a Basic structure, b examples of application (by [13])",
        "texts": [
            " The use of this motor in direct drive means that the complete function is obtained without any additional gear mechanism (for speed reduction, or for converting rotation in translation). Optics is probably the domain where the use of the piezoelectric motors is the most advanced. The most famous example, is the Canon camera, which includes an auto focus zoom based on a piezoelectric ultrasonic motor (USM) since 1992 (Fig. 6.21) [12]. Several other concept have been developed since then; few of them have found industrial applications. The motor from Elliptec is using a multilayer component, encased in a structure to couple two flexural modes of the beam (Fig. 6.22a) [13]. The stator includes a play recovering mechanism in the form of a spring that: \u2013 applies the preload force between the vibrating stator and the moving member; \u2013 guides the stator; \u2013 decouples the vibrations in the stator from the ground. Such a vibrating stator can be implemented in various ways (Fig. 6.22b). Several concepts of quasistatic motors exists as well. One of them is using at least one pair of amplified piezo actuators. The basic working principle of the Cedrat stepping piezomotor concept is illustrated through a simplified linear model based on a pair of APAs (Fig. 6.23). The displacements and forces produced by the APAs are transferred to the slider or the rotor by friction. At least one pair of APAs is used in the following conditions: held by their centre, the APAs are actuated in opposite phase"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001960_s12239-009-0053-x-Figure5-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001960_s12239-009-0053-x-Figure5-1.png",
        "caption": "Figure 5. Nomenclature of a gear tooth.",
        "texts": [
            " The important tangential force has the following relationships: (1) (2) where r is the gear radius; \u03b8 is the tooth rotational angle (bending slope) due to bending at a moving contact point during one rotation; the subscripts g and p are a gear and a pinion, respectively; Ft is the transmitted tangential load (N); n is speed (rpm); T is torque (Nm); and V is pitch line velocity (m/s). The equivalent tooth bending stiffness k in equation (1) can be obtained analytically by Castigliano\u2019s theorem, which is based on a single gear tooth (Figure 5). (3) where , , l2 is the clearance under a base circle, l1 is the working depth, E is Young\u2019s modulus and b is the gear face width. Consequently, the transmitted tangential force (Ft) may be calculated by ADAMS (with the option of a function input) in conjunction with k in equation (3), and it varies as gears rotate. Considering k, the influence of tooth bending deflections can be included in the equivalent model. 2.2. Torsional Stiffness of Shafts In the equivalent model, the torsional stiffness of transmission shafts is represented in ADAMS as several torsional springs with torsional rates at each shaft section: (4) where L is the solid shaft length between two gears, G is the shear modulus, and J is the polar area moment of inertia"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002309_s11012-010-9380-2-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002309_s11012-010-9380-2-Figure1-1.png",
        "caption": "Fig. 1 Redundantly full-actuated planar 2RPR/RP manipulator",
        "texts": [
            " A model for redundantly actuated PKM is introduced in Sect. 3. The resolution of ac- tuation redundancy and its use for secondary tasks is discussed in Sect. 4. Section 5 is dedicated to the model-based control. It shown that kinematic uncertainties lead to interferences of the drives. To cope with these effects a corrected augmented PD and computed torque control scheme is introduced. The motivation for the introduction of redundant actuators is best clarified by means of a simple example such as the 2RPR/1RP PKM in Fig. 1. The PKM has a DOF 2, and can be uniquely positioned by controlling two of the prismatic joints. The actuation of all three prismatic joints leads to redundant actuation. 2.1 Manipulability/dexterity/singularities The possibility of avoiding singularities might have been the catalyst for the interest in redundant actuation of PKM. A non-redundant PKM is in a singularity if the EE-motion cannot be determined by the actuated joints (this is by no means a precise definition, but sufficient for the following\u2014more complete treatises are [18, 21])"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003001_00405001003696464-Figure7-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003001_00405001003696464-Figure7-1.png",
        "caption": "Figure 7. FE model of woven SMA fabric: (a) the mesh, and (b) a contact pair.",
        "texts": [
            " As the crimp of the wires does not change much during uniaxial or biaxial deformation, the \u201cafter-plateau\u201d modulus is the same for both The well-defined multilinear character of the tensile diagrams of SMA fabrics, which preserves the multilinear character of the wire tensile diagram, is the result of approximations, inherent to the approximate model: the wire is modelled as a material line with certain bending rigidity. This assumption takes into account the distribution of strain in the wire cross-section only approximately. By the integral effect, this distribution has on the overall bending rigidity. The FE modelling, discussed next, will reveal more \u201csmoothened\u201d tensile diagrams. Building and solving an FE model of woven fabric Figure 7(a) shows ANSYS FE model of the woven SMA fabric. Note that the fabric geometry is symmetrised for the transition to FE model from WiseTex. The material model is defined as multilinear elastic, with the tensile diagram shown in Figure 2(b). The elements used are 3D structural eight-node solid elements of the type SOLID45 (prismatic). Figure 7. FE model of woven SMA fabric: (a) the mesh and (b) a contact pair.Contact pairs are defined at the contact zones on the yarns, as shown in Figure 7(b). There is important interpenetration of the volumes of the yarns. This is a common difficulty for the FE modelling of textile structures. Geometrical models, like the models used here, use several simplifying assumptions. The shape of the yarn\u2019s central line prescribes the positions of the centres of the cross-sections. The model calculates the dimensions of the cross-sections, ensuring that the distance between the central lines of the contacting yarn is equal to the sum of their dimensions"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002945_s10846-011-9649-2-Figure6-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002945_s10846-011-9649-2-Figure6-1.png",
        "caption": "Fig. 6 Output distribution of each state in Markov chain for N = 100 and aRcom = 5, b Rcom = 10, c Rcom = 12, and d Rcom = 15",
        "texts": [
            " The convergence for varied initial distributions can be added to these graphs (Figs. 4 and 5) by finding \u2223 \u2223\u03bd Pt \u2212 \u03bcPt \u2223 \u2223 with respect to t for any of the experimental cases. It is important to note that for each scenario, any initial distribution (\u03bd) will fall below the corresponding line and hence converges faster than the contraction coefficient graph of the Markov kernel as in the proof of Theorem 2. Figures 6 and 7 represent the possible outcome percentages of each state in the Markov chain for varying communication ranges and numbers of nodes, respectively. In Fig. 6, the probability of being in the ideal state for a mobile node is higher than the other states (44%) when Rcom = 5. It demonstrates that nearly half of the mobile nodes reach the state where they have the desired number of neighbors and locations that result in minimal external force. The probability of reaching the stop state with a non-ideal number of neighbors is approximately 10%. The remaining states that are not explicitly labeled with values ranging from 1% to 8% represent states where the node is moving and has an ideal number of neighbors",
            " As seen from these results, to reach and stay at the desired state for a mobile node is less probable when communication range increases. It is an expected result since larger communication range means more local neighborhood information and more neighboring nodes that results in a less stable position (aggregated force on a mobile node is not zero). Figure 7 shows the possible outcome percentages of each state in the Markov chain for the experiment comparing varying numbers of nodes with Rcom = 10 and N = 125 and 150, respectively. Referring back to Fig. 6b, Rcom = 10 and N = 100, the probability of being in the ideal state for a mobile node is 32%, the probability of being in the stopped, non-ideal state is only 20% and the sum of all remaining moving states is 48%. When there are a total of 125 mobile nodes in a manet, the probability of being in the ideal state is 27% as seen in Fig. 7a. When N = 150 it can be seen that the probability of being in the ideal state is 18%, the probability of being in the stopped, non-ideal state is now 30% and the sum of all remaining moving states is now 42%"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001916_ssp.164.339-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001916_ssp.164.339-Figure1-1.png",
        "caption": "Fig. 1. The model of complex systems vibrating transversally in rotational transportation",
        "texts": [
            " Specified algorithm of derivation of the dynamical flexibility of analyzed systems is based on equations of motion derived by classical methods and orthogonalization of these equations and searching solution by Fourier series (a sum of products of known eigenfunctions of displacement and eigenfunctions of time variable). The paper is a part of research work series [1, 7, 8] considering the transportation effect and it is a part of these works where the methods and theory of dynamical systems in transportation were presented. The following symbols were used in Fig. 1 for individual beams: \u03c1 \u2013 mass-density, A \u2013 crosssection, l \u2013 length of beam, x \u2013 location of analyzed cross-section, b \u2013 damping factor, M \u2013 mass of the beam, \u03c9 \u2013 angular velocity, \u2126 \u2013 frequency, Q \u2013 rotation matrix, S \u2013 position vector, Fd \u2013 damping force, F \u2013 harmonic excitation force, E \u2013 Young\u2019s modulus, IZ \u2013 geometric moment of inertia, w \u2013 vector of displacement. The presented model (Fig. 1) is based on simple system models (mechanical and mechatronical). The starting element is applied as the beam fixed on a rotational table (in place of origin global reference frame) and consecutive systems treated as the free-free ones (with known boundary values at the end and the beginning of system) with maintaining 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_12_0000983_00207170802654410-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000983_00207170802654410-Figure1-1.png",
        "caption": "Figure 1. Potential forces acting at CG and CB and the righting arm for a submerged spherical vehicle.",
        "texts": [
            " Writing this in matrix form we can express the potential forces from gravity and buoyancy as accelerations by G#P\u00f0 0\u00f0t\u00de\u00de \u00bc 1 m1 \u00f0W B\u00des 1 m2 \u00f0W B\u00dec s 1 m3 \u00f0W B\u00dec c 1 j1 \u00f0\u00f0 yGW yBB\u00dec c \u00f0zGW zBB\u00dec s \u00de 1 j2 \u00f0\u00f0zGW zBB\u00des \u00f0xGW xBB\u00dec c \u00de 1 j3 \u00f0\u00f0xGW xBB\u00dec s \u00fe \u00f0 yGW yBB\u00des \u00de 2666666666666666666664 3777777777777777777775 : \u00f013\u00de Note that if CG 6\u00bc CB, the two opposing restoring forces will induce a torque, referred to as the righting moment, if the vehicle rotates. The righting arm GZ depends on the distance between CG and CB and the list angle as seen in Figure 1. On the other hand, D ow nl oa de d by [ U ni ve rs ity o f B or as ] at 1 5: 12 0 5 O ct ob er 2 01 4 if CG\u00bcCB, then the vehicle will experience no torque that opposes orientation displacements. Not all external forces interacting with a submerged rigid body can be derived from a potential function. One example is the external force due to the viscosity of the surrounding fluid, or in our specific case, viscous drag. Such a force is called a dissipative force since it dissipates energy from the system"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001823_j.rcim.2010.06.010-Figure7-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001823_j.rcim.2010.06.010-Figure7-1.png",
        "caption": "Fig. 7. Tool path and cutting f",
        "texts": [
            " We define kinematic capacity as the ability of the cell to give the effector a speed in a given direction. This capacity depends first of all on maximum articular speeds. The behavior of our cell being non-isotropic because of its architecture, this capacity also depends on the pose of the effector and on the considered direction. In our application we seek to improve this kinematic capacity with two objectives: to decrease articular requests in order to preserve the structure and to maintain a margin for improvement with respect to the advance speed Vf relative to the process (Fig. 7). Our objective is to improve machine behavior while exploiting the degrees of freedom introduced by the architecture redundancy. We formalize this step as an optimization of the machine and kinematics capacities of the cell under process constraints. The total degree of redundancy r is the sum of functional redundancy rf and structural redundancy rs [10]. The degree of functional redundancy rf is defined by rf \u00bc dim\u00f0EO\u00de dim\u00f0ET \u00de \u00f02a\u00de where rf is the difference between the operational space EO dimension and the task space dimension",
            " They will be defined with respect to the characteristics of the task. 4. Finishing constraints 4.1. Constraints related to the part The constraints related to the part come from the functional requirements: form defect, morphological constraint and surface quality: volume and accessibility. 4.2. Constraints related to the process The process constraints are related to the selected strategy. For machining and polishing, we use a strategy of 5-axis sweeping, which defines the advance direction d ! , advance speed V ! f and the angle of the tool [14] (Fig. 7). The trajectory is a curve, discretized by all of the poses defined by the position of a characteristic point of the tool and by the direction of the tool axis n ! . The choice of the strategy parameters is already a problem which is constrained with respect to the geometrical surface specifications. 4.3. Machining characteristics In machining, material removal is materialized by the envelope of the cutting edge path. Reducing the advance speed involves a reduction in the tooth advance fz. This does not modify the overall path envelope but can cause a deterioration in cutting quality, resulting in premature wear, an oscillatory mode or a degradation of the surface quality",
            " 8): c1 \u00bc q1\u00feq2\u00feq3 3 \u00f06\u00de Indeed, the deformation induced by the solicitation is relatively proportional to the median leg length. The lower this value, the greater the rigidity of the substructure. The lever arm being weak relative to the weight applied to the wrist, we neglect its influence on the calculation of the criterion. 6.1.2.2. Cutting pressures. Modeling the cut [2] makes it possible to evaluate the direction of the efforts during machining. An experimental study on a Kistler table enabled us to determine the resultant F ! nc(Fig. 7) in the plane normal to the tool axis n ! . This is directed on average at 1111 compared to the advance direction d ! . We seek to minimize the torque induced by the component F ! nc cutting pressures on the last axis associated with q6. We then define a second criterion by c2 \u00bc : u ! Fnc 4u , 6: \u00f07\u00de with u ! Fnc the unit directing vector of the resultant of the cutting pressures in the normal plane and u ! 6 the directing vector of the last axis of the wrist q6. The purpose of the kinematic study is to express a performance criterion related to the improvement of kinematic behavior"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002342_j.mechmachtheory.2011.06.006-Figure8-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002342_j.mechmachtheory.2011.06.006-Figure8-1.png",
        "caption": "Fig. 8. A multibody model of a Puma manipulator [41].",
        "texts": [
            " The corresponding parameters required for the calculation of reaction forces in joints 4 and 5 are listed as follows: 3 \u03f1\u21923 = 0; 0;\u2212\u20183\u00bd T ; 4 \u03f1\u21924 = 0; 0;\u2212\u20184\u00bd T ; 5 \u03f1\u21925 = 0;0;\u2212\u20185\u00bd T ; 3 \u03f1\u2192C3 = 0; 0; \u20183 =2\u00bd T ; 4 \u03f1\u2192 C4 = 0; 0; \u20184 =2\u00bd T ; 5 \u03f1\u2192 C5 = 0; 0; \u20185 =2\u00bd T ; \u0398 1\u00f0 \u00de = 1\u00bd ; \u0398 2\u00f0 \u00de = 1; 1\u00bd T ; \u0398 3\u00f0 \u00de = 1; 1; 1\u00bd T ; \u0398 4\u00f0 \u00de = 1; 1; 1; 1\u00bd T ; \u0398 5\u00f0 \u00de = 1; 1; 1; 1; 1\u00bd T : Taking the above relations into account, the reaction forces in joints 4 and 5 are determined by: R4\u03b73 = q 3 \u20183 m4 + m5\u00f0 \u00de + q 1 m4 + m5\u00f0 \u00decos q3 + \u20184 2 q 3\u2212q 4 m4 + 2m5\u00f0 \u00decos q4 + q 2 + g m4sin q3 + \u20184 2 q 3\u2212q 4 2m4sin q4 + m5 2 h q 3\u2212q 4 + q 5 \u20185cos q4\u2212q5\u00f0 \u00de + 2 q 2 + g sin q3 + 2 q 3\u2212q 4 2 \u20184sin q4 + q 3\u2212q 4 + q 5 2 \u20185sin q4\u2212q5\u00f0 \u00de i \u2212N\u03b75 cos q4\u2212q5\u00f0 \u00de\u2212N\u03b65sin q4\u2212q5\u00f0 \u00de: \u00f060\u00de R4\u03b63 = q 2 3 \u20183 m4 + m5\u00f0 \u00de + q 2 + g m4 + m5\u00f0 \u00decos q3 + \u20184 2 q 3\u2212q 4 2 m4 + 2m5\u00f0 \u00decos q4\u2212q 1m4sinq3 + \u20184 2 q 4\u2212q 3 m4sin q4 + m5 2 q 3\u2212q 4 + q 5 2 \u20185cos q4\u2212q5\u00f0 \u00de\u22122q 1sin q3 + 2 q 4\u2212q 3 \u20184sin q4\u2212 q 3\u2212q 4 + q 5 \u20185sin q4\u2212q5\u00f0 \u00de h i + N\u03b75 sin q4\u2212q5\u00f0 \u00de\u2212N\u03b65cos q4\u2212q5\u00f0 \u00de: \u00f061\u00de R5\u03b74 = m5 2 q 3\u2212q 4 + q 5 \u20185cosq5 + 2 q 3\u2212q 4 \u20184 + q 1cos q3\u2212q4\u00f0 \u00de + q 3\u20183cosq4 + q 2 + g sin q3\u2212q4\u00f0 \u00de\u2212q 23\u20183sinq4 \u2212 q 3\u2212q 4 + q 5 2 \u20185sinq5 h i \u2212N\u03b75 cosq5 + N\u03b65 sinq5: \u00f062\u00de R5\u03b64 = m5 2 q 3\u2212q 4 + q 5 2 \u20185cosq5 + 2 q 3\u2212q 4 2 \u20184 + q 2 + g cos q3\u2212q4\u00f0 \u00de + q 23\u20183cosq4\u2212q 1sin q3\u2212q4\u00f0 \u00de + q 3\u20183sinq4 + q 3\u2212q 4 + q 5 \u20185sinq5 h i \u2212N\u03b75 sinq5\u2212N\u03b65cosq5: \u00f063\u00de Using the substitutions q3=\u03c62, q3\u2212q4=\u03c63, and q3\u2212q4+q5=\u03c64 as well as the relations 0A3 3 R \u2192 4 and 0A4 4 R \u2192 5, in which it is taken that 0 N \u2192 = 0; Ny; Nz T =0A5 5 N \u2192 , the results that coincidewith Blajer's results [17] are obtained (there are only differences in the notations of some constant system parameters). 6.2. A Puma manipulator Let us consider a Puma manipulator [40] as a second example of the application of the algorithm. According to [41], the kinematic scheme of this robot without the robot's hand is shown in Fig. 8. The corresponding parameters of the robot, required for the calculation of joint reactions, are listed as follows (see [41]): 1 \u03f1\u2192C1 = \u22120:14; 0; 0\u00bd T ; 2 \u03f1\u2192C2 = \u22120:175; 0; 0\u00bd T ; 3 \u03f1\u2192C3 = \u22120:13; 0; 0\u00bd T ;1 \u03f1\u21921 = 0:14; 0; 0:5\u00bd T ; 2 \u03f1\u21922 = 0:35; 0; 0\u00bd T ; 3 \u03f1\u21923 = 0:36; 0; 0\u00bd T ;1 e\u21921 = 0; 0; 1\u00bd T ; 2 e\u21922 = 0; 1; 0\u00bd T ; 3 e\u21923 = 0; 0; 1\u00bd T ; \u0398 1\u00f0 \u00de = 1\u00bd ; \u0398 2\u00f0 \u00de = 1; 1\u00bd T ; \u0398 3\u00f0 \u00de = 1; 1; 1\u00bd T ;0\u2192F1 = 0; 0;\u2212m1g\u00bd T ; 0 F \u2192 2 = 0; 0;\u2212m2g\u00bd T ; 0 F \u2192 3 = 0; 0;\u2212m3g\u00bd T ; 0M \u2192 1= 0M \u2192 2= 0M \u2192 3 = 0; 0; 0\u00bd T ;1 I1 = diag 0; 0; 0:16\u00f0 \u00de;2 I2 = diag 0:00377; 0:0353; 0:0353\u00f0 \u00de; 3I3 = diag 0:00031; 0:00488; 0:00488\u00f0 \u00de;m1 = 5; m2 = 2:45; m3 = 0:723; wheremi, \u03f1 \u2192 i, \u03f1 \u2192 Ci , e\u2192i, and Ii are given in units of kg,m,m,m, and kg\u22c5m2, respectively, and where it is taken that the gravitational acceleration is g=9.81 m/s2. It is obvious from Fig. 8 that A3 *= I. Applying the relations from Section 3.2, the joint reaction force and the torque of joint reaction couple in joint 3, expressed in local frame C3\u03be3\u03b73\u03b63, read: R3\u03be3 = \u22127:09263sin q2 cos q3\u22120:16629q 2 3 \u22120:25305q 22 cos q3\u22120:16629q 22 cos2q3 + 0:10122q 1 sin q3 + 0:25305q 1 cos q2 sin q3\u22120:33258q 1q 3 cos q2\u22120:5061q 1q 2 sin q2 sin q3\u22120:33258q 1q 2 cos q3 sin q2 sin q3 \u22120:10122q 21 cos q2 cos q3\u22120:25305q 21 cos2 q2 cos q3\u22120:16629q 21 + 0:16629q 21 cos2 q3 sin2q2; R3\u03b73 = 0:16629q 3 + 7:09263 sin q2 sin q3 + 0:16629q 1 cos q2 + 0:10122q 1 cos q3 + 0:25305q 1 cos q2 cos q3 \u22120:5061q 1q 2 cos q3 sin q2\u22120:33258q 1 q 2 cos2 q3 sin q2 + 0:25305q 22 sin q3 + 0:16629q 22 cos q3 sin q3 \u22120:16629q 21 cos q3 sin2 q2 sin q3 + 0:10122q 21 cos q2 sin q3 + 0:25305q 21 cos2 q2 sin q3; R3\u03b63 = \u22120:25305q 2 \u22120:16629q 2 cos q3 + 7:09263 cos q2\u22120:16629q 1 sin q2 sin q3\u22120:10122q 2 1 sin q2 \u22120:33258q 1 q 3 cos q3 sin q2 + 0:33258q 2 q3 sin q3\u22120:25305q 21 cos q2 sin q2\u22120:16629q 21 cos q2 cos q3 sin q2; MR3\u03be3 = 0:00031q 2 q 3 cos q3\u22120:00031q 1 q 2 cos q2 cos q3 + 0:00031q 1 q 3 sin q2 sin q3\u22120:00031q 1 cos q3 sin q2 + 0:00031q 2 sin q3; MR3\u03b73 = \u22121:6313 cos q2 + 0:0582q 2 + 0:43127q 2 cos q3 + 0:43127q 1 sin q2 sin q3 + 0:02328q 2 1 sin q2 + 0:00031q 1q 2 cos q2 sin q3 + 0:08594q 1 q 3 cos q3 sin q2\u22120:08594q 2 q 3 sin q3 + 0:0582q 2 1 cos q2 sin q2 + 0:04282q 21 cos q2 cos q3 sin q2: See [15] for the expressions of reaction forces in joint 2"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001347_bfb0109668-Figure3.1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001347_bfb0109668-Figure3.1-1.png",
        "caption": "Fig. 3.1. Front wheel drive car with a camera mounted above the center O.",
        "texts": [
            "8) of ~3 is of 3 dimensions. Actually, for a general curve which is not necessarily of linear curvature, one still can show that the shape of the image curve is controllable only up to its linear curvature parameters ~3: T h e o r e m 3.2. ( G e n e r a l C u r v e C o n t r o l l a b i l i t y ) The locally reachable space of ~ under the motion of an arbitrary ground-based mobile robot has at most 3 dimensions. Similar results can be obtained for the model of a front wheel drive car as shown in Figure 3.1. The kinematics of the front wheel drive car (relative to the spatial frame) is given by = sin 9ul = cos/~Ul /~ = 1-1 tan au l ~ - - u 2 (3.2) Comparing (3.2) to the kinematics of the unicycle, we have: w ---- l-1 tan c~ul, v -- Ul. From the system (2.6), the dynamics of the image of a ground curve under the motion of a front wheel drive car is given by 0 ) (1) ---- 1-1 t a n ~ f l -t- f2 ul + u2 ---- /1~1 -~- L U 2 . (3.3) By calculating the controllability Lie algebra for this system, one can show that the controllability for the front wheel drive car is the same as the unicycle"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002634_s12541-012-0036-0-Figure7-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002634_s12541-012-0036-0-Figure7-1.png",
        "caption": "Fig. 7 Curing angle according to the intensity distribution of the patterned beam",
        "texts": [
            " 1 and 2, the lateral magnification MT can be expressed as the ratio of s\u2032 to s: ' ' T y s M y s = = (3) To reduce the image height through the objective lens, the image plane must be positioned within the effective focal length (s1). If the resin surface is set on the image plane, the side of the cured layer is inclined by curing angle \u03b8. Figure 6 shows \u03b8 according to the position of the resin surface. The curing angle affects the precision of the layered microstructure. The curing angle also varies due to Gaussian superposition of the patterned beam. In Fig. 7(a), the position of the resin surface corresponding to the image plane is written as RS1 and RS2, and the intensity distribution of the patterned beam is drawn with a dotted curve. Figures 7(b) and (c) represent the curing angle when a patterned beam is formed at the position of RS1. In Fig. 7(b), the critical exposure energy Ec is low, and the curing angle \u03b81 occurs when the minimum exposure energy Emin is higher than the critical exposure energy of the resin. In Fig. 7(c), the critical exposure energy is high, and the curing angle \u03b82 occurs when the minimum exposure energy Emin is lower than the Ec. Figures 7(d) and (e) represent the curing angle when a patterned beam is formed at the position of RS2. The conditions for Ec, Emin, and Emax are the same as shown in Figs. 7(b) and (c), respectively. The curing angle in Fig. 7(d) is the same as the angle of the incident ray. On the other hand, the curing angle shown in Fig. 7(e) is smaller than the angle of the incident ray due to the high value of Ec. When using a photoabsorber such as Tinuvin, the curing angle can be reduced. The curing angle is also proportional to the layer thickness due to the amount of exposure energy. As previously mentioned, due to the patterned beam intensity, the curing angle is affected by the position of the resin surface and the curing properties. The curing angle can vary with the amount of photo-initiator and photo-absorber. By controlling the curing angle, the amount of stair-step can be reduced"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000164_aina.2006.24-Figure3-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000164_aina.2006.24-Figure3-1.png",
        "caption": "Figure 3. Multi-actor/multi-sensor (MAMS) model",
        "texts": [
            " On the other hand, in the multi-actor (MA) model [1], each sensor sends sensed information to one actor but some pair of sensors in an event area may send sensed information of an event to different actors. In order for each actor to make the decision, more number of sensors are required. In the SA and MA models, each sensor sends sensed information to one actor in each event area. In this paper, we propose a multi-actor/multi-sensor (MAMS) model to realize the faulttolerant WSAN. Each sensor sends sensed information to multiple actors and each actor receives sensed information from multiple sensors in an event area [Figure 3]. Each sensor broadcasts a message of sensed information in wireless medium. An area where each sensor can deliver a message with wireless medium is referred to as broadcast cell. Durresi et al. [5] discuss how to distribute sensors in a area so that every event to occur can be sensed by some number of sensors. An actuation device is modeled to be an object in this paper. An action is modeled to be the execution of a method on an object. On receipt of a method issued by an actor, the method is performed on an actuation device object"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001065_13506501jet487-Figure6-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001065_13506501jet487-Figure6-1.png",
        "caption": "Fig. 6 Forces acting on big-end bearing",
        "texts": [
            " The acceleration of the reciprocating part aais given by aa = d2 dt 2 S = \u2212Ra\u03c9 2 ( cos(\u03b8)+ \u03bb cos(2\u03b8) + \u03bb3 sin4(\u03b8)\u221a (1\u2212\u03bb2 sin2(\u03b8))3 ) where \u03bb = crank length/connecting rod length and Ra = crank radius. Details of the equation can be found in Mufti and Priest [30]. The reciprocating inertial force can be expressed as Frec = ( mp + 1 3 mc ) aa (2) Also the rotating inertial force is given by Frot = 2 3 mcRa\u03c9 2 (3) as mc = connecting rod mass and mp = piston assembly mass. Using the connecting rod two-point mass system, the side thrust force due to the obliquity of the connecting rod acting through the piston normal to the cylinder liner can be calculated according to Fig. 6 as FT = (FG + Frec) tan(\u03c8) (4) where \u03c8 = angle between the connecting rod axis and the cylinder liner axis. The resultant of the vector sum of Frec, FG, FT, and Frot forces (Fig. 6) provides the instantaneous load on the big-end bearing \u2018FB\u2019 at any particular crank angle. In this analysis all the forces are drawn relative to the axes fixed to the connecting rod centre line. The big-end bearing loads are relatively simple to calculate, whereas the main bearing loading is slightly complicated as it consists of the force partly from the reaction of the big-end bearing loading FB and partly from the out-of-balance forces of the crankshaft (crank webs + crank pin). This is an indeterminate problem because both the crankshaft and the crankcase have finite stiffness and the interactions of big-end bearings loading and crankshaft out-of-balance forces make the calculations very complex"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002173_robio.2009.5420838-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002173_robio.2009.5420838-Figure1-1.png",
        "caption": "Fig. 1 Coordinate system representation of UVMS",
        "texts": [
            " Based on this discussion, a new redundancy scheme is proposed for minimizing the vehicle restoring moment and joint torques. Section IV shows the simulation results of the proposed algorithm. Finally, Section V concludes the paper. R 978-1-4244-4775-6/09/$25.00 \u00a9 2009 IEEE. 1393 In this section, the kinematic modeling of an underwater vehicle-manipulator system is briefly reviewed. Consider an underwater vehicle with an n-link mounted arm. In mdimensional space, the general vectors, and are utilized to define the ocean vehicle state vector. Using an underwater coordinated frame as in Fig. 1, the former vector is expressed in an inertial fixed frame, i while the latter is defined in a vehicle fixed frame, v. The end-effector fixed frame is denoted by ee. The vector is represented by , where is the vehicle position vector and is the vehicle orientation vector. The vector is represented by , where is the linear velocity vector and is the angular velocity vector. Similarly, the two vectors, and ! ! ! , which consist of the endeffector position and orientation vectors, are used to define the end-effector state vector in the inertial-fixed frame and vehicle-fixed frame, respectively"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000224_1-4020-2933-0_13-Figure5-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000224_1-4020-2933-0_13-Figure5-1.png",
        "caption": "Figure 5. A spherical-spherical joint",
        "texts": [
            " Point Pi on body i and point Pj on body j are defined along the joint axis, where they must remain together. Vector is is defined along the joint axis on body i and vectors ja and jb are defined on body j perpendicular to the joint axis. The following constraints can be written: 3 5 1 1 T 1 1 T 0 0 s P P i j r n i j n i j \u23a7 \u2261 \u2212 = \u23aa\u23aa\u2261 \u2261 =\u23a8 \u23aa \u2261 =\u23aa\u23a9 r r 0 s a s b ( , ) ( , ) ( , ) ( , ) . Similarly, other types of kinematic constraints, such as cylindrical, prismatic, universal can be constructed. Another useful constraint is the spherical-spherical joint shown in Figure 5. The rigid link between the two bodies is not represented as a body\u2014it is represented as the following constraint in order to lower the number of coordinates in a model: 1 T 21 0 2 s s l\u2212 \u2261 \u2212 =d d( , ) ( ) . (8) The coefficient 1/2 is recommended in order to eliminate the 2 coefficients in the time derivatives. This constraint can be combined with Eqs. (5)\u2013(7) to model more complex joints. The first time derivative of a position constraint yields the corresponding velocity constraint. For our four fundamental constraints, the velocity constraints are: 1 1 T T 2 1 T T 3 1 T 0 s 0 0 n j i i i j j n P P i i i j j j i i i s P P i i i j j j s s P P j j j i i i \u2212 \u2261 \u2212 \u2212 = \u2261 \u2212 + \u2212 \u2212 + = \u2261 \u2212 \u2212 + = \u2261 \u2212 \u2212 + = s s s s d s r s r s r s r s 0 d r s r s ( , ) ( , ) ( , ) ( , ) ( ) ( ) , We consider an array of velocities for the two bodies containing the translational and rotational velocities for bodies i and j, in that order"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003821_ls.1184-Figure5-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003821_ls.1184-Figure5-1.png",
        "caption": "Figure 5. Schematic view of the experimental set-up for measuring material permeability.",
        "texts": [
            " In this case, on the basis of ISO 4022:1987,15 the pressure drop of the fluid passing through the porous medium is measured and the coefficients estimated. The equipment is composed of a cylindrical chamber, manufactured with a polyvinyl chloride tube with a length of 1m. The air flow in the tube is perpendicular to the sample, which is positioned inside a holder. The holder is located in the middle distance between the pressure Copyright \u00a9 2012 John Wiley & Sons, Ltd. Lubrication Science (2012) DOI: 10.1002/ls gauges fixed along the chamber, allowing measurements independently of the sample thickness. Figure 5 shows the schematic view of the experimental set-up, indicating the position of flowmetres and the electric pressure gauges. Two air filters are located in the pressure line next to the assembly, one for pre-filtering with 40mm grid and another for coalescence with 0.01mm grid. During the experimental tests, the room temperature remained around 25 C. The air density and the air dynamic viscosity were considered to be 1.079 kg/m3 and 1.83 10 5 Pa s, respectively. Measurements were made in duplicates for each experimental test to get an average value of the pressure drop and to guarantee data repeatability"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002425_vppc.2012.6422784-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002425_vppc.2012.6422784-Figure1-1.png",
        "caption": "Fig. 1 (a) shows the winding configuration of the motor under the inter turn short. Rr denotes a contact resistance of the turns short fault. The contact resistance Rr has a low resistance value. The turn short fault winding makes another fault circuit loop which is composed of the contact resistance, fault winding, and the fault back emf. Since the fault circuit loop is connected to the main circuit, the overall circuit analysis is very complicated. To simplify the circuit, we assumed that Rr is almost zero impedance. The main circuit and the fault circuit can be modeled as shown in Fig. 1 (b). Since they are not electrically connected to each other, the whole analysis is much simpler. But, they are coupled magnetically, the cross coupling effects should be considered in the analysis.",
        "texts": [],
        "surrounding_texts": [
            "back emf and an impedance unbalance. Those unbalance causes a torque ripple and an escalating partial heat defect. This paper proposes a fault detection scheme based on the PMSM dynamic model under the inter turn short which is derived in the positive and the negative sequence synchronous reference frames. To validate the proposed scheme, the finite elements method (FEM) with electric circuit is conducted on the PM SM.\nI. INTRODUCTION\nThe reliability is an issue of the vehicle propulsion system since it is directly related to the passengers safety. Hence, every propulsion system part has to be trouble-free and fault detectable. Hence, the PMSM, which is commonly used for the propulsion part of the electric vehicle (EV) and hybrid EV, also need to be reliable and easily fault detectable.\nOne common fault of the PMSM is a winding turn short fault caused by the coil insulation breakdown. Since the coil insulation material is under high voltage and temperature stress, it degrades gradually and reaches to the inter turn short fault. When the inter turn short happens, the shorted turns composes an extra circuit loop which is coupled with others motor winding flux and the rotor magnet flux. The couple flux induces the high fault current in the turn fault winding. The fault current generates heats up the shorted winding. The heat is conducted to the near winding and weakens the insulation material. Hence, the inter turn fault easily expands to the near winding.\nWhen the inter turn short occurs, the phase impedance and the back emf are reduced in the turn short phase winding. Hence, the impedance and back emf of the three phase winding are not balanced. As a result, the negative sequence terms are generated in the motor current or the voltage [1 ]-[ 4]. Note further that the normal field-oriented control yields torque ripple at the presence of negative sequence terms. In the previous works, the inter turn fault were developed mostly in the abc-frame [2]-[8]. They are complex since they include the negative and the positive sequences. One previous work has shown in the d-q reference frame [1]. The dynamic model is simple and easy to implement in the filed-oriented control scheme. But, it has focused on the healthy winding excluding the faulted inter turn loop.\nIn this paper, dynamic d-q model of the PMSM under the inter turn short fault is derived in the positive and the negative sequence rotating frame. Based on the model, a new fault detection scheme is proposed also. The proposed\n978-1-4673-0954-7/12/$31.00 \u00a920121EEE",
            "The turn number of main circuit a-phase is reduced by the fault. Hence, the motor model changes depending on the faulted turn number. To establish the main circuit model, the equivalent magnetic circuit is used.\nwhere Nshort denotes the fault turn number. By the magnetic circuit of Fig. 2(b), the flux density at the faulted phase is derived as\n(2)\nNl(%P-Ij-Nl( f -2 j - XNl IA,3,Pl2 =\nR(%P-lj +R\n(3)\nBy eq. (2) and (3), the self magnetizing inductance of the turn short phase Lmt is obtained such that\n(4)\nIf the PMSM has no fault (\ufffd= 0), the self magnetizing inductance is Lm = PN2 I 3R. But, if the PMSM has an inter turn short (1 Z \ufffd > 0), the decrease of the self magnetizing inductance has a quadratic function. Fig. 3 shows the decrease of the magnetizing inductance along the fault ratio.\nWith (4), the flux linkage of the PMSM under the a-phase winding inter turn fault is given by\nwhere Aahc = [Aa Ah Ac y, iahc = k ih ic y, \"'ahc =If/m [(1-\ufffd)cosg cos(g- 2;j cos(g + 2;jJ ,\nLn,+L[ -\ufffdL -\ufffdL 2 m 2\nm I I Lsym = -'2Lm Lm+LJ -'2Lm\n-\ufffdL -\ufffdL L L 2 2 m m+ ["
        ]
    },
    {
        "image_filename": "designv11_12_0001337_iccas.2007.4407011-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001337_iccas.2007.4407011-Figure1-1.png",
        "caption": "Fig. 1 3-D formation geometry",
        "texts": [
            " The body-fixed reference coordinate is a rotating coordinate attached to the cg position of the leader, and is aligned with its instantaneous velocity. Since the flight path of the formation flight mainly lies on a horizontal plane, it can be assumed that 0i\u03b3 \u2248 . Therefore formation flight control problem can be decomposed into two decoupled problems, which are the horizontal and vertical tracking problem. The formation geometry can be also decoupled with horizontal geometry and vertical geometry as shown in Fig. 1. The parameters of the horizontal formation geometry are the forward clearance dx and lateral clearance dy from the reference frame of the leader. The forward distance error xe and lateral distance error ye in an inertial, Earth-fixed frame can be expressed as 1 2 1 1 1 2 1 1 x d y d e xX X C S e Y Y S C y \u03c7 \u03c7 \u03c7 \u03c7  \u2212 \u2212      = \u2212       \u2212         (12) where C and S denote the cosine and sine of the related angles, respectively, and 1\u03c7 represents the heading angle for the leader aircraft. The vertical distance error he can be represented as 1 2h de H H h= \u2212 \u2212 (13) where dh is the vertical distance clearance that is the desired height difference as shown in Fig. 1. The specific energy error Ee can be obtained as 2E de E E= \u2212 (14) where 1 1( , , )d dE f V H h= is a desired energy state. The general objective of formation control is to minimize the distance errors , ,x y he e e using the throttle, yaw acceleration, pitch acceleration command. That is, the throttle and yaw acceleration commands are used to control the forward distance and lateral distance, and the pitch acceleration command is used to control the vertical distance. In this study, to obtain a fuel efficient controller the pitch acceleration command is used to control the forward distance using the energy maneuverability method"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001487_j.rehab.2009.12.001-Figure3-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001487_j.rehab.2009.12.001-Figure3-1.png",
        "caption": "Fig. 3. Syste\u0300me de freinage pour le prototype. a : poigne\u0301es de frein fixe\u0301es sur les poigne\u0301es de bras de pousse\u0301e ; b : les e\u0301triers de frein fixe\u0301s sur le cha\u0302ssis derrie\u0300re la roue arrie\u0300re.",
        "texts": [
            " Handling ability (related to the prototype-user interface) and other parameters were evaluated with a user self-questionnaire. Data was collected after 2 days of practice on the new prototype. Our prototype is an adjusted non-folding, lightweight Quickie GT1, integrating a system of drive chains and freewheels linked to a two-handed, push bars placed in front of the user. The push-bars are linked to the wheelchair\u2019s rear wheel axles on each side (Fig. 1 and 2). Brake calipers are fitted on the chassis behind the back wheel and can be activated by brake levers fitted onto the push-bars (Fig. 3). None of the other mechanical components of the conventional wheelchair were modified. The prototype was a subject of a publication elsewhere [23]. The objectives are described as follows: to identify the prototype\u2019s advantages and disadvantages in different conditions; help us in improving the prototype\u2019s design; to evaluate the user\u2019s overall satisfaction with the prototype\u2019s push-bar propulsion system. Seventeen volunteers (11 men and six women) has participated in the study. Participant\u2019s examination has confirmed the absence of any shoulder problems and has all a minimum of 2 years experience in sport wheelchair",
            " Notre prototype a e\u0301te\u0301 conc\u0327u a\u0300 l\u2019aide d\u2019un syste\u0300me de cha\u0131\u0302nes de bicyclette relie\u0301 aux deux poigne\u0301es de pousse\u0301e. Ces deux poigne\u0301es de pousse\u0301e servent comme des leviers a\u0300 la porte\u0301e de l\u2019utilisateur qui les prend par ses deux mains. Ces leviers ont e\u0301te\u0301 fixe\u0301s et relie\u0301s aux deux axes des roues-arrie\u0300re principales du fauteuil roulant (Fig. 1 et 2). Des e\u0301triers de frein sont monte\u0301s sur le cha\u0302ssis derrie\u0300re la roue arrie\u0300re et active\u0301s par des poigne\u0301es de frein monte\u0301es sur les poigne\u0301es de pousse\u0301e (Fig. 3). Les accessoires et les composantes me\u0301caniques du fauteuil, type Quickie GT1 : ultrale\u0301ger et non-pliable, sont reste\u0301s les me\u0302mes. Notre prototype a e\u0301te\u0301 l\u2019objet d\u2019une e\u0301tude expe\u0301rimentale publie\u0301e [23]. Les objectifs sont de\u0301crits comme suit : identifier les points forts et les points faibles du prototype pour les diffe\u0301rents terrains et dans les diffe\u0301rentes conditions d\u2019utilisation. solliciter la participation des utilisateurs de\u0301pendants du fauteuil roulant pour contribuer a\u0300 l\u2019ame\u0301lioration du prototype"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000137_la0624945-Figure5-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000137_la0624945-Figure5-1.png",
        "caption": "Figure 5. Schematic illustration of the response (change in size) of doubly thermosensitive core-shell microspheres to changes (increases) in temperature.",
        "texts": [
            " The narrow size distributions of all the particles also suggest that the nucleation of new particles during shell addition is negligible. Should any linear polymers have been formed during seed polymerization, they may have entangled with the crosslinked matrix of the microspheres to form the core-shell structure, or they were removed during the purification process. Temperature-dependent shrinking behaviors were studied by means of optical transmittance (Figure 2B) and DLS (Figure 4B), indicating quite clearly a two-step shrinkage upon rising the temperature as depicted in Figure 5. The first transition corresponds to the volume phase transition (VPT) of the thermosensitive outer layer of P(nPA-co-S) near 13-15 \u00b0C, while the second transition is attributed to the VPT of PDEA, PiPA, or PiPMA shells near 28, 32, or 42 \u00b0C, respectively, as illustrated by Figure 5. The thicknesses of the shells on PnPA50S50 seed particles were calculated to be 180, 235, and 175 nm for PDEA, PiPA, and PiPMA respectively, whereas the thicknesses of the shells on PnPA25S75 were estimated to be 275, 350, and 205 nm for PDEA, PiPA, and PiPMA shells, respectively. Furthermore, Figure 4B shows that the deswelling extent of the coreshell microspheres at the first VPTT compares well with that of the parent core particles. For clarity of presentation, only the core-shell particles with a PnPA25S75 core are shown in Figures 2B and 4B, while other microspheres with different chemical compositions behaved similarly",
            " Such interpenetration could be the reason that the VPTT of PnPA outer layer of the core microspheres was shifted to higher temperatures, which is more remarkable in the case of PnPA25S75-iPMA (Figure 4B) and PnPA50S50-iPMA (see Supporting Information). The extent of interpenetration is expected to change depending on the core particles used and the nature of the thermosensitive monomer shell. The swelling ratios of core-shell microspheres were calculated according to eq 5 and shown in Figure 4C. Water expelled from the microspheres upon shrinking results in 80-90 vol % loss, compared to the fully swollen state. The two-step deswelling process depicted in Figure 5 is observed in Figure 4C with each step corresponding to the deswelling properties of the corresponding homopolymers. It shows that the PiPA and PiPMA R ) Vshrunk Vswollen ) (R h 313.2K Rh 283.2K)3 (5) shells of core-shell microspheres, which have larger volumes of expelled water (>90 vol %), shrink more efficiently than others. Doubly thermosensitive core-shell microspheres, which have a core of copolymer of nPA with styrene and a shell of N-alkyl (meth)acrylamide, can be prepared by a one-pot synthetic procedure"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003076_cphc.201300600-Figure5-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003076_cphc.201300600-Figure5-1.png",
        "caption": "Figure 5. a) Normal bending of a rod of length l = 1, and with different modes: b) G is the center of inertia and Gx is the neutral axis of the cross-sectional area, S. c) For Mz>0, the stress sxx<0 for y>0 (compression), and sxx>0 for y<0 (tension). d) Stress sxx and strain tensor exx components. Adapted from ref.[29] .",
        "texts": [
            " The deformation is assumed to be small, and linear isotropic elasticity described the passive solid. We present two characteristic cases that highlight the key aspects of the device: 1) normal bending of a rod (moment axis is perpendicular to the axis of the rod), and 2) off-axis bending of a beam with a rectangular cross-section; the solid mechanics follows the treatment of [29] . A cylindrical rod of length 1, with its axis along the x direction, is subjected to a bending moment of M(x=1) M(x=0) = Mzdz, see Figure 5 a. Normal bending denotes the perpendicular tilting of the areal cross-section with respect to the curvature of the axis of the rod. The stress vectors, Td y \u00bc Td z \u00bc 0, at x = 0 and x = 1 are zero. The origin of the axis is the point \u201cO\u201d, which is located at the center of inertia of the cross-section at 2014 Wiley-VCH Verlag GmbH & Co. KGaA, Weinheim ChemPhysChem 2014, 15, 1405 \u2013 1412 1409 x = 0. The membrane contact area, A, is much smaller than the lateral area of the rod; therefore, the effect of curvature of the cross-section on the glued flexoelectric membrane can be neglected. The elastic modulus of the rod is E, the Poisson ration is n, and its principal moment of inertia is Iz. The uniaxial stress-tensor field s(y) and strain-tensor field, the components of which are illustrated in Figure 5 b\u2013d, can be represented by Equation (22): s y\u00f0 \u00de \u00bc M Iz ydxdx; e y\u00f0 \u00de \u00bc Mz Iz y dxdx n dydy \u00fe dzdz \u00f022\u00de s(y) is independent of the cross-section and is a linear function of y, and it creates compression above the neutral axis x and tension below it (Figure 5 c). The curve along the centers of inertia of the cross-sections is the mean fiber of the rod with a curvature Czx given by Equation (23): Cs z \u00bc Mz EIz \u00f023\u00de The constant curvature describes the circular bending of all of the lines parallel to the Ox axis in circular arcs that lie in planes parallel to the Oxy plane. The curvature vector, w, describes the tilting of the cross-section attributed to bending [Eq. (24)]: w x\u00f0 \u00de \u00bc w x\u00f0 \u00dedz \u00bc Cs z xdz \u00f024\u00de w is normal to the mean fiber and coincides with the z-axis in Figure 5 c; therefore, it divides each cross-section into uppercompression and lower-tension zones. In terms of curvature, s(y) follows Equation (25): s y\u00f0 \u00de \u00bc Cs z Eydxdx \u00f025\u00de By placing the flexoelectric membrane on the top, y = h and centered at z = 0, the surface stress in the solid is given by Equation (26): sxx y \u00bc h\u00f0 \u00de \u00bc Cs z Eh \u00f026\u00de and because the flexoelectric membrane is glued to the beam, their curvatures are equal, as shown in Equation (27): Cm z \u00bc Cs z \u00bc sxx Eh \u00f027\u00de By using the relationship between bending stress, sxx, and displacement (D = 2 eH) we obtain Equation (28): D \u00bc e Eh sxx \u00bc k ffiffiffiffiffiffiffiffiffiffi ebYm p Eh "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003713_978-1-4614-2419-2-Figure2.4-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003713_978-1-4614-2419-2-Figure2.4-1.png",
        "caption": "Fig. 2.4 Schematic of the experimental setup",
        "texts": [
            " For the second measurement, two pairs of chronographs were used to measure the projectile speed before and after impact; one pair was located in front of the target, while the second was located behind the target. 8 A. Seyed Yaghoubi and B. Liaw 2 Experimental and Numerical Approaches on Behavior of GLARE 5 Beams. . . 9 Thirdly, a high-speed camera was used to monitor the bullet motion during the test. In this study, the high-speed camera was set perpendicular to the projectile\u2019s ballistic trajectory. Using the captured high-speed video, the bullet speed was then determined before and after the impact. A schematic of the setup is shown in Fig. 2.4. After determining the speeds of the bullet before and after impact, the incident projectile impact velocity was plotted versus the residual velocity for each type of the specimens listed in Table 2.1. The experimental data was then fitted by least-square regression according to the classical Lambert-Jonas equation [10] for the positive residual velocity values: VP R \u00bc A VP I VP 50 \u00bc A VP I B (2.2) where A and B are two regression coefficients and P is power. VR and VI are the residual and incident velocities of the projectile, respectively, while V50 is the ballistic limit velocity, which is defined as the velocity required for a projectle to perforate the target 50% of the time"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003501_2041302510394742-Figure5-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003501_2041302510394742-Figure5-1.png",
        "caption": "Fig. 5 Structure of thrust and attitude controls",
        "texts": [
            " The advantage of the rotating wing is that transition to stall can be delayed, and the co-axial counter-rotating propellers can provide lift force to keep the height invariant although the wing will still stall as forward velocity goes to zero. The pair of co-axial counter-rotating propellers provides up thrust when the aircraft takes off or lands, and forward thrust when the aircraft flies forward. Moreover, the counter-rotating torques of the two propellers can be counteracted to each other. In Fig. 5, directional control includes: roll, pitch, and yaw controls. The co-axial counter-rotating propellers 1 are driven by an engine 10. The tail counterrotating propeller 9 is driven by motor 11. There exist four groups of pairings 5\u20138 generating pitch and yaw moments during modes transition. The tail counterrotating propeller 9 with the pairings provide pitch moments and yaw moments during mode transition. Moreover, in hover, roll dynamics can be controlled Proc. IMechE Vol. 225 Part G: J. Aerospace Engineering at The University of Manchester Library on April 23, 2015pig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001960_s12239-009-0053-x-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001960_s12239-009-0053-x-Figure2-1.png",
        "caption": "Figure 2. 3-D gear train model of a manual transmission.",
        "texts": [
            " The reactive forces on teeth and bearings were calculated and compared using three different methods developed for this study - an equivalent model, a rigid-body model, and a frequency-based model. A realistic multibody dynamic model for a transmission system, which may reflect real working conditions, should be constructed for an accurate loading analysis. An FF manual transmission consists of a clutch, input and main shafts, mating helical gears, final-drive gears in the differential section, and housing. The 3-D six-speed manual transaxle model, which combines the manual transmission, final drive gearing, and differential into a single unit, is shown in Figure 2. A multibody dynamic analysis model for Figure 2 was constructed using MSC/ADAMS and is represented in Figure 3. This model is based on the following three assumptions: (1) shafts and gear teeth are flexible, and bearings are considered to be bushings with 6 degrees of freedom; (2) gear meshing stiffness due to bending varies along a moving contact point between two helical teeth; and (3) fluctuating torque or acceleration is transmitted through the clutch input to the transmission input shaft. 2.1. Bending Stiffness of a Gear Tooth Figure 4 shows a schematic of three components of forces acting against a helical gear tooth"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000710_j.aca.2007.03.052-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000710_j.aca.2007.03.052-Figure1-1.png",
        "caption": "Fig. 1. Scheme of the PMME.",
        "texts": [
            " pH value of solution as measured using a Mettler Toledo DELTA 320 meter Mettler-Toledo, Shanghai, China). The HPLC system (Agient Technologies, Boblingen, Germany) consisted of Agilent 100 Series components including a quaternary pump, degasser, heodyne 7725i six-port valve with a 20 L loop (Cotati, A, USA) and a fluorescence detector. Fluorescence detection avelengths were set at \u03bbex/\u03bbem = 498/507 nm. Chromatoraphic separations were achieved with ZORBAX Eclipse DB-C18 column (150 mm \u00d7 4.6 mm, 5 m, Agilent Technoloies, Boblingen, Germany). The PMME extraction device as shown in Fig. 1 was escribed in detail in the previous work [23]. The poly(MAAGDMA) monolithic column was formed inside a fused silica apillary by a polymerization method that has been described reviously [19,20]. The monolithic capillary tube used in this 1 Chim 2 o b o P ( w d u a t P 2 P l 0 t 0 t F a i p a e m t a 2 f c p t c fl c a c 2 p p s p h i t a 3 3 m O s a z s t o T e r t s y r a a b d d c t e s u t n a t z c w a d t o t t d t a c 18 K.-J. Huang et al. / Analytica .3. Derivatization procedure Some components in the blood may influence the reaction f DAMBO with NO"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003761_9781118818428-Figure7.1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003761_9781118818428-Figure7.1-1.png",
        "caption": "Figure 7.1 The Archibald Test \u2013 schematic representation",
        "texts": [
            " 87), differ little from the Schmidt when both are computed for common values of the respective parameters \u2013 temperature ratio, volume ratio, volume phase angle, and dead-space ratio(s). Either (a) it does not matter which of the two extreme values is chosen for h or (b) extant formulations of the \u2018adiabatic\u2019 model do not fully reflect the implications of h = zero. If either extreme were correct then the thermal lag engine could not function, so reality must lie with some intermediate value. A simple bench test sheds light out of all proportion to the experimental overhead. Figure 7.1 indicates a sealed cylinder, some 50 mm (2 inches) diameter and 75 mm (3 inches) long, containing air and attached to a simple Bourdon-tube pressure gauge. With container and contents pre-cooled to zero deg. C, pressure is atmospheric and the gauge indicates zero. The container is heated to, say, 400 \u25e6C, plunged into a water/ice mix and the time for air temperature T to return to the start value noted, as indicated by the fall in pressure. Since John Archibald notified the experiment by e-mail, the author has described it on numerous occasions and challenged hearers to estimate temperature decay time"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001631_1.3442472-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001631_1.3442472-Figure1-1.png",
        "caption": "Fig. 1. According to our choice of the initial condition, the pendulum starts its motion from =0 with a nonzero initial velocity v0, directed from the left to the right.",
        "texts": [
            "1 This approximate solution works well for 7\u00b0, for which T0 gives an error less than 0.1% compared to the exact period.2\u20134 For larger amplitudes the nonlinear nature of the pendulum oscillations becomes apparent. Although the sinusoidal solution remains a good approximation to the exact solution of the nonlinear equation of motion for small amplitudes, the period increases rapidly with the amplitude.5 Although the oscillatory motion is more common, there are two other possibilities. When the pendulum starts its motion from =0 with an initial velocity v0 0, as indicated in Fig. 1, one of the following three regimes is observed.6 1 The oscillatory regime occurs when the total energy E is less than 2mg and is thus insufficient for the bob to reach the top position the point P . 2 Perpetual ascent occurs when the pendulum has exactly the energy needed to reach the top, that is, E=2mg . This energy corresponds to an initial velocity v0=2 g , which we call the critical velocity vc for the transition from the oscillatory to the nonoscillatory regime. 3 The nonoscillatory regime occurs when E 2mg , a case in which the pendulum rotates periodically without velocity inversions",
            " An exact analytic solution, as well as an exact expression for the period, both in terms of elliptic integrals and functions, is known only for the oscillatory 1146\u00a9 2010 American Association of Physics Teachers ject to the terms at: http://scitation.aip.org/termsconditions. Downloaded to IP: 5 Apr 2014 15:26:41 This art \u2018regime.8,11 To derive the exact analytic solution of Eq. 1 for the nonoscillatory regime, we use the conservation of energy approach in Refs. 2, 5, and 11. By taking the lowest point of the circular trajectory as our reference level for the potential energy, the total energy a constant of motion reads E=mv2 /2+mgh, where v is the speed of the bob and h is its height see Fig. 1 . Because h = 1\u2212cos and v= \u0307, where \u0307 d /dt, we have E = 1 2m\u03072 2 + mg 1 \u2212 cos . 2 By solving for \u03072 and using the identity sin2 /2= 1 \u2212cos /2, we find \u03072 = 2E m 2 \u2212 4 g sin2 2 . 3 Equation 3 is valid for all . For simplicity, we adopt the initial condition 0 =0 and \u0307 0 =v0 / .13,15 The counterclockwise direction of motion will be taken as positive, as indicated in Fig. 1. In this regime 0 E 2mg , and the bob reaches a maximum height corresponding to the amplitude max 0 max , stops, returns to its lowest position, and repeats this motion symmetrically for negative values of . By applying Eq. 3 to the highest position, where \u0307=0, we have 2E m 2 = 4 g sin2 max 2 . 4 Because the potential energy is zero at =0, E=mv0 2 /2. Therefore, v0 = 2 g sin max 2 . 5 From Eq. 3 , we find 1147 Am. J. Phys., Vol. 78, No. 11, November 2010 icle is copyrighted as indicated in the article"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003415_j.mechmachtheory.2011.03.006-Figure5-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003415_j.mechmachtheory.2011.03.006-Figure5-1.png",
        "caption": "Fig. 5. The construction of a 6R linkage from two subtractive Goldberg 5R linkages by removing the common Bennett linkage in dash lines.",
        "texts": [
            " (3), we can use link-pair 45\u201351 to form another common Bennett linkage, which is different from Bennett linkage A or B. For two subtractive Goldberg 5R linkages with the following geometric parameters, and aL12 = aL34 = a;aL23 = d\u2212c;aL45 = c;aL51 = d; \u03b1L 12 = \u03b1L 34 = \u03b1;\u03b1L 23 = \u03b4\u2212\u03b3;\u03b1L 45 = \u03b3;\u03b1L 51 = \u03b4; aR12 = aR34 = b;aR23 = c\u2212a; aR45 = a;aR51 = c; \u03b1R 12 = \u03b1R 34 = \u03b2;\u03b1R 23 = \u03b3\u2212\u03b1;\u03b1R 45 = \u03b1;\u03b1R 51 = \u03b3; \u00f010\u00de k-pair 34\u201345 of linkage L and link-pair 45\u201351 of linkage R are the same as a/\u03b1 and c/\u03b3. Therefore, a 6R linkage can be the lin obtained by constructing a common Bennett linkage from these two link-pairs, see in Fig. 5. Thus, the geometry conditions of this 6R linkage are a12 = a34 = b;a23 = c\u2212a;a45 = d;a56 = a;a61 = d\u2212c; \u03b112 = \u03b134 = \u03b2;\u03b123 = \u03b3\u2212\u03b1;\u03b145 = \u03b4;\u03b156 = \u03b1;\u03b161 = \u03b4\u2212\u03b3; sin\u03b1 a = sin\u03b2 b = sin\u03b3 c = sin\u03b4 d ; Ri = 0 i = 1; 2;\u2026; 6\u00f0 \u00de: \u00f011\u00de According to Eq. (9), the closure equations of the subtractive Goldberg 5R linkages L and R are \u03b8L1 + \u03b8L4 = 2\u03c0;\u03b8L2 + \u03b8L3 + \u03b8L5 = 2\u03c0; tan \u03b8L1 2 tan \u03b8L2 2 = sin \u03b4 + \u03b1 2 sin \u03b4\u2212\u03b1 2 = m1; tan \u03b8L1 2 tan \u03b8L3 2 = sin \u03b3 + \u03b1 2 sin \u03b3\u2212\u03b1 2 = m2; \u00f012\u00de \u03b8R1 + \u03b8R4 = 2\u03c0;\u03b8R2 + \u03b8R3 + \u03b8R5 = 2\u03c0; tan \u03b8R1 2 tan \u03b8R2 2 = sin \u03b3 + \u03b2 2 sin \u03b3\u2212\u03b2 2 = m3; tan \u03b8R1 2 tan \u03b8R3 2 = sin \u03b1 + \u03b2 2 sin \u03b1\u2212\u03b2 2 = m4; \u00f013\u00de tively. respec Here mi(i=1, 2, 3 and 4) are set to simplify the representations of different relationships of twists for later derivation. The relationship between the revolute variables of the objective 6R linkage and the subtractive Goldberg 5R linkages L and R are \u03b81 = \u03b8R1\u2212\u03c0;\u03b82 = \u03b8R2 ;\u03b83 = \u03b8R3 ; \u03b84 = \u03b8L3\u2212\u03c0 + \u03b8L5\u2212\u03c0 \u2212 \u03c0\u2212\u03b8R4 = \u03b8L3 + \u03b8L5 + \u03b8R4\u22123\u03c0; \u03b85 = \u03b8L1;\u03b86 = \u03b8L2: \u00f014\u00de For the common Bennett linkage shared by these two subtractive Goldberg 5R linkages, as shown in dash lines in Fig. 5, the compatibility relationship, should \u03b8L4 + \u03b8R5 = 2\u03c0; \u00f015\u00de be preserved. show that the newly formed 6R linkage has mobility one. Because of the construct method, it can be called as the double which subtractive Goldberg 6R linkage. Physical models are also made to validate this linkage. Fig. 6 is the constructing process of the double subtractive Goldberg 6R linkage shown in physical models. Fig. 7 shows the full circlemovement of the constructed linkage, whose input\u2013output curves are presented in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001763_robot.2009.5152405-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001763_robot.2009.5152405-Figure1-1.png",
        "caption": "Fig. 1. (a) Experimental field. (b) Modified e-Puck with RFID writer/reader. (c) Data carrier.",
        "texts": [
            " 3 REALIZATION OF AN ARTIFICIAL PHEROMONE SYSTEM BY USING DATA CARRIERS Advancements in communication technologies have provided many advantages to researchers, particularly in the field of robotics. Radio frequency identification (RFID) technology is used in our research. This technology uses radio frequency or magnetic field variations for communication. (2) (3) (4) The two components of RFID system are tags and readers. A tag is an identification device attached to an item that is to be identified. A reader is a device that can recognize the presence of RFID tags and read the information stored in them [17]. We use a tag as a data carrier (shown in Figure 1(c)). This data carrier is wireless and battery-free. Each data carrier is marked with a unique identifier and is equipped with a tiny memory that allows it to store data. This tag has a 32-bit unique ID that is factory-programmed and an 896-bit user memory that is organized into 28 blocks of 32-bits each. Data carriers can be used to store and retrieve information several times. Multiple data carriers can be read simultaneously. A reader is attached to each autonomous robot. In this research, data carriers are used to store information during the construction of pheromone potential fields",
            " We use passive data carriers; hence, the storing process occurs only if robot communicates with the data carrier. The mobile robots used in this project are based on the e-Puck (mini mobile robot) that was originally developed for educational purpose at Swiss Federal Institute of Technology in Lausanne (EPFL). An RFID reader/writer is added to the e-Puck for the reading and writing processes. An antenna is employed for this project. Serial communication is used between the e-Puck and RFID processing unit in order to control the reading and writing processes. Figure 1(b) shows an image of a modified e-Puck used in this experiment. The robot moves around the data carriers in the experimental field, as shown in Figure 1(a). The unique identification (UID) of the data carrier is read, and it is verified whether or not the tag is new when the data carrier is detected. If the robot finds that the tag is new, the robot reads data from the tag. The robot calculates the current density from the information stored inside the data carrier and the latest information obtained from the robot. Subsequently, the robot stores the current density information into the current data carrier. The autonomous robots read and write on the data carriers and construct the potential field of the pheromone by using the content of each data carrier in the experimental field",
            " 5(b) to guarantee the stability of the system. From Fig. 6, we can see that the system is stable and goes on to a steady state if we use the parameters from Fig. 4 that follows 1 1 )(2 = \u2212 + \u03be \u03b7\u03b6 . 5 EXPERIMENT By using an artificial pheromone system composed of data carriers and autonomous robots, as explained in section 3, we conducted an experiment to verify our model by using real hardware. In the experiment, we used data carriers in a random manner. We used one autonomous robot and 25 data carriers for this purpose. Figure 1(a) shows the experimental field employed to realize an artificial pheromone system. In the experiment, the autonomous robot moved randomly from one data carrier to the next data carrier. The robot stored information obtained from the previous data carrier and stored them in the current data carrier. At the beginning of the experiment, there was no information on the density of neighboring data carriers stored in each data carrier. Information on the density of neighboring data carriers was updated after every time step when the robot communicated with the data carrier"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001321_tsmcc.2007.900660-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001321_tsmcc.2007.900660-Figure1-1.png",
        "caption": "Fig. 1. Two inverted pendulums connected by a spring.",
        "texts": [
            " However, \u2016ri(0)\u2016 can be set to zero by appropriately initializing the reference trajectory qid(0). 2) Large adaptation gain lwi, l\u03b4i, ldi, which attenuate the effects of initial parameter error is vital for good L\u221e performance. 3) Large parameter \u03bbs can also lead to improvement in performance. This implies that increasing Kvi can achieve small tracking error. V. SIMULATION EXAMPLE In this section, we illustrate the proposed decentralized RBF control by means of a practical example of inverted double pendulums connected by a spring as shown in Fig.1. The equations that describe the motion of the pendulums are defined by \u03b8\u03081 = ( m1gr J1 \u2212 kr2 4J1 ) sin(\u03b81)+ kr 2J1 (l\u2212b)+ u1 J1 + kr2 4J1 sin(\u03b82) (42) \u03b8\u03082 = ( m2gr J2 \u2212 kr2 4J2 ) sin(\u03b82)\u2212 kr 2J2 (l\u2212b)+ u2 J2 + kr2 4J2 sin(\u03b81) (43) where \u03b81 and \u03b82 are the angular displacements of the pendulums from vertical. This system can be represented in a form of (1) with Mi = Ji Vmi = 0 Gi = ( migr \u2212 kr2 4 ) sin(\u03b8i) Fi = 0 \u03c4di = (\u22121)i+1 kr 2 (l \u2212 b) Zi = kr2 4 sin(\u03b8i). It can be seen that the disturbance and interconnections satisfy the properties"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000894_iros.2008.4651029-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000894_iros.2008.4651029-Figure2-1.png",
        "caption": "Fig. 2. Two possible assembly configurations of the concerning device.",
        "texts": [
            " Two different aspects of this property can be underlined: the flexure, in the hinges chosen as joints, assure a functional compliance, while a collateral compliance, that allows the compensation of high and unexpected forces and helps the motion, can be identified from the global architecture flexibility Strong points of this architecture are: the possibility to realize also wide motions, the positioning accuracy obtained through more close kinematic chains, and the structure capacity of fronting the arise of resistant and also finite forces while the platform remains always parallel to itself. T 978-1-4244-2058-2/08/$25.00 \u00a92008 IEEE. 735 Further essential elements are the high programmability concerning with trajectories and the function softness assured by the motion profiles at the actuators imposable. Once coupled the manipulator to a common cloth-glove (figure 2) joined to the mobile platform of the robot, through controlling the meso-manipulator, it is possible to drive the patient\u2019s finger in the required movement, to impose external loads to the involved tendons and to define those external forces as time constant, following a desired profile or a random one not overtaking pre-definable limits. It can be thought to insert force or displacement sensors between the platform center and the frame center, through spherical connection joints or similar configuration so that, during the progress of a rehabilitation treatment, repeated measurements of the forces acting on the robot can be executed"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001730_elan.200900540-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001730_elan.200900540-Figure1-1.png",
        "caption": "Fig. 1. Scheme of the assembly and operation of the GODBOD-MWCNT-Au electrode. By switching on the catalytic activity of GOD in the presence of glucose enhanced oxygen consumption depletes the concentration near the electrode surface. Thus bioelectrocatalytic oxygen reduction is decreased resulting in a diminished cathodic current.",
        "texts": [
            " For the amperometric experiments the potentiostat CHI800 (USA) was employed. A magnetic stirrer with a rotation speed of 2500 rpm was used to achieve a constant diffusion layer thickness. For the construction of the bienzyme sensor a BODMWCNT-electrode is combined with an oxygen-consuming enzyme. The oxidase recognises and converts the analyte which is accompanied by oxygen conversion. This oxygen depletion is then detected by the BOD-electrode via biocatalytic oxygen reduction and direct electron transfer (Fig. 1). For the sensitive detection of an analyte the oxidoreductase has to consume the oxygen near the BODelectrode effectively. In order to construct such a molecularly defined system the enzymes are covalently linked to the BOD which is covalently fixed to the MWCNTs. 1582 www.electroanalysis.wiley-vch.de 2010 Wiley-VCH Verlag GmbH & Co. KGaA, Weinheim Electroanalysis 2010, 22, No. 14, 1581 \u2013 1585 After preparation of the bienzyme sensor by coupling GOD to the BOD-electrode the sensor is examined by linear sweep voltammetry and amperometry. The voltammetric examination reveals that the catalytic oxygen reduction is dependent on the glucose concentration and thus GOD activity on the surface can be shown (Fig. 2). The presence of catalytically active GOD has a pronounced influence on the oxygen reduction current in a wide potential range from 0.45 V to 0.2 V. This demonstrates that the principle illustrated schematically in Figure 1 can be realised by this assembly on the electrode. GOD can obviously successful compete for the oxygen conversion at the BOD. Furthermore there is no shift of the start potential for oxygen reduction by crosslinking of GOD to the BOD-MWCNTAu electrode. The start potential for both BOD-MWCNTAu electrodes with and without GOD is about \u00fe0.5 V vs. Ag/AgCl, 1 M KCl. The differential voltammograms with and without glucose show that the response to glucose is not dependent on the electrode potential around zero polarisation"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000388_14644193jmbd97-Figure5-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000388_14644193jmbd97-Figure5-1.png",
        "caption": "Fig. 5 Loci groove centres of curvature after three dimensional displacement of the inner ring centre (x, y, and z)",
        "texts": [
            "comDownloaded from inner and outer ring raceway grooves at ith ball before and after the preload is equal to the contact deflection and can be found from Fig. 3 and equation (4), where Fig. 4 is a simplified two-dimensional representation of the view normal to a plane at the angle ui from the x-axis. d0 \u00bc Bdb cos\u00f0a0\u00de cos\u00f0aP\u00de 1 \u00f010\u00de The preloaded contact angle was obtained from the following formula aP \u00bc tan 1 Bdb sin\u00f0a0\u00de \u00fe z0 Bdb cos\u00f0a0\u00de \u00f011\u00de The actual deflections caused by the vibrations in the x, y and z directions will be dealt with from this point onwards. Figure 5 describes the situation after this three-dimensional movement of the inner ring centre has taken place. The deflections along x and y axes can always be combined as a single radial deflection. The radial deflection for the ith ball is given by (Fig. 5) dr \u00bc x cos\u00f0ui\u00de \u00fe y sin\u00f0ui\u00de \u00f012\u00de From Fig. 6 the deflection is given by d0i \u00bc \u00bd\u00f0Bdb sin\u00f0a0\u00de \u00fe z0 \u00fe z\u00de2 \u00fe \u00f0Bdb cos\u00f0a0\u00de \u00fe dr\u00de 2 1=2 Bdb \u00f013\u00de Proc. IMechE Vol. 222 Part K: J. Multi-body Dynamics JMBD97 # IMechE 2008 at UNIV PRINCE EDWARD ISLAND on August 5, 2015pik.sagepub.comDownloaded from and the contact angle is obtained from a0 i \u00bc tan 1 Bdb sin\u00f0a0\u00de \u00fe z0 \u00fe z Bdb cos\u00f0a0\u00de \u00fe dr \u00f014\u00de Deflection of the inner ring centre due to the rocking motion of the spindle may also be considered in terms of deflections in the three axes"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001461_s0022-0728(72)80262-6-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001461_s0022-0728(72)80262-6-Figure1-1.png",
        "caption": "Fig. 1. Equi l ibr ium concns, of various species vs. coord inat ion number for various init ia] N H 3 concns. ]=or a]] solns, ini t ial N i 2+ concn, is 10 -2 M. (a) 1 M NH4C], 0.4] M N H 3 ; (b) 0.5 M NH4CI , 0.205 M N H 3 ; (c) 0.1 M NH4C], 0.041 M NH3.",
        "texts": [
            " Before the calculation could be performed it was necessary to have some knowledge of the equilibrium N H 3 + H + ~ N H +. The value of K Nn3 N H 4 = [NH2 ] / [H + ] [NH 3] was taken to be log (KN H3/1 mol - 1) = 9.413 for an ionic strength of 0.1 mol 1-1 at 20 \u00b0 C 9. It was also necessary to measure the pH of each solution used. With the use of the above data, simple matrix algebra produced a 7 x 7 determinant which on evaluation gave values for the equilibrium concentration of each species present. This calculation was performed using the IBM 360 computer. Figure 1 illustrates how the equil ibrium concentrations of the various nickel species present depends on the concentration of ammonia used to prepare the solution. In order to attempt an understanding of the kinetics of the system it is assumed that the equilibria present in the solution are of the form\" MX 2+ ~ MX 2+ ~ MX~ + ~ M X 2 + ~ M 2+ . . . . . . . . u-1 . . . . (1) where in the present case n = 6. The greater the number of ammine ligands, the more labile the complex is likely to be t\u00b0. Thus, it may be seen that, providing the complex species are in sufficiently rapid equilibria, the effective concentration of the species being reduced will be E~_ 1 2 + MXj "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003501_2041302510394742-Figure6-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003501_2041302510394742-Figure6-1.png",
        "caption": "Fig. 6 Gearing arrangement of pitch control for the free wings",
        "texts": [
            " The tail counterrotating propeller 9 is driven by motor 11. There exist four groups of pairings 5\u20138 generating pitch and yaw moments during modes transition. The tail counterrotating propeller 9 with the pairings provide pitch moments and yaw moments during mode transition. Moreover, in hover, roll dynamics can be controlled Proc. IMechE Vol. 225 Part G: J. Aerospace Engineering at The University of Manchester Library on April 23, 2015pig.sagepub.comDownloaded from Xinhua Wang and Hai Lin 17 15 2 2 1 16 15 16 4 3 18 In Fig. 6, a pair of threaded screw rods 15 is, respectively, within the fixed-wing root sections 3 of the fuselage 4 in the threaded contact with a pair of threaded collars 16. Screw rods 16 extend longitudinally within the left- and right-fixed wing sections, respectively, and support in an axially fixed manner at the opposite ends thereof. These screw rods 16 are rotated about their longitudinal axes to longitudinally translate the thread collars 16 along their respective rods. A link 17 having opposite ends pivotally mounted to the associated collar 16 and free wing 2 transmits longitudinal movement of the collars to pivot free wing 2 relative to fuselage 4"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001517_j.mechmachtheory.2009.12.004-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001517_j.mechmachtheory.2009.12.004-Figure2-1.png",
        "caption": "Fig. 2. A Watt six-bar linkage.",
        "texts": [
            " A sub-branch refers to a linkage configuration space, in which transformation between configurations may be accomplished without reaching a singularity or dead center position, where the linkage may lose control [10\u201312]. In a sub-branch, an input value corresponds to one and only one linkage configuration and the order of motion of a linkage is determined by the magnitude order of the input value [12]. A Stephenson six-bar linkage, as shown in Fig. 1, consists of a four-bar loop ABCDA and a five-bar loop ABEFGA. A Watt six-bar linkage, as shown in Fig. 2, consists of two four-bar loops. A Watt six-bar linkage may also be regarded as one comprised of a four-bar loop ABCDA and a six-bar loop ABCEFGA or even a five-bar loop by the stretch and rotation of the second four-bar loop [11]. Two other six-bar linkages are shown in Figs. 3 and 4. In the following discussion, any link in the linkage may be used as the reference link. Hence, the treatment is valid for any linkage inversion. The formation of the mobility of a multiloop linkage is governed by the individual single loops and also the interaction between them",
            " (23) with s3 as the unknown parameter is D3 \u00bc sin2 c: \u00f024\u00de Since sin2 c P 0 is always true, Eq. (24) can be automatically satisfied. If sin c \u2013 0, for each pair of h2 and h3 values, there is only one solution for s3 in Eq. (23), i.e. only one linkage configuration. Thus, the branch of the four-bar chain is also the branch of the whole linkage in Fig. 4. If sin c = 0, there are infinite solutions to s3 and such linkage is useless. Thus, sin c cannot be allowed to be zero. Virtual five-bar loop: A Watt six-bar chain consists of two four-bar loops and a six-bar loop ABCEFGA (Fig. 2). The six-bar loop may be also transformed to a five-bar loop through the stretch rotation of a four-bar loop [11]. Thus, a Watt six-bar linkage may be regarded as a degenerate Stephenson six-bar linkage and the mobility analysis method for Stephenson six-bar linkages is fully applicable. To demonstrate the versatility and simplicity of the discriminant method, no stretch rotation is performed in the following discussion. Among the three loops in a Watt six-bar linkage, any two of them may be regarded as the independent loops from which the third loop can be derived. The mobility of a Watt linkage is determined by the mobility of two loops and also the interaction between both loops. Let ABCDA and ABCEFGA be the two independent loops. The loop equation for ABCEFGA in Fig. 2 can be expressed as a2eih2 \u00fe a3eih3 \u00fe a9ei\u00f0h4 b c\u00de a7 a5eih5 \u00bc a6eih6 : \u00f025\u00de Since cos h4 and sin h4 are determined by the four-bar loop ABCDA (Fig. 2), as shown in Eqs. (2) and (3), the above six-bar loop equation is virtually equivalent to a five-bar loop, which can be also realized through stretch rotation [11]. Eliminating eih6 and using the tangent half-angle formula of Eq. (14), Eq. (25) can be written in the form of Eq. (15) and subsequently equations in the form of Eqs. (17), (18) and (20) can be obtained, in which P2 \u00bc a2 2 \u00fe a2 3 \u00fe a2 5 a2 6 \u00fe a2 7 \u00fe a2 9 2a5a7 \u00fe 2a2a3 cos\u00f0h2 h3\u00de 2a1a9 cos\u00f0h4 b c\u00de \u00fe 2a5a9 cos\u00f0h4 b c\u00de \u00fe 2\u00f0a2a5 a2a7\u00de cos h2 \u00fe 2a2a9 cos\u00f0h4 h2 b c\u00de \u00fe 2a3a9 cos\u00f0h4 h3 b c\u00de \u00fe 2\u00f0a3a5 a3a7\u00de cos h3 \u00f026a\u00de Q2 \u00bc 4a2a5 sin h2 4a3a5 sin h3 4a5a9 sin\u00f0h4 b c\u00de \u00f026b\u00de R2 \u00bc a2 7 \u00fe a2 2 \u00fe a2 3 \u00fe a2 5 a2 6 \u00fe a2 9 \u00fe 2a7a5 \u00fe 2a2a3 cos\u00f0h2 h3\u00de 2a7a9 cos\u00f0h4 c\u00de 2a5a9 cos\u00f0h4 b c\u00de 2\u00f0a2a5 \u00fe a7a2\u00de cos h2 \u00fe 2a2a9 cos\u00f0h4 h2 b c\u00de \u00fe 2a3a9 cos\u00f0h4 h3 b c\u00de 2\u00f0a3a5 \u00fe a7a3\u00de cos h3 \u00f026c\u00de S1 \u00bc \u00bda2 cos h2 \u00fe a3 cos h3 \u00fe a9 cos\u00f0h4 b c\u00de a7 2 \u00fe \u00bda2 sin h2 \u00fe a3 sin h3 \u00fe a9 sin\u00f0h4 b c\u00de 2 \u00f0a5 \u00fe a6\u00de2 \u00f027a\u00de S2 \u00bc \u00bda2 cos h2 \u00fe a3 cos h3 \u00fe a9 cos\u00f0h4 b c\u00de a7 2 \u00fe \u00bda2 sin h2 \u00fe a3 sin h3 \u00fe a9 sin\u00f0h4 b c\u00de 2 \u00f0a5 a6\u00de2 \u00f027b\u00de It is noted that substituting Eqs. (2) and (3) into Eq. (26) yields the expression in the form of Eq. (17), which represents the JRS of h2 and h3 of the six-bar loop ABCEFGA (Fig. 2). Since such a six-bar loop can be treated like a five-bar loop in terms of its mobility, it is regarded as a virtual five-bar loop. With h2 as the input, when the four-bar loop or the five-bar loop is at singularities, the whole linkage becomes singular.The above discussion reveals the fundamentals and properties of 4R, 3R1P, 5R, 4R1P, 3R2P, and virtual five-bar chains. These chains compose the planar single-DOF double-loop linkages. Thus, the singleDOF double-loop linkages must satisfy the necessary conditions determined not only by the 4R or 3R1P chains but also by the 5R or 4R1P or 3R2P chains",
            " The mobility issues of single-DOF double-loop linkages are determined by its mathematical fundamentals and hence the proposed method for the branch, singularity and sub-branch identification of single-DOF double-loop linkages are identified directly from the loop equations and their corresponding mathematical fundaments. Therefore, this method is simple and straightforward and it is ideally for the implement with computer-aided program. The algorithm for general and automated mobility analysis of single-DOF six-bar linkages is shown in Fig. 10 and illustrated with the examples below. Example 1. Given the dimensions for the Watt six-bar linkage in Fig. 2 as follows, a1 = 8.0, a2 = 6.0, a3 = 6.0, a4 = 7.0, a5 = 3.0, a6 = 7.0, a7 = 5.27, a9 = 2.43, a10 = 7.0, a = 37.62 . With the above given dimensions, the plots for the Watt linkage are shown in Fig. 9, where the four-bar curve is drawn from Eq. (4) and the motion domain (shade area) is drawn from Eq. (17). The mobility analysis of Watt linkage with the proposed method above can be carried out as follows, (1) Branch identification of four-bar chain: From Eq. (8) or Eq. (10) with D = 0, there are two dead center positions m( 98"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003211_amr.505.154-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003211_amr.505.154-Figure1-1.png",
        "caption": "Fig. 1 Elements of friction clutch",
        "texts": [
            " There are two basic methods available for design of friction clutch, namely the uniform pressure and uniform rate of wear. In this paper, these methods were used to study the effect of non-dimensional radius (R) on inner and outer radius, the axial force, tangential velocity and heat flux. The assumption of a uniform pressure distribution at the interface between mating surfaces is valid for an unworn accurately manufactured clutch with rigid outer discs. The area of an elemental annular ring on a disc clutch is, rrA \u03b4\u03c0\u03b4 2= as shown in Fig. 1. The differential force acting on the disc is, rrpF \u03b4\u03c0\u03b4 2= . Then the total axial force on the clutch is found by integrating the differential force between the limit (inner radius ri and outer radius ro) [12]. )( 22 io rrpF \u2212= \u03c0 (1) The total frictional torque for multiple-disc is, )1()( 3 2 33 1 33 RrcrrnpT oio \u2212=\u2212= \u00b5\u03c0 (2) Where, n is the number of friction surfaces in clutch and       = 3 2 1 np c \u00b5\u03c0 The dimensionless radius ratio (R) is, ( )oi rrR = (3) Substituting Eq. 1 into Eq. 2 and re-arranged yield:- )( )( 3 2 22 33 io io rr rr nFT \u2212 \u2212 = \u00b5 (4) Rearranging Eq"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002513_3.4984-Figure4-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002513_3.4984-Figure4-1.png",
        "caption": "Fig. 4 Vehicle model consisting of two tethered solid rectangular parallelpipeds.",
        "texts": [
            " Hence, A \u2014 3A may be positive even for positive values of A thereby indicating the possibility of a stable rather than an unstable arrangement. As A [see Eq. (25)] characterizes the tether length and the tether connection geometry, it may be seen that the tether offsets the destabilizing effect of the gravitational torque acting on each end body for the cases where /i > 72. To illustrate some numerical results that may be obtained by following the procedure outlined in the stability section and to provide further insight into the relationships among the vehicle parameters, one can choose a model (see Fig. 4) consisting of two identical, but arbitrary, solid, rectangular parallepipeds tethered with the connection points on the surfaces of the solids. The dimensions of the end bodies can now be varied to obtain all possible combinations of moment of inertia ratios, A and A- As the tether connection points are specified as being on the surfaces of the solids, the parameter A varies according to the dimensions of the solid and can be shown to be A = + A)(l + A)/(l + A A)] (32) Without loss of generality, one may let A = 106 and D4 = 1 (33) and see the effects on stability of varying A, A, and A"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000388_14644193jmbd97-Figure13-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000388_14644193jmbd97-Figure13-1.png",
        "caption": "Fig. 13 Vectorial presentation of the speeds due to movements of the spindle center",
        "texts": [],
        "surrounding_texts": [
            "The equations of motion are based on the second law of dynamics, which leads to the net force or moment on a mass balanced by the acceleration of mass. Proc. IMechE Vol. 222 Part K: J. Multi-body Dynamics JMBD97 # IMechE 2008 at UNIV PRINCE EDWARD ISLAND on August 5, 2015pik.sagepub.comDownloaded from The equations of motion in five degrees-offreedom system can be written as follows (Figs 12 to 14) M \u20acx \u00fe Pm i\u00bc1 \u00f0Ki\u00f0di\u00de 3=2 L cos\u00f0ai\u00deL cos\u00f0ui\u00deL\u00de \u00fe Pm i\u00bc1 \u00f0Ki\u00f0di\u00de 3=2 R cos\u00f0ai\u00deR cos\u00f0ui\u00deR\u00de \u00feQx Mg \u00bc 0 \u00f032\u00de M \u20acy \u00fe Pm i\u00bc1 \u00f0Ki\u00f0di\u00de 3=2 L cos\u00f0ai\u00deL sin\u00f0ui\u00deL\u00de \u00fe Pm i\u00bc1 \u00f0Ki\u00f0di\u00de 3=2 R cos\u00f0ai\u00deR sin\u00f0ui\u00deR\u00de \u00fe Qy \u00bc 0 \u00f033\u00de M \u20acz \u00fe Pm i\u00bc1 \u00f0Ki\u00f0di\u00de 3=2 L sin\u00f0ai\u00deL\u00de \u00fe Pm i\u00bc1 \u00f0Ki\u00f0di\u00de 3=2 R sin\u00f0ai\u00deR\u00de \u00feQz \u00bc 0 \u00f034\u00de where ui is as defined in equation (25) and therefore equations of motion are time dependent. Table 1 Spindle properties Diameter of spindle between bearings 0.04 m Length of the spindle 0.55 m Mass of the spindle 5.5 kg Distance between the support bearings 0.215 m Position of the LHS bearing from the centre of gravity 0.0875 m Position of the RHS bearing from the centre of gravity 0.174 m Moment of inertia of the spindle about x- or y-axis 0.051 77 kg m2 Moment of inertia of the spindle about z-axis 0.0044 kg m2 JMBD97 # IMechE 2008 Proc. IMechE Vol. 222 Part K: J. Multi-body Dynamics at UNIV PRINCE EDWARD ISLAND on August 5, 2015pik.sagepub.comDownloaded from"
        ]
    },
    {
        "image_filename": "designv11_12_0000020_jjap.45.4241-Figure4-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000020_jjap.45.4241-Figure4-1.png",
        "caption": "Fig. 4. Measurement setup cross section of a chip fabricated for evaluation of selectivity.",
        "texts": [
            " In our study, a new type of chip has been developed in which only 4 ml of blood is necessary to run the analysis. In the case of ammonia sensors, measurements are carried out firstly in an open-field environment followed by closed-field measurement on the chip. Freshly made stock 1M NH4Cl solution diluted to 1mM, 0.01M and 0.1M has been used. In this study, a saturated calomel reference electrode and double-junction reference electrode containing 0.3M NH4NO3 are used to reduce interferences that might be caused by the leakage of saturated KCl solution from the calomel reference electrode (see setup, Fig. 4). The double-junction configuration design is used to allow saturated KCl solution balancing with NH4NO3 in the beginning without significantly disturbing the targeting solution. Recently, BUN sensors based on ion-sensitive field effect transistors (ISFETs) used for detecting pH change have been developed.13) However, due to some difficulties, the development of BUN has been moved to the ion-selective membrane field. In the history of development, preconditioning of the membrane is still necessary"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002481_tpas.1969.292344-Figure5-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002481_tpas.1969.292344-Figure5-1.png",
        "caption": "Fig. 5. Quadrature-axis damper-circuit current.",
        "texts": [],
        "surrounding_texts": [
            "This paper presents a mathematical analysis of unbalanced operations of synchronous machines with additional field circuits. The proposed method is aimed to be more rigorous than that initially proposed by Doherty, Nickle, and Concordia [5] and should offer solutions with a higher degree of accuracy. The derived expressions for short-circuit currents, torque, and openphase voltage for a double-line-to-ground fault apparently are new and of immediate practical use. The proposed method"
        ]
    },
    {
        "image_filename": "designv11_12_0003160_1.57112-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003160_1.57112-Figure2-1.png",
        "caption": "Fig. 2 Radar deception scenario.",
        "texts": [
            " 6, November\u2013December 2012 1730 D ow nl oa de d by U N IV E R SI T Y O F SY D N E Y o n M ay 2 0, 2 01 3 | h ttp :// ar c. ai aa .o rg | D O I: 1 0. 25 14 /1 .5 71 12 of LOSG, kinematic studies are presented in Sec. III and the analysis of specific phantom target trajectory patterns is carried out in Sec. IV. Multiple phantom target generation is studied in Sec. V. Simulation studies are carried out in Sec. VI followed by some concluding remarks in Sec. VII. II. Problem Definition Consider n ECAVs flying in an area protected by n radars and generating a single coherent phantom target as shown in Fig. 2. The ECAVs are assumed as point masses and the effect of wind disturbances in neglected. The ECAVs are assumed to be stealthy. Considering the point mass model, the stealth is assumed independent of the ECAVaspect and range to the radar. There are two fundamental constraints in this formation-flying of theECAVs. First, the radar-ECAVLOS lines should be concurrent and secondly the individual phantom targets should all be generated at this point of concurrence. Here i, ri, and Ri are the LOS angle, ECAV radial position, and phantom target radial position, respectively, with respect to the corresponding engaged ith radar with i 1; 2 ",
            " The phantom target is generated along the LOS of the ECAVwith respect to the deceived radar. As the ECAVs move along their trajectories so does the phantom target. The motion kinematics can be analyzed from the LOSG point of view wherein the ECAV is always placed between the phantom target and the radar. Kinematic relations for position, velocity, and lateral acceleration of the ECAVs is derived in the following subsections. A. Position Relations Consider the radar deception scenario as shown in Fig. 2 where all the individual phantom targets should be coherent for a successful deception. Therefore, matching the coordinates of individual phantom targets generated by each ECAV, it can be deduced that xpt xr1 R1 cos 1 xr2 R2 cos 2 . . . xrn Rn cos n (3) and ypt yr1 R1 sin 1 yr2 R2 sin 2 . . . yrn Rn sin n (4) where xpt; ypt and xri; yri are the position coordinates of the phantom target and the ith radar. Note that Eqs. (3) and (4) are 2n equations which can be solved for deceiving parametersRi and i for each ECAV"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003761_9781118818428-Figure21.4-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003761_9781118818428-Figure21.4-1.png",
        "caption": "Figure 21.4 Notation for collision algorithm (Continued overleaf )",
        "texts": [
            " The scan needs to cover every possible pair of molecules j and k \u2013 but there is scope for economy: suppose the first molecule of a scan, molecule j = 1, has been inspected for possible contact with all molecules k up to k = 1500 and a collision eventually registered. Molecule 1 is not allowed further collisions, so the scan recommences \u2013 but not at j= 2, because the potential collision j = 1, k = 2 has already been tested, but with molecule j = 2. Each successive scan in the sequence of nmol scans per \u0394t is thus shorter by one than its predecessor in the sequence. Further economy is possible, as explained after an account of collision mechanics. Figure 21.4 shows molecule j at the instant of collision with molecule k. Under present assumptions respective components of momentum in the tangential direction are unchanged by the impact. Components of momentum along a line through the origins follow the law for collinear impact: uj \u2032 = \u2212uk; uk \u2032 = \u2212uj (21.1) The prime (\u2032) indicates values after impact. Ultimate Lagrange formulation? 241 A way of implementing these principles is first to arrest one of the particles. In Figure 21.4 molecule k is arrested by subtracting uk from all x components of velocity and vk from all y-components. The step is equivalent to transformation to a new reference frame moving with components \u2212uk and \u2212vk. A line through both origins defines the point of contact and the directions of normal and tangential components of momentum exchange. With subscript c to denote coordinates at the instant of impact: \u03b8 = atan{(ykc \u2212 yjc)\u2215(xkc \u2212 xjc)} (21.2) (The numerical value of the argument of Equation 21.2 \u2013 even when properly signed \u2013 is ambiguous"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002632_3.4868-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002632_3.4868-Figure1-1.png",
        "caption": "Fig. 1 Definition of geometrical symbols.",
        "texts": [
            "4 86 8 1788 AIAA JOURNAL VOL. 6, NO. 9 Rigid Body Motions In Ref. 7, it was pointed out that rigid body motions can be represented in the curved field by 0 0 0 cos<\u00a3 \u2014 si 0 sin$ cos<\u00a3 0 \u2014 r(cos$ \u2014 cos/3) \u2014 r si \u2014 r(l \u2014 cos$ cos/3) \u2014 \u00a3 sin<\u00a3 \u00a3 cos$ r cos/3 sin$ \u00a3 cos$ \u00a3 sin</> where $s, $y, &, 6XJ 6y, dz are, respectively, the three components of a general rigid body translation and the three components of a general rigid body rotation of small amplitude in the system of reference (x, y, z) (see Fig. 1). Using (8) in (6) for each of the theories, we get the following results. Theory (a) leads to \u2014 r\"\"1 cos/ (9) showing that the classical approximate theory is going to introduce artificial constraints in the element under two components of rigid translation and all three components of rotation. With such a theory, rigid body motions can never be completely strain-free, even when the displacement field is constructed in such a way as to contain an accurate description of these motions. Theory (b) leads to KI, = r~l sm4>6y \u2014 r~l cos<\u00a303 (10) showing that this theory introduces two artificial constraints for rotation components Oy and 6Z"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000041_j.mechmachtheory.2006.05.006-Figure3-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000041_j.mechmachtheory.2006.05.006-Figure3-1.png",
        "caption": "Fig. 3. Vector representation of the cutting force.",
        "texts": [
            " It is well known that for fully parallel LAPs, the kinematically admissible virtual displacement of the driving joints is related to the virtual displacement and rotation of the moving platform by [15] dQ \u00bc \u00bdJp dbX; \u00f010\u00de where dQ is a 6 \u00b7 1 vector containing the virtual displacement of the joint variables, and [Jp] is the 6 \u00b7 6 overall Jacobian matrix of the robot. Therefore, the virtual work done by the driving forces can be written as dW q e \u00bc dbXT\u00bdJp Ts; \u00f011\u00de where s is a 6 \u00b7 1 vector containing the driving forces provided by the linear actuators. On the other hand, as shown in Fig. 3, the external force exerted on the tip of the tool bit during a machining process can be written as [22] f \u00bc f \u00f0ut \u00fe krun\u00de; \u00f012\u00de where f is the magnitude of the cutting force and ut is a unit vector along the tangential direction of the cutting path; kr is a constant coefficient, the value of which is dependent on various factors such as the tool shape, chip thickness, feed rate, spindle speed, etc., and un = w \u00b7 ut is a unit vector which denotes the direction normal to the cutting path. Consequently, the virtual work done by the cutting force can be written as dW f e \u00bc dbXTbFf ; \u00f013\u00de where bFf is a 6 \u00b7 1 composite vector given by bFf \u00bc f ut \u00fe krun \u2018t\u00f0un krut\u00de \u00fe rtw ; \u00f014\u00de in which \u2018t is the distance from the center of mass of the moving platform to the tip of the tool bit, and rt is the radius of the circular cross-section of the tool bit"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003761_9781118818428-Figure17.6-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003761_9781118818428-Figure17.6-1.png",
        "caption": "Figure 17.6 For given \u03b4, the response to halved flow passage length Lr is a doubling of temperature gradient intensity \u2013 offset by doubled free-flow area Aff. With rh\u2215Lr unaltered, Re is un-changed, as is NTU value and solution independent of \u2202Tw\u2215\u2202x!",
        "texts": [],
        "surrounding_texts": [
            "180 Stirling Cycle Engines",
            "Wire-mesh regenerator \u2013 \u2018back of envelope\u2019 sums 181\nPressure p and local mass rates m\u2032 vary with crank angle as per the well-known Schmidt algebra.\nInstantaneous Re is calculated from local, instantaneous m\u2032. Instantaneous St follows in terms of Re from published steady-flow correlation appropriate to volume porosity.\nMatrix dead volume \u03b4r = Vdr\u2215Vsw is fixed while examining the effect of varying hydraulic radius and stack length. Unexpected benefits arise in relation to heat exchange intensity, which is that arising when a particle flows at velocity u relative to matrix temperature gradient \u2202Tw\u2215\u2202x. In the absence of heat exchange, temperature difference \u0394T = T \u2212 Tw when tracking the particle increases at rate D\u0394T\u2215dt = \u2212u\u2202Tw\u2215\u2202x. Under present assumptions \u2212u\u2202Tw\u2215\u2202x \u2248 \u2212u(TC \u2212 TE)\u2215Lr. Varying Lr (without compensatory change in TE and/or TC) smacks of moving the goal posts. With",
            "182 Stirling Cycle Engines\ndead volume ratio \u03b4r fixed, however, halving Lr (thereby doubling \u2202Tw\u2215\u2202x) is precisely offset (from the point of view of NTU = StLr\u2215rh) by the doubling of free-flow area Aff \u2013 and corresponding halving of u. (NB: expressed in terms of m\u2032, Reynolds number Re \u2013 and thus St \u2013 is independent of u \u2013 Equation 17.10).\nWithin the above constraints, thermal design reduces to selection of hydraulic radius rh and length Lr. The algebra becomes particularly compact when formulated in terms of dimensionless variables (similarity variables) Lr\u2215Lref and rh\u2215Lr.\nWhat ostensibly calls for return to first principles has been short-circuited here by the high\ntemperature recover ratio required by the application, achievable only by a combination of high NTU and large thermal capacity ratio TCR\u2217. This amounts to a limiting case for compre-\nhensive regenerator solutions (e.g., those of the author, 1997, Chapter 7), which point to temperature distributions in both fluid and matrix of invariant, linear gradient \u2202T\u2215\u2202x \u2248 \u2202Tw\u2215\u2202x \u2248 (TC \u2212 TE)\u2215Lr. With the exception of short distances close to x = 0 and to x = Lr, instantaneous temperature difference \u0394T (= T \u2212 Tw) is independent of x. This allows the cyclic variation in \u0394T, Tw \u2013 and hence T \u2013 to be defined in simple algebra. Cycle variations of mass ratesme \u2032 andmc \u2032 at exit from expansion and compression exchangers, together with those of instantaneous pressure p (assumed uniform throughout the matrix) are\nthose of the Schmidt algebra.\nThe start point is the standard thermodynamic relationship ds = cpDT\u2215T \u2013 RDp\u2215p. The total differential operator D (= \u2202\u2215\u2202t + u\u2202\u2215\u2202x) indicates that the relationship applies while following unit mass of gas. In the absence of viscous dissipation ds = dq\u2215T. Recalling that R = cp(\u03b3 \u2212 1)\u2215\u03b3:\ndq = cpDT \u2212 cp[(\u03b3 \u2212 1)\u2215\u03b3]TDp\u2215p\nHeat rate per unit mass q\u2032 = dq\u2215dt can be expressed in terms of (variable) heat transfer coefficient h, and instantaneous local temperature difference \u0394T = T \u2212 Tw:\n\u2212hpw\u0394Tdt\n\u03c1cpAff\n= DT \u2212 [(\u03b3 \u2212 1)\u2215\u03b3]TDp\u2215p (17.14)\nIn Equation 17.14 pw is wetted perimeter [m].\nBeing uniform with x at any instant, Dp\u2215p can be written dp\u2215p. DT is replaced by DT \u2212 DTw + DTw, and thus by D\u0394T + DTw. By definition DTw\u2215dt = \u2202Tw\u2215\u2202t + u\u2202Tw\u2215\u2202x, where, by earlier hypothesis \u2202Tw\u2215\u2202x \u2248 (TC \u2212 TE)\u2215Lr and where, in consequence of the same hypothesis \u2202Tw\u2215\u2202t = dTw\u2215dt, so that DT\u2215dt \u2248 D\u0394T + \u2202Tw\u2215\u2202t + u(TC \u2212 TE)\u2215Lr. Substituting into Equation 17.14:\n\u2212hpw\u0394T \u03c1cpAff = D\u0394T\u2215dt + dTw\u2215dt + u(TC \u2212 TE)\u2215Lr \u2212 [(\u03b3 \u2212 1)\u2215\u03b3]Tp\u22121Dp\u2215dt (17.15)\nEquation 17.15 is shaping up to be a first-order, total differential equation in \u0394T of the form D\u0394T\u2215dt + P\u0394T = Q. This will be susceptible to incremental solution over a succession of finite time increments \u0394t. The process will require a value for dTw\u2215dt at each step, dealt with"
        ]
    },
    {
        "image_filename": "designv11_12_0002008_robot.2010.5509480-Figure4-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002008_robot.2010.5509480-Figure4-1.png",
        "caption": "Fig. 4. (i) Schematic of a robot intended to follow a straight path inclined at an angle \u03d5, (ii) Simulation of the robot trajectory using an autonomous wall following algorithm. Initial condition: (x, y, \u03b8) = (\u22120.5, 0, \u03c0/2)",
        "texts": [
            " These programs interface with an on-board library which has access to current position, orientation of the robot, obstacle free area around it, and infant\u2019s joystick inputs. The training path, shown in Fig. 3, is chosen to consist of three straight lines interspersed with a right and a left turn. A robot could autonomously follow this path using the wall following strategy described below. The research challenge is if a special needs infant driver will learn such a wall following strategy, when assisted by the force feedback joystick. Fig. 4(i) shows the schematic of a robot with the goal to follow a wall inclined at \u03d5 from the horizontal. The kinematic model of a differentially driven mobile robot has the following form:    x\u0307c = v cos \u03b8 y\u0307c = v sin \u03b8 \u03b8\u0307 = \u03c9. (1) Here, xc and yc are coordinates of the robot center and \u03b8 is its orientation. d is the normal distance between the robot center and the inclined path. The inputs to the robot are the translational speed v and rotational speed \u03c9. In the figure, the current heading of the robot is shown at an angle \u2206\u03b8 from the wall. A wall following algorithm, such as [15], is an error correcting control law that specifies the inputs v and \u03c9 such that d \u2192 0 and \u2206\u03b8 \u2192 0 as time increases. This control law is given by { v = vdes \u03c9 = \u2212 k1d vdes cos\u2206\u03b8 \u2212 k2 tan \u2206\u03b8 . (2) We divide the training area into three regions: I, II, and III (Fig. 3). The robot will switch to track the next line if it is inside the corresponding region. Fig. 4(ii) shows simulation of a path when such a strategy is applied autonomously to the mobile robot. v and \u03c9 computed using the wall following strategy can be viewed as ideal commands for an autonomously driven robot. However, in experiments, the speed commands are given by the infant driver through the joystick. Hence, v and \u03c9 commands need to be mapped on to the motion of the joystick. A joystick has predominantly two motions - forward/backward and left/right. We map pure forward/backward motion of the joystick to forward/backward motion of the vehicle along the heading direction"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000375_02286203.2008.11442485-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000375_02286203.2008.11442485-Figure1-1.png",
        "caption": "Figure 1. View of the harmonic drive gear transmission components.",
        "texts": [
            " The simulation and experimental results with the model validation, discussion, and conclusions are given in Sections 6\u20138. The harmonic drive system considered for our analysis is composed of a motor actuator, a harmonic drive gear and an inertial load. The harmonic drive gear consists of the mechanical assembly of three components: a rigid circular spline, an elliptical wave generator, and a nonrigid flexible spline or flexspline, which form together a compact, high-torque, high-ratio, in-line gear mechanism as shown in Fig. 1. A harmonic drive test apparatus was designed and built at Rice University as a platform to perform various types of experiments on the harmonic drive and to characterize the different errors inherent in its operation while preventing any external error component from being imposed [8, 16]. The system is shown in Fig. 2. It has its major axis of motion in the vertical plane to avoid the radial loading problem. A special design of vertical support plates and circular steel pipe sections with a highly stable platform was also used to maintain torsional integrity of the system"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003592_jmes_jour_1965_007_030_02-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003592_jmes_jour_1965_007_030_02-Figure1-1.png",
        "caption": "Fig. 1",
        "texts": [],
        "surrounding_texts": [
            "Equations (I), (2) and (5), with the condition 1GC = 0, yield 1 1 G+1 #, - # = - -z2 tan2 u cos z,hc + higher terms 6,- 8 = - z2 tan2 u sin $,+higher terms 1: 28, Ga 1 G+1 2 Ga By means of these equations we can examine the contours of constant separation s between the tooth surfaces in the neighbourhood of the point of contact C. Let D be any point on the surface of the pinion tooth near to C, lying in a transverse section of co-ordinate z. The points S and Q will be regarded as coincident in what follows (Fig. 2). In the section of co-ordinate z let M be the point on the pinion surface such that MS makes an angle 4, with HS, the perpendicular to AB. Let 0 be the angle subtended at S by MD. Then for co-ordinates of D \\ ,z tan d G2w2a2 . (4) R, = - .. 6, sinz ucos u' ~ * . J 0 U R N AL iM E C H AN I C A L E N G I N E E R I N G S C I E N C E VoI 7 No 2 1965 at UNIV OF VIRGINIA on June 5, 2016jms.sagepub.comDownloaded from"
        ]
    },
    {
        "image_filename": "designv11_12_0002335_detc2010-28926-Figure5-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002335_detc2010-28926-Figure5-1.png",
        "caption": "Figure 5. Example of several linkages corresponding to the same determinate truss. (a) The Assur Graph. (b), (c), (d) The corresponding linkages.",
        "texts": [
            " From the above it follows that once we have all the Assur Graphs it is possible to construct all different determinate trusses by composing different Assur Graphs, each time in a different order. The transformation from determinate trusses into planar linkages is easy and is done by just augmenting a driving 3 Copyright \u00a9 2010 by ASME Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 08/12/2018 Terms of Use: http://www.asme.org/about-asme/terms-of-use 4 Copyright \u00a9 2010 by ASME link, each time to a different ground vertex of the corresponding determinate truss. In Figure 5 we can see the three planar linkages in which the driving link is augmented, each time to a different ground vertex. A B Although the paper is aiming towards synthesis of linkages through AGs, in this section it is explained briefly that the concept of decomposition of a system into AGs enabling at each step to analyze a small component of the system. The method relies on the properties of the decomposition graph. In the first sub-section we show how to analyze mechanisms using bottom-up method, i"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000246_978-3-540-88518-4_120-Figure5-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000246_978-3-540-88518-4_120-Figure5-1.png",
        "caption": "Fig. 5. The schematic diagram of the mother robot",
        "texts": [
            " A tracked structure with double units is adopted for the mother robot to increase the adhesion to ground. Also it supplies a \u201cpocket\u201d used in carrying the baby robot which can enter and get out of the mother robot through the \u201cpocket\u201d. Because the mother robot communicates with and gets power from the remote control center via a cable, the mother robot has a mechanism of dragging and releasing the cable. Moreover, it carries various detecting sensors. The schematic diagram of the mother robot is shown in Fig. 5. The baby robot also is an independent and whole unit. As discussed above, the robot will experience complicated terrain under the mine well, so it is necessary to bring forth much more strict requirements for its ability of climbing over obstacles. Therefore, the baby robot is serially connected by multiple joints just like a snake. Otherwise, the baby robot takes communication nodes to improve the performance of wireless communication under the mine well. The baby robot generally includes a head unit, a tail unit, a control unit, a communication node cabinet and a power unit, as shown in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002129_s0263574709005426-Figure3-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002129_s0263574709005426-Figure3-1.png",
        "caption": "Fig. 3. Vector diagram of a spatial displacement.",
        "texts": [
            " In this subsection, we describe a method to represent the location of a rigid body in a kinematic chain with respect to a coordinate frame, based on the successive screw displacement concept. First, we present the transformation matrix associated with a screw displacement, and then describe the concept of the resultant screw of two successive screw displacements. representation. The Chasles theorem states that the general spatial displacements of a rigid body are a rotation about and a translation along some axis. Such a combination of translation and rotation is called a screw displacement.10 Below, we derive a homogeneous transformation that represents a screw displacement.8 Figure 3 shows a point P of a rigid body, which is displaced from a first position P1 to a second position P2 by a rotation \u03b8 about a screw axis followed by a translation of t along the same axis. The rotation brings P from P1 to P r 2 , and the translation brings P from P r 2 to P2. In the figure, s = [sx sy sz]T denotes a unit vector along the direction of the screw axis, and so = [sox soy soz]T denotes the position vector of a point lying on the screw axis. The rotation angle \u03b8 and the translation t are called the screw parameters"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003801_acs.analchem.5b02804-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003801_acs.analchem.5b02804-Figure1-1.png",
        "caption": "Figure 1. (a) Schematic of the cross-section of the carbon ultramicroelectrodes, where d is the interelectrode distance, a is the electrode radius, and t is the Al2O3 layer thickness. Top-view scaled representations of (a): (b) macro, (c) 1.54CUA, (d) 11CUA, and (e) 90CUA electrodes with their d/a values indicated.",
        "texts": [
            " Absorbance spectra were acquired with an Agilent Instrument 8453 UV\u2212vis\u2212NIR spectrometer. A homemade Teflon spectroelectrochemical cell was used, and contact to the working electrode was made with copper tape. A CHI 440 potentiostat (CH Instruments Inc.) was used to perform the CVs. \u25a0 RESULTS Geometry of CUAs. The carbon ultramicroelectrode arrays (CUAs) are here labeled as 1.54CUA, 11CUA, and 90CUA, with the numeric value specifying the micron-sized diameter of the polystyrene spheres (PSS) used during their fabrication, 1.54, 11, and 90 \u03bcm, respectively. Figure 1 graphically displays the cross-sectional as well as the scaled unit-cell geometry of these electrodes, where d is the interelectrode distance, a is the electrode radius, and t is the alumina layer thickness. As t \u226a a, it can be assumed that the electrochemical behavior of these electrodes will resemble that of a coplanar electrode. The scaled representations in Figure 1c\u2212e indicate that the alumina blocking layer dominates the arrays\u2019 geometry, with only a minute area occupied by the exposed carbon electrodes, which is indicated by the black dots (circles represent the residual alumina that had coated the PSS). As the schematic is a scaled geometric representation, it can be seen from Figure 1c\u2212 DOI: 10.1021/acs.analchem.5b02804 Anal. Chem. XXXX, XXX, XXX\u2212XXX B D ow nl oa de d by U N IV O F C A M B R ID G E o n Se pt em be r 15 , 2 01 5 | h ttp :// pu bs .a cs .o rg P ub lic at io n D at e (W eb ): S ep te m be r 14 , 2 01 5 | d oi : 1 0. 10 21 /a cs .a na lc he m .5 b0 28 04 e that the exposed carbon area is very small compared to the alumina layer for the CUAs. In fact, it is difficult to even resolve the individual electrodes for 90CUA in this scaled representation. The geometric parameters for each of these arrays are displayed in Table 1",
            " Accordingly, a 10 nm alumina coated PPF was electrochemically characterized for its specific capacitance by integrating the area of its CV curve in background solution (0.1 M PBS), and it was found to be 5.6 \u00d7 10\u22127 F cm\u22122, which is within a 10% error of the expected theoretical value indicated above. Although this specific capacitance is small compared to that obtained from a bare PPF carbon electrode (1.1 \u00d7 10\u22125 F cm\u22122), its contribution to a CUA\u2019s total capacitance is significant due to alumina\u2019s relatively large area when compared to that of the exposed carbon electrode area, as mentioned previously and shown schematically in Figure 1. Its effect on the total capacitance is graphically shown in Figure 4, where the capacitance value obtained experimentally from the arrays was proportioned between that from the exposed carbon (hashed bars) and from the alumina layer (solid bars). The exposed carbon capacitance values were calculated using the experimental specific capacitance obtained from the bare PPF carbon and the total exposed carbon area from Table 1 for each CUA. While the alumina capacitance values were obtained using the specific capacitance of alumina coated PPFs with varying thicknesses of 10, 20, and 50 nm and the alumina\u2019s areal coverage for each CUA, it is assumed here that the capacitances are additive since they act as capacitors connected in parallel"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002185_bit.260090407-Figure3-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002185_bit.260090407-Figure3-1.png",
        "caption": "Fig. 3. Schematic diagram of sparge pipe assembly.",
        "texts": [
            " The air and nitrogen werc separately metered into the vessel arid sterilized by filtration. The oxygen probe was inserted through the base plate of the vessel. The filtered and dried effluent gas was supplied to an infrared type carbon dioxide analyzer (model SBK, Hilger & Watts Ltd., London NWl) and to a paramagnetic oxygen analyzer (model OA 137, Servomex Controls Ltd., Crowborough, Sussex). The gas flow diagram for the system is shown in Figure 2. The sparge pipe consists of a 1/4-in. O.D. stainless steel pipe, fitted a t its bottom end The sparge pipe assembly is shown in Figure 3. 5 18 11. S. BLYNN ANT) h4. 13. LILLY with a plug through which a 1/16-in. hole has been drilled. The drive of a variable-speed reversing motor having a full load output of 2-10 rpm (type KQR, Citenco Ltd., Boreham Wood, Herts.) is suitably reduced by a gear box, one output shaft of which is connected to the sparge pipe by a rack and pinion gear. With the present system at maximum motor speed the sparge pipe takes 17 min. to move from 0 to 1 0 0 ~ o immersion. The lower guide, which serves also as a seal in the stainless steel top plate of the vessel, corisists of a I\u2019TFE rod screwed into the culture vessel top plate arid reamed internally to make a tight fit around the sparge The upper guide for the sparge pipe is a PTFE bush",
            " However, the limiting factor at present in the rate of response and degree of precision obtainable with this controller is riot the characteristics of the controller itself but the transfer lag of the control loop, i.e., the slow response of the oxygen probe. In addition, these oxygen probes have a high temperature coefficient and therefore, as can be seen from Figure 10a, the control precision obtained is affected by the precision of temperature control. Unlike the vortex-aer:tt,ed systcm uscd hy nI:icT,ennan and Pirt7 the supply side ctlpucihrcc. of the syste~n is rr1:~tively 101v arid is givcii roughly by tlw volume oI' gab held up i i i tlie liquid (see Fig. 3 which shows thc reltlt ivrly small contribution of top air to oxygen transfer, but i t should be noted that a stirred, gas sparged liquid has a more turbulent surface than one merely stirred). BIOTECHNOLOGY AND I<IOENGINEERING, VOL. IX, ISSUE 4 OXYGEN TENSION I N RlICltOBIAL CULTURES 5\u201929 Against this advantage may be set the possible disadvantage that the oxygen solution rate is not a linear function of sparge pipe immersion (Fig. 5). This would be expected to make the controller inherently difficult to tune since the optimum setting of motor speed will vary according to the range of immersion in which the sparge pipe is working"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002035_j.bspc.2009.12.001-Figure5-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002035_j.bspc.2009.12.001-Figure5-1.png",
        "caption": "Fig. 5. Schematic representation of the limb fixation and force measurement system.",
        "texts": [
            " T he first ks trains are used to establish the initial fatigue model, and then the parameters, a1(p), b1(p) and b2(p) are updated in real-time. The exact value of ks can be adjusted based on some lot is the zoom of the left plot with a\u03021 2 \u00bd0:9;1 . criteria, for example, ks is the value that the fatigue phenomenon is initially observed. Fourteen subjects with chronic SCI (>2 years) provided written informed consent, as approved by the University of Iowa Human Subjects Institutional Review Board. A detailed description of the stimulation and the force transducing systems has been previously reported [8\u201310] (Fig. 5). In brief, the subject sat in a wheelchair with the knee and ankle positioned at ninety degrees. The foot rested upon a rigid metal plate, and the ankle was secured with a soft cuff and turnbuckle connectors. Padded straps over the knee and forefoot ensured isometric conditions. The tibial nerve was supra-maximally stimulated in the popliteal fossa using a nerve probe and a custom computer-controlled constant-current stimulator. Stimulation was controlled by digital pulses from a data-acquisition board (Metrabyte DAS 16F, Keithley Instruments Inc"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002042_j.actaastro.2010.07.010-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002042_j.actaastro.2010.07.010-Figure1-1.png",
        "caption": "Fig. 1. Orbital and satellite coordinate systems.",
        "texts": [
            " Of course, the dynamical model of the satellite in elliptic orbits is relatively complicated due to the time varying orbital angular velocity and the satellite position from the Earth. Simulation results are presented which show that the NCEA law accomplishes precise attitude control of the satellite in an elliptic orbit, despite large parameter uncertainties. The organization of the paper is as follows. The mathematical model of the satellite in elliptic orbit is described in Section 2. Section 3 presents an adaptive control module and an estimator is designed in Section 4. Finally, Section 5 presents the simulation results. Fig. 1 shows an unsymmetrical satellite with its center of mass S moving in an elliptic orbit around the Earth\u2019s center 0. The apparent position of the Sun is indicated by the solar aspect angle f, measured from the line of nodes. Two identical, highly reflective, lightweight solar flaps P1, and P2 are mounted on the satellite for generating control moments for the control of the pitch motion. The control moments are nonlinear functions of the rotation angles \u00f0d1,d2\u00de of the flaps. The principal body-fixed frame has x-axis along the local vertical, the z-axis (not shown here) is normal to the orbital plane, and y-axis represents the third axis of this right-handed frame"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003990_wcica.2012.6359128-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003990_wcica.2012.6359128-Figure2-1.png",
        "caption": "Fig. 2. Moments resulting from the rotor flapping (Gavrilets et al, 2001)",
        "texts": [
            " 1) Main rotor forces and moments: The forces and moments produced by the main rotor can be expressed in terms of rotor flapping. The forces are the projections of the rotor thrust vector on the hub plane under an assumption. It is assumed that the direction of thrust vector is perpendicular to the rotor disc. The force components along the helicopter body axes can be written as \ud835\udc4b\ud835\udc5a\ud835\udc5f = \u2212\ud835\udc47\ud835\udc5a\ud835\udc5f\ud835\udc60\ud835\udc56\ud835\udc5b(\ud835\udefd\ud835\udc59\ud835\udc54) \u2248 \u2212\ud835\udc47\ud835\udc5a\ud835\udc5f\ud835\udefd\ud835\udc59\ud835\udc54 \ud835\udc4c\ud835\udc5a\ud835\udc5f = \ud835\udc47\ud835\udc5a\ud835\udc5f\ud835\udc60\ud835\udc56\ud835\udc5b(\ud835\udefd\ud835\udc59\ud835\udc61) \u2248 \ud835\udc47\ud835\udc5a\ud835\udc5f\ud835\udefd\ud835\udc59\ud835\udc61 \ud835\udc4d\ud835\udc5a\ud835\udc5f = \u2212\ud835\udc47\ud835\udc5a\ud835\udc5f\ud835\udc50\ud835\udc5c\ud835\udc60(\ud835\udefd\ud835\udc59\ud835\udc54)\ud835\udc50\ud835\udc5c\ud835\udc60(\ud835\udefd\ud835\udc59\ud835\udc61) \u2248 \u2212\ud835\udc47\ud835\udc5a\ud835\udc5f (7) where \ud835\udefd\ud835\udc59\ud835\udc54 and \ud835\udefd\ud835\udc59\ud835\udc61 are the longitudinal and lateral rotor flapping angles, respectively, \ud835\udc47\ud835\udc5a\ud835\udc5f denotes the thrust vector of the main rotor. Fig. 2 shows the moments resulting from the rotor flapping. The rolling and pitching moments \ud835\udc40\ud835\udc58 can be expressed as \ud835\udc40\ud835\udc58 = \ud835\udc3e\ud835\udefd\ud835\udefd, where \ud835\udc3e\ud835\udefd described in [11] is a constant stiffness coefficient with which the restraint can be approximated by means of a linear torsional spring. The longitudinal and lateral moments can be written as \ud835\udc40\u210e = \ud835\udc47\ud835\udc5a\ud835\udc5f\u210e\ud835\udc5a\ud835\udc5f\ud835\udefd. The total rolling and pitching moments produced by the main rotor are expressed as \ud835\udc3f\ud835\udc5a\ud835\udc5f = (\ud835\udc47\ud835\udc5a\ud835\udc5f\u210e\ud835\udc5a\ud835\udc5f +\ud835\udc3e\ud835\udefd)\ud835\udefd\ud835\udc59\ud835\udc61 \ud835\udc40\ud835\udc5a\ud835\udc5f = (\ud835\udc47\ud835\udc5a\ud835\udc5f\u210e\ud835\udc5a\ud835\udc5f +\ud835\udc3e\ud835\udefd)\ud835\udefd\ud835\udc59\ud835\udc54 (8) During flight, the main rotor produces a torque effect that turns the fuselage of the helicopter in the opposite direction of the rotation of the rotor"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003847_ecc.2013.6669739-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003847_ecc.2013.6669739-Figure1-1.png",
        "caption": "Fig. 1. The inertial coordinate system O and the vehicle coordinate system V, used to describe the dynamics of the system.",
        "texts": [
            " The objective of this experiment is to balance an inverted pendulum on a quadrocopter while the vehicle flies horizontal circles. The derivations from [13] are reproduced here in abbreviated form for the purpose of completeness; the reader is referred to the previously published paper for a more thorough discussion. The quadrocopter is modeled as a rigid body with six degrees of freedom: Its position (x, y, z) in the inertial coordinate system O, and its attitude, represented by the rotation between the inertial coordinate system O and the body-fixed coordinate system V, as shown in Figure 1. The rotation is parameterized by three Euler angles, representing rotations about the z-axis (\u03b1), the y-axis (\u03b2) and the xaxis(\u03b3), executed in this order: O VR(\u03b1, \u03b2, \u03b3) = Rz(\u03b1) Ry(\u03b2) Rx(\u03b3) . (1) The control inputs are the rotational rates of the vehicle about the three body axes (\u03c9x, \u03c9y, \u03c9z) and the collective, mass-normalized thrust applied by the vehicle along its body z-axis, (a; in units of acceleration). It follows that the differential equations governing the vehicle motion are   x\u0308 y\u0308 z\u0308   = O VR(\u03b1, \u03b2, \u03b3)   0 0 a  +   0 0 \u2212g   (2)   \u03b3\u0307 \u03b2\u0307 \u03b1\u0307   =   cos \u03b2 cos \u03b3 \u2212 sin \u03b3 0 cos \u03b2 sin \u03b3 cos \u03b3 0 \u2212 sin \u03b2 0 1   \u22121   \u03c9x \u03c9y \u03c9z   (3) where g denotes gravitational acceleration"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001628_elan.201000267-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001628_elan.201000267-Figure1-1.png",
        "caption": "Fig. 1. OTTLE cell and holder. (A) Assembly of cell front and side view. (B) Cell holder, 3D and top down view.",
        "texts": [
            " Here we introduce a new OTTLE cell to be used with conventional spectrophotometers for both absorbance and fluorescence measurements. The cell consists of two parts: an OTTLE and a unique cell holder. The cell holder is designed to position the OTTLE, a reference electrode, and an auxiliary electrode for both absorbance and fluorescence measurements and to standardize the distance between electrodes. The OTTLE cell and holder provide a reproducible way to perform spectroelectrochemical experiments in instruments compatible with a standard 1 cm cuvette. The unique OTTLE cell holder (Figure 1B) was designed and fabricated using rapid prototyping technology. This technique was first introduced in the late 1980s and is now used extensively to produce small models and parts [24,25]. Rapid prototyping is based on taking a computer-designed model and automatically transforming it into a physical object by the sequential delivery of material to specified points in space. This process creates a physical model identical to the virtual design. The use of computer-aided design (CAD) software makes creating virtual models of specified dimension quick and easy",
            "5 carbonate buffer/20 % EtOH. The OTTLE cell is constructed from a quartz glass slide (ESCO products) cut to (1.90 1.00 cm), 0.018 cmthick silicone spacers (Specialty Manufacturing Inc., Pineville, NC), and ITO-coated glass slides (Corning 1737F and 7059, 11\u201350 W/square, 130-nm-thick film on 1.1-mm glass, Thin Film Devices, Anaheim, CA) with dimensions of 4.00 1.00 cm. Silicone spacers cut to approximately 1.90 0.20 cm are placed onto the edges of the ITO glass slide and sandwiched between a quartz slide and the ITO (Figure 1A). Two-part quick-set epoxy (Loctite) is applied along the edges of the spacers and allowed to cure for 2 h to hold the components together. OTTLE cells made in this manner were capable of approximately 8 hr of continuous use. The exposed ITO above the quartz slide is used for electrical contact. For all experiments the electrochemical cell consisted of a Pt wire auxiliary electrode, a miniature Ag/AgCl reference electrode (3 M KCl, Cypress Systems), and an OTTLE. Thin layer cyclic voltammetry and coulometry of ferricyanide/ferrocyanide in 1",
            " Laser power was attenuated to 0.5 mW and the sample was exposed to the laser light only during data acquisition to minimize photodegradation. Electroanalysis 2010, 22, No. 19, 2162 \u2013 2166 2010 Wiley-VCH Verlag GmbH & Co. KGaA, Weinheim www.electroanalysis.wiley-vch.de 2163 OTTLE cells have previously been shown to be useful for optical characterization of electrochemical reactions. The majority of these cells have focused on absorbance based spectroelectrochemical detection. By incorporating this unique cell holder (Figure 1B), created using SolidWorks software and rapid prototyping technology, the OTTLE cell can be easily used for both absorbance and fluorescence based measurements. The outer dimensions of the holder are identical to the standard cuvette commonly used in spectrophotometers (1 1 4 cm). However, unlike a standard cuvette, all four side walls have windows for light passage. Two of the inside diagonal corners have slots for the OTTLE cell. This design positions the thin layer cell at a 458 angle with respect to each of the holder s walls"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000171_acc.2006.1656632-Figure2.2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000171_acc.2006.1656632-Figure2.2-1.png",
        "caption": "Figure 2.2: Decomposition of a quantization function for",
        "texts": [
            " Note that \u03c1i(\u00b7) determines coarseness of the quantizer qi(\u00b7, \u00b7) and ai(\u00b7) determines the size of the deadzone for each ui, i = 1, . . . , m. It is important to note that the logarithmic quantizer (3) can be characterized as a class of time-varying sector-bounded memoryless input nonlinearities Q which is given by Q {q : R \u00d7 R m \u2192 R m : q(\u00b7, 0) = 0, [q(t, u) \u2212 (Im \u2212 \u0394(t))u]T[q(t, u) \u2212 (Im + \u0394(t))u] \u2264 0, u \u2208 R m\\(\u2212a(t), a(t))m, a.e. t \u2265 0, and q(\u00b7, u) is Lebesgue measurable for all u \u2208 R m}, (4) where \u0394 diag[\u03b41, . . . , \u03b4m] is such that 0 < \u0394 < Im and \u03b4i = (1 \u2212 \u03c1i)/(1 + \u03c1i), i = 1, . . . , m (Figure 2.2(a)). Note that the sector condition characterizing Q is implied by the scalar sector conditions (1 \u2212 \u03b4i(t))u 2 i \u2264 qi(t, ui)ui \u2264 (1 + \u03b4i(t))u 2 i , ui \u2208 R\\(\u2212a(t), a(t)), a.e. t \u2265 0, i = 1, . . . , m. (5) m = 1 Since \u03c1i(\u00b7) = (1\u2212\u03b4i(\u00b7))/(1+\u03b4i(\u00b7)), i = 1, . . . , m, the coarseness of the input quantizer qi(\u00b7, \u00b7) is determined by \u03b4i(\u00b7) for each i = 1, . . . , m. Though the time variation of q(t, \u00b7) is due solely to the variation of \u0394(t), we write q(t, u(t)) instead of q(\u0394(t), u(t)) for simplicity of exposition. To design controllers for (1) we decompose the quantization function q(\u00b7, \u00b7) into a linear part and a nonlinear part so that q(t, u) = u + qs(t, u), (6) where qs : R \u00d7 R m \u2192 R m. Note that the transformed nonlinearity qs(\u00b7, \u00b7) belongs to the set Qs given by (see Figure 2.2(b)) Qs {qs : R \u00d7 R m \u2192 R m : qs(\u00b7, 0) = 0, qT s (t, u)qs(t, u) \u2212 uT\u03942(t)u \u2264 0, u \u2208 R m\\(\u2212a(t), a(t))m, a.e. t \u2265 0, and qs(\u00b7, u) is Lebesgue measurable for all u \u2208 R m}. (7) In this paper, we assume that the functions f(\u00b7) and/or G(\u00b7) in (1) are unknown. In light of this assumption, we have no information a priori on how fine the quantizer should be, that is, how small \u0394 should be to stabilize the closed-loop system. In the following section, we propose a method of updating the quantization function on-line along with an adaptation law of a controller gain matrix such that the closed-loop system is stable for the nonlinear uncertain plant"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000558_j.1467-9450.1977.tb00279.x-Figure5-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000558_j.1467-9450.1977.tb00279.x-Figure5-1.png",
        "caption": "Fig. 5. The optical arrangement (the letters defined in the text).",
        "texts": [
            " Fig. 1 above gives two examples of proximal stimuli with horizontal \u201cdistal\u201d stimuli, and Fig. 4 demonstrates the construction of the proximal stimulus for a non-horizontal \u201cdistal\u201d stimulus. Stimuli of the former kind moved 0.27 cycleslsec, stimuli of the latter kind 0.24 cycleslsec. The subject was allowed as many cycles as he wanted. Experimental arrangement The stimuli were produced on an oscilloscope (Tektronix Type 565) fed by a small computer (LINC 8). The optical arrangements are shown in Fig. 5 . From the oscilloscope ( A ) , the stimuli were projected onto a focusing screen (B). The subject ( S ) sat on the opposite side of the screen looking monocularly, one eye being occluded by a piece of black cardboard (C) through a large collimator lens ( D ) at focal distance from the screen. This transferred the event on the screen to optical infinity and made it appear in an empty space, perceptually within a distance of a few meters. The peripheral parts of the collimator lens were occluded by a black screen close to the lens"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002739_icra.2011.5980299-Figure5-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002739_icra.2011.5980299-Figure5-1.png",
        "caption": "Fig. 5. The interaction results for 3 different cases from Fig. 7 are shown. Object angle is always kept as \u221245 \u25e6 but the approach angle \u03b1 is changed.",
        "texts": [
            " The X-means algorithm was used to find channel-specific effect categories, and Support Vector Machine (SVM) classifiers were employed to learn effect category prediction. For the power-grasp behavior, 4 clusters were found to represent whole effect space as shown in Table I. Large objects could not be lifted resulting in not-lifted effect. Small objects could be lifted so the height is increased and touch sensor is activated as shown in prototype of lifted effect. In some cases, the grasp was not stable, so the object slided from robot\u2019s hand during lifting but remained in contact with the hand, creating unstable-lifted effect (Fig. 5 (b)). In this effect, the vertical position of the object was not increased (significantly), however the touch sensor remained activated. The disappeared effect was created by the spheres that roll away during interaction. For the precision-grasp behavior, 3 effect categories were obtained as shown in Table II. Because the robot inserted one of its fingers through the aperture of the handle, the grasps were more stable once the object is hold. After the discovery of effect categories, the mapping from the initial object features to these categories was learned for each behavior bj (Predictorbj ()) by multi-class Support Vector Machines (SVMs)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003188_20110828-6-it-1002.01128-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003188_20110828-6-it-1002.01128-Figure1-1.png",
        "caption": "Fig. 1. Lay-out of the \u2018Helicopter\u2019.",
        "texts": [
            " This is done to make possible the convergence of the gains. Given a non-zero reference y\u2217 6= 0, if the system is asymptotically stable, y \u2212 y\u2217 and the \u2018derivative\u2019 y\u0302 will converge to zero while the integral \u222b t 0 (y(\u03c4)\u2212 y\u2217)d\u03c4 will converge to a constant value. No adaptive scheme is therefore possible based on a adaptation law depending of the square of this signal. And indeed with the proposed structure of adaptation there is no such term. The adaptation will stop as soon as the system has converged to the reference. The 3DOF helicopter (Figure 1) is manufactured by Quanser Consulting Inc., www.quanser.com. This setup was modified under demand of LAAS-CNRS for implementation and testing of robust control laws. It 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 a third encoder. Two motors with propellers mounted on the helicopter body can generate a force proportional to the voltage applied to them. The force, generated by the propellers, causes the helicopter body to lift off the ground and/or to rotate about the pitch axis. The system is also equipped with a motorized lead screw that can drive a mass along the main arm in order to impose known controllable disturbances (the so-called Active Disturbance Option, ADO). The following notation is used (see Fig. 1): \u03b8 is the pitch angle; \u03b5 is the elevation angle; \u03bb is the travel angle; vf , vr is control voltages of the front and the rear motors. Denote u = vf \u2212 vr and v = vf + vr. The control voltages vf and vr, applied to the front and rear motors are calculated based on the command signals u, v as follows: vf = 0.5(v + u), vr = 0.5(v \u2212 u). (13) We make the following simplifying assumptions: the \u2018Helicopter\u2019 is considered as a rigid body, i.e. it is assumed that the long arm of the \u2018Helicopter\u2019 rotates about the long axis together with the crossbar, and bending of structural components is neglected; the gyroscopic torques, developed by motor/propeller pairs are neglected; dependence of motor/propeller force gain on the \u2018Helicopter\u2019 airspeed is neglected; influence of aerodynamical pressure forces on the \u2018Helicopter\u2019 body is neglected; dry friction in the pivots is neglected"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002342_j.mechmachtheory.2011.06.006-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002342_j.mechmachtheory.2011.06.006-Figure1-1.png",
        "caption": "Fig. 1. (a) A tree structure multibody system; (b) kinematics of a joint.",
        "texts": [
            " The method from [23] and ([24], pp. 349\u2013353), which is based on Lagrange's equations, is adapted for the use of Kane's equations. Spatial and planar multibody systems are particularly considered. Applications of the algorithm are illustrated through examples. 2. Kinematic description of a system Consider a system of n rigid bodies interconnected by revolute or prismatic frictionless joints. The multibody system has the form of a tree-like multibody structure which moves in a uniform gravitational field (see Fig. 1(a)). In Fig. 1(a), labelling of the bodies is done as in [25] so the body is connected to the fixed reference base body (V0) by either a revolute joint or a prismatic joint which has the notation (V1) while any other body in the system has the subscript greater than the subscript of its preceding adjacent body on a direct path from body (V1). In this case, the direct path between bodies (V1) and (Vi) means that such a path passes through the bodies only once. Note that a joint has the same index as the subscript of the following adjacent body in the pair of bodies connected by this joint",
            " For the numbering procedure the subscripts of bodies satisfy the following relation [25]: 1 b 2 b\u2026 b j b j + 1 b\u2026 b p1 b j1 b\u2026 b p2 b j2 b\u2026 b p3 = n : \u00f01\u00de According to the above mentioned description of the system, the following matrices can be introduced [25]: \u0398 i\u00f0 \u00de \u2208 Ri 1 ; i = 1;\u2026;n \u00f02\u00de whose components are \u0398k; i\u00f0 \u00de = 1; if body Vk\u00f0 \u00de is in the direct path from V1\u00f0 \u00de to Vi\u00f0 \u00de 0; otherwise ; \u00f03\u00de where 1\u2264k\u2264 i. The other descriptions of tree-like multibody systems are given in [20,26\u201328]. The considered multibody system has n degrees of freedom and its motion is described by generalised coordinates qi(i=1,\u2026, n). In the case of a prismatic joint, the coordinate qi represents the relative linear displacement of body (Vi) with respect to its preceding adjacent body (Vh) (see Fig. 1(b)) measured along the joint axis determined by the unit vector e\u2192i, while in the case of a revolute joint this coordinate describes the relative rotation of body (Vi) with respect to (Vh) carried out about the axis e\u2192i. The vector e\u2192i is fixed to the preceding adjacent body (Vh). In Fig. 1(b), the symbol \u03c7i represents an identifier of the joint type. If the ith joint is revolute, then \u03c7i=0, and if prismatic, then \u03c7i=1. Thereby \u03c7i = 1\u2212\u03c7i. Points Oi | and Oi, belonging to body (Vi), are placed on the axes of joints i and i+1, respectively. Point Ci denotes the mass centre of body (Vi) while the vectors \u2192 O j i Oi \u2261 \u03f1\u2192i and\u2192 OiCi \u2261 \u03f1\u2192Ci are fixed to body (Vi). Together with the inertial reference frame, the local coordinate frames Ci\u03bei\u03b7i\u03b6i(i=1, \u2026, n) that are fixed to bodies (V1), \u2026, (Vn), respectively, are introduced",
            " Taking into account the equivalence between Kane's and Lagrange's equations (see for details [29]) as well as computational efficiency of Kane's equations (see e.g. [30\u201332]), the adaptation of the method from [23] and ([24], pp. 349\u2013353) to the case of Kane's equations is given further below. To determine reaction forces in the ith joint, the joint is cut imaginary. After the ith joint has been cut the number of degrees of freedom of the system is increased by five in the multibody system depicted in Fig. 1(a). In regard to this, the additional set of five coordinates are introduced in such a manner that the zero values of these coordinates lead to the configuration of the system before cutting of the ith joint. After cutting of the ith joint, the motion of the system can be observed as the motion with the redundant coordinates qn+1, \u2026,qn+5 subject to the following constraints: fr \u2261 qn+ r = 0; r = 1;\u2026;5; \u00f06\u00de where qn+ r(r=1, \u2026,5) represent relative displacements between two bodies after cutting of the ith joint corresponding to the relative degrees of freedom that are not allowed by the joint before cutting",
            " 0 = 0; r = 1;2 \u0398i; p\u00f0 \u00de \u03bb \u2192 ; r = 3 \u0398i; p\u00f0 \u00de \u03bc \u2192 ; r = 4 \u0398i; p\u00f0 \u00de \u03bd \u2192 ; r = 5; 8>>< >>: \u00f036\u00de \u2202 v\u2192 Cp \u2202q n+r 0 @ 1 A 0 = \u0398i; p\u00f0 \u00de \u03bb \u2192 ; r = 1 \u0398i; p\u00f0 \u00de \u03bc \u2192 ; r = 2 \u03bb \u2192 \u00d7 \u03f1\u2192i + \u03f1\u2192Ci ; r = 3; p = i \u0398i; p\u00f0 \u00de \u03bb \u2192 \u00d7 \u2211 p \u2018= i \u0398i; \u2018\u00f0 \u00de \u03f1 \u2192 \u2018 + \u2211 p \u03b3= i+1 \u0398i; \u03b3\u00f0 \u00de\u03c7\u03b3q\u03b3 e\u2192\u03b3 + \u03f1\u2192Cp ! ; r = 3; p N i \u03bc\u2192 \u00d7 \u03f1\u2192i + \u03f1\u2192Ci ; r = 4; p = i \u0398i; p\u00f0 \u00de \u03bc \u2192 \u00d7 \u2211 p \u2018= i \u0398i; \u2018\u00f0 \u00de \u03f1 \u2192 \u2018 + \u2211 p \u03b3= i+1 \u0398i; \u03b3\u00f0 \u00de\u03c7\u03b3q\u03b3 e\u2192\u03b3 + \u03f1\u2192Cp ! ; r = 4; p N i \u03bd\u2192 \u00d7 \u03f1\u2192i + \u03f1\u2192Ci ; r = 5; p = i \u0398i; p\u00f0 \u00de \u03bd \u2192 \u00d7 \u2211 p \u2018= i \u0398i; \u2018\u00f0 \u00de \u03f1 \u2192 \u2018 + \u2211 p \u03b3= i+1 \u0398i; \u03b3\u00f0 \u00de\u03c7\u03b3q\u03b3 e\u2192\u03b3 + \u03f1\u2192Cp ! ; r = 5; p N i: 8>>>>>>>>>>>>>>>>>< >>>>>>>>>>>>>>>>>: \u00f037\u00de 4. Planar multibody systems Assume that the multibody system depicted in Fig. 1 represents a planar multibody system that is placed in the vertical coordinate plane Oyz. Next, let the coordinate planes Ci\u03b7i\u03b6i(i=1,\u2026, n) of the local coordinate frames defined in Section 2 coincide with theplaneOyz. In further considerations, all the conventions and expressionsdefined in Section2 aswell as Eqs. (19)\u2013(22)hold. 4.1. Determination of constraint reactions in revolute joints Suppose the ith planar joint is revolute. Let coordinate frame Oh\u03be*\u03b7*\u03b6* be fixed to body (Vh) at point Oh so that axis \u03be* coincides with joint axis e\u2192i (see Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002150_med.2009.5164662-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002150_med.2009.5164662-Figure2-1.png",
        "caption": "Fig. 2. Line following.",
        "texts": [
            " The on-board real-time control system, developed in C++, is based on GNU/Linux and run on a Single Board Computer (SBC) which supports serial and Ethernet communications and PC-104 modules for digital and analog I/O. The steering equation of Charlie can be described with (1) where r is yaw rate, \u03c8 is heading, \u03c4N commanded yaw torque, and parameters to be identified are yaw inertia Ir , and drag kr|r| (see [1] for details on model parameters). Ir r\u0307 = \u2212k\u0303r|r|r|r| + \u03c4N \u03c8\u0307 = r (1) For Charlie ASV, the yaw torque control is described with \u03c4N = n2\u03b4 where \u03b4 is the rudder angle and n is propeller revolution rate. The line following approach is shown in Fig. 2. The aim is to steer the vehicle moving at surge speed ur in such a way that its path converges to the desired line. If \u03b3 is orientation of the line that should be followed, a new parameter \u03b2 = \u03c8 \u2212 \u03b3 (vehicle\u2019s orientation relative to the line) is defined. Having this in mind, the line following equations (2) - (5) can be written, where \u03bd is drift due to external disturbances which is perpendicular to the direction of the desired path. r\u0307 = \u2212kr|r| Ir r|r| + 1 Ir \u03c4N (2) \u03c8\u0307 = r (3) \u03b2\u0307 = r (4) d\u0307 = ur sin\u03b2 + \u03bd (5) The nonlinearities of the line-following model appear in (2) and (5)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003033_s11191-012-9502-4-Figure16-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003033_s11191-012-9502-4-Figure16-1.png",
        "caption": "Fig. 16 Surface Force Apparatus, SFA (Richetti and Drummond 2005, p. 235)",
        "texts": [
            "1 New Instruments of Observation and Measurement The new instruments were the surface force apparatus (SFA), the scanning tunnelling microscope (STM), the atomic force microscope (AFM), the lateral or friction force microscope (LFM or FFM), and the quartz crystal microbalance (QCM). The SFA (Surface Force Apparatus) was already developed in 1969 to measure the normal forces between two surfaces dipped into a liquid or a gas and it had been used for measurements of friction since 1973 but only later was it improved and became very precise. Two thin layers of an adapted material (Fig. 16), usually mica, were stuck to the surface of two glass or silicon cylinders (diameter approximately 1 cm), having their axes at a right angle, so forming a small circular zone of contact of radius between 10 and 40 lm. On mica surfaces it was possible to deposit layers of other materials and lubricant to be studied. The whole system could be plunged into a gaseous or liquid medium, at a fixed temperature and pressure, inside a closed box. The frictional force and the normal force could be measured with sensitivity up to 10 nN and the lateral displacements with a resolution of a micrometer"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003635_0954410012450550-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003635_0954410012450550-Figure1-1.png",
        "caption": "Figure 1. Planar relative motion of missile and target.",
        "texts": [
            " Thus, it is more practical to design a dimension reduction observer-based guidance law accounting for second-order dynamics of the missile autopilot. In this article, such a guidance law is designed using the dynamic surface control method. In the design, no derivatives of the LOS angular rate occur in the expression of guidance law. The stability of the guidance system is verified. Simulation results are proposed to demonstrate the excellent property of the proposed guidance law. The planar relative motion of a missile and a target is shown in Figure 1. The relative range between the missile and target is denoted by R. The derivative of R with respect to time is denoted as _R. The velocities of the target and missile are denoted by Vt and Vm, respectively. The LOS angle is denoted by q and the derivative of q with respect to time is denoted as _q The flight path angles of the target and missile are denoted by \u2019t and \u2019m, respectively. Assume that the missile and target are point masses moving in the plane and the velocities of the missile and target are constant. Then, the relative motion shown in Figure 1 can be expressed by the following equations _R \u00bc Vt cos\u00f0q \u2019t\u00de Vm cos\u00f0q \u2019m\u00de \u00f01\u00de R _q \u00bc Vt sin\u00f0q \u2019t\u00de \u00fe Vm sin\u00f0q \u2019m\u00de \u00f02\u00de Let VR \u00bc _R and Vq \u00bc R _q. Substituting them into equations (1) and (2) and then differentiating them with respect to time yields _VR \u00bc V2 q R \u00fe aTR aMR \u00f03\u00de _Vq \u00bc VRVq R \u00fe aT aM \u00f04\u00de where aTR and aMR denote the accelerations of the target and missile along the LOS, respectively; aT and aM denote the accelerations of the target and missile normal to the LOS, respectively. Substituting Vq \u00bc R _q and VR \u00bc _R into equation (4) gives \u20acq \u00bc 2 _R R _q 1 R aM \u00fe 1 R aT \u00f05\u00de In practical applications, the thrust of missile cannot be accurately adjusted or there are no thrusts in the terminal guidance processes"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000810_015003-Figure9-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000810_015003-Figure9-1.png",
        "caption": "Figure 9. A schematic diagram illustrating the pasting of an arbitrary polygon into a stretched polygon at a cut-plane. In this example, a polygon K of span s0 in a slab Ls0 is translated and then concatenated at the cut-plane into a stretched polygon, increasing the Z-span of the stretched polygon by s0 + 1.",
        "texts": [
            " There is a fixed f0 such that for any fixed finite positive value of f > f0 there exists a b \u2208 (0, 1/2] such that lim n\u2192\u221e[Z\u2217 n(f ; bn )]1/n = lim n\u2192\u221e[Zn(f ; bn )]1/n = eF(f ). The significance of this result is that it is in particular valid for large values of f > f0. In this section the discussion will be limited to stretched polygons with a force f > f0, where f0 is defined in theorem 3.5. In these circumstances, corollary 3.6 is applicable, and a class of stretched polygons with a positive density of cut-planes determines the limiting free energy in the model. The basic construction in this section is illustrated in figure 9: a polygon K is translated to a cut-plane of a stretched polygon. The stretched polygon is decomposed by deleting the pair of cut-edges in the cut-plane and its components are moved apart to create space for inserting K, possibly rotated by 90\u25e6 about the Z-axis to line up edges which must be concatenated. K is inserted by concatenating it to a sequence of edges in the Z-direction along one of the lines which contained the original cut-edges. When the polygon K is concatenated into a stretched polygon in this way, it maintains all its edges, save one, and it is called an event",
            " Moveover, \u00b5b(K, f ) is concave on [0, 1/2) and right continuous at b = 0. Proof. Suppose that K is the (irreducible) event that a polygon of length k and Z-span s0 occurs. Concavity and right continuity at b = 0 follow from arguments similar to theorem 3.3. It remains to show that \u00b5b(K, f ) = limn\u2192\u221e [Zn(f ; bn , 0K)]1/n < eF(f ). Consider polygons counted by pn(s; bn , 0K). If A is such a polygon of length n, then it has bn cut-planes. Let 0 < \u03b4 < b, and choose \u03b4n cut-planes. Perform the construction in figure 9 at each of the chosen cut-planes. This increases the length of A by k \u03b4n (for some fixed k) and increases the Z-span by (s0 + 1) \u03b4n . The number of cut-planes is also increased by \u03b4n . Thus ( bn \u03b4n ) pn(s; bn , 0K) pn+k \u03b4n (s + (s0 + 1) \u03b4n ; bn + \u03b4n , \u03b4n K). Multiply this by ef s and sum over s:( bn \u03b4n ) Zn(f ; bn , 0K) e\u2212(s0+1) \u03b4n f Zn+k \u03b4n (f ; bn + \u03b4n , \u03b4n K). Take the power 1/n and the limit of the left-hand side as n \u2192 \u221e. The partition function on the right-hand side is bounded above in the limit by the partition function of all stretched polygons, and so the result is that[ bb e(s0+1)\u03b4f ek\u03b4F(f ) \u03b4\u03b4(b \u2212 \u03b4)b\u2212\u03b4 ] lim n\u2192\u221e[Zn(f ; bn , 0K)]1/n eF(f ) after terms have been rearranged",
            " Next, choose n large enough, and \u03c10 small enough, such that lim sup m\u2192\u221e [Zm(f ; \u03c1 K)]1/m < (1 + \u03b5)[Zn(f ; \u03c1n )] 1 n+2 , for the value of \u03b5 above (which is independent of n). This shows that lim sup m\u2192\u221e [Zm(f ; \u03c1 K)]1/m < (1 \u2212 \u03b52) eF(f ) for 0 \u03c1 \u03c10, for some \u03c10 > 0. This completes the proof. Observe that theorem 4.6 states that if K is an event and f > f0, then K will occur with a positive density in almost all polygons of length n, and with positive density with probability 1 as if n \u2192 \u221e. The occurrence of K is in the sense illustrated in figure 9. This result may strengthened as follows: we say that a self-avoiding walk L is a stretched pattern if there exists a stretched polygon which contains L as a sub-walk. We say that a stretched pattern L occurs in a stretched polygon P if P contains a sub-walk which is a translate of L: that is, L is identical to a sub-walk in P. P contains Lm times if m copies of L can be translated onto m distinct sub-walks of P. Define Zn(f ;mL) to be the partition function of stretched polygons of length n at force f which contains exactly m copies of a stretched pattern L"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002335_detc2010-28926-Figure9-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002335_detc2010-28926-Figure9-1.png",
        "caption": "Figure 9. The process of analyzing the series of Assur Graphs composing the determinate truss appearing in Figure 8a.",
        "texts": [
            " Search for an Assur Graph that can be removed and at least one force, which is acting on one of its vertices. 2. Remove this AG, add ground vertices to its outer vertices with ground vertices and calculate the forces in its internal edges. 3. Replace all the ground edges of the removed AG with external forces with the same magnitude and direction forces in its ground edges. 4. Go to 1. An example of applying this analysis process appears in Figures 8 and 9. The determinate truss for which the analysis is applied consists of three triads and one dyad, as appears in Figure 8a. Figure 9 depicts the process of analyzing the determinate truss appearing in Figure 8a, each time an AG is being analyzed. First, the triad (A,B,C) can be removed and since on one of its vertices, vertex B, acts an external force thus this AG is the first to be removed and analyzed (Figure 9b). The inner forces in the three ground edges, (AK), (CD) and (B,K), of the latter AG become external forces that act on the remaining determinate truss: PBK, PCD, PAD, as shown in Figure 9c. This process continues and is applied on the dyad K (Figure 9c), then triad (G,H,I) as shown in Figure 9d, and ends with the analysis of the triad (E,D,F) upon which four external forces act (Figure 9e). 5 Copyright \u00a9 2010 by ASME Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 08/12/2018 Terms of Use: http://www.asme.org/about-asme/terms-of-use 6 Copyright \u00a9 2010 by ASME GRAPHS IN 2D In this section we show that it is possible to derive all the AGs in 2D by only two operations. All the AGs, although there is an infinite amount, are arranged in a very unique order as shown in the map appearing in this section. This map is proved to be complete and sound, i.e., all the AGs appear in this map and all the graphs that appear in the map are AGs"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003398_156855112x629531-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003398_156855112x629531-Figure1-1.png",
        "caption": "Figure 1. Fabrication of micro-braided yarn. (a) Schematic view of a braiding machine. (b) Micro-braided yarn.",
        "texts": [
            " Specimens were molded under various conditions and tested to clarify the effects of molding condition on the mechanical properties of the continuous bamboo fiber reinforced PLA composites. The materials used in this study were a bamboo rayon fiber (Tex: 18.5) and PLA fiber yarns (Tex: 8.2). Micro-braided yarn was fabricated as fiber volume fraction of 50% with a tubular-braiding machine. In the present study, since bamboo rayon fiber was very fine and was tender, four fibers were bundled with a quiller machine and were used in the micro-braiding process. Fabrication and structure of microbraided yarn are illustrated in Fig. 1. Micro-braided yarn obtained was wound on the metallic frame 50 \u00d7 2 times. Then, the micro-braided yarns wound on the frame were placed on the pre-heated mold die and were molded with a hot-press system. Figure 2 shows a schematic view of the molding process. After a given molding cycle, the heating platens were cooled by the water flow in the pipes equipped through the platens. During cooling, molding pressure was applied until the temperature of the platens became 40\u00b0C and the molded pieces were obtained"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000597_978-3-540-73129-0-Figure6.49-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000597_978-3-540-73129-0-Figure6.49-1.png",
        "caption": "Fig. 6.49. Principle of the wireless linear micromotor",
        "texts": [],
        "surrounding_texts": [
            "T. Fukuda [65] has opened the field of miniature magnetostrictive actuators and motors taking advantage of wireless magnetic excitation. He has experimented with two small self-moving linear motors (some of cubic centimeter dimensions) based on a conversion-mode principle. The first linear micromotor, based on magnetostrictive thin films deposited on a 7\u00b5m polyamide film, was built in Japan in 1994 [66]. The 13mm long prototype used a 200Hz vibration induced by magnetostriction to obtain one-way motion at 5mm/s. This is a mode conversion ultrasonic motor (MCUM) according to the Japanese classification of piezoelectric motors. The torsion-based, drift-free microactuator [67], invented by CNRS Grenoble, is basically a unimorph structure composed of a single magnetostrictive film deposited on a passive substrate. The new feature is a square shape maintained by hinges at three corners (Fig. 6.46). The useful displacement due to magnetostriction is obtained at the fourth (free) corner and without thermal displacement. The different deformed shapes are due to the anisotropy of magnetostrictive strains and the isotropy of thermal strains. Modeling with Atila (Fig. 6.47) has permitted the design of appropriate microhinges. Prototypes have been realised by micromachining a Silicon substrate and by depositing a magnetostrictive film by sputtering. Measurements using laser interferometry have confirmed the modeling expectations. Several standing-wave ultrasonic motors (SWUMs), have been designed at Cedrat [68] (Figs. 6.48 and 6.49). A linear motor is a self-moving silicon plate including magnetostrictive film. It is submitted to a 10mT dynamic field produced by an external coil, which may be placed at some centimeters distance from the motor. At resonance, this field excites a flexure mode, producing vibrations in the plate, which in turn induces by friction a motor motion at 10 . . . 20mm/s. A rotating version has been also created (Fig. 6.48b) that uses a slightly different principle [68]: the vibrating rotor is based on a 100\u00b5m thick by 20mm diameter plate with 10\u00b5m deposited magnetostrictive films, which are wireless and excited by a small coil. Typical performance is a rotating speed of 30 rpm and a torque of 1.6\u00b5Nm, with a 20mT excitation field. These examples demonstrate some of the special advantages of magnetostriction, especially the fact that the moving parts are wireless. The disadvantage is the coil, which is difficult to miniaturise because of the field requirements. These considerations are driving the development of films with magnetostriction at low fields. Note that, as these devices are very small, the price of the material is not a problem, and so such actuators could find large scale applications \u2013 for instance in optics, in medicine, or in the automobile industry."
        ]
    },
    {
        "image_filename": "designv11_12_0002472_icelmach.2012.6350085-Figure3-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002472_icelmach.2012.6350085-Figure3-1.png",
        "caption": "Fig. 3. FE meshes.",
        "texts": [
            " Therefore, to accurately calculate the core loss of the IM, it is expected to be essential to consider the skin effect in the electrical steel sheets. However, the skin depth depends on not only the order of the time harmonics but also the permeability that varies with flux density [4]. From these viewpoints, we have developed a combination of the 2-D and 1-D FEMs [4], [7], [8]. In this method, the 1- D FEM along the thickness of the electrical steel sheet is employed for the post calculation of the main 2-D FEM, as shown in Fig. 2. Fig. 3 shows the FE meshes for this analysis. Table III lists the discretization data. First, the electromagnetic field in the IM is analyzed by the multi-sliced 2-D FEM that considers the rotor skew and the inverter carrier; the formulations are expressed, as follows: t cage    A JJJA   121) 1 ( (3) m m m IR dt d V 11 1 1    (4) where A is the magnetic vector potential;  is the permeability considering the nonlinear B-H curve of the laminated core; cage is the conductivity of the secondary conductor by considering the resistance of end-rings; J1 and J2 are the primary and secondary current densities, respectively; V1m and I1m are the primary voltage and current of the m-th phase, respectively;1m is the primary flux linkage; R1 is the primary resistance"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003924_1.4030612-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003924_1.4030612-Figure2-1.png",
        "caption": "Fig. 2 Cross section of bearing and damper",
        "texts": [
            " The displacement vector of each node is (x, y, z, hx, hy, and hz). The shaft is assumed to be axially symmetric and modeled with Timoshenko beam elements for lateral directions (x, y, hx, hy) and a linear element for axial and torsional directions (z, hz). The built on components, such blades, disks, and impellers, are modeled as lumped masses plus the properties of polar and transverse mass moments of inertia. An additional node with 2DOF (xsd, ysd) is added for each SFD to indicate the location of bearing center, which whirls inside the SFD, as shown in Fig. 2. Gyroscopic effects, which are functions of both rotational speed and acceleration, are taken into consideration as well. Bearings or SFDs are treated as nonlinear components, 1Corresponding author. Contributed by the Dynamic Systems Division of ASME for publication in the JOURNAL OF DYNAMIC SYSTEMS, MEASUREMENT, AND CONTROL. Manuscript received November 7, 2014; final manuscript received April 28, 2015; published online June 24, 2015. Assoc. Editor: Yongchun Fang. Journal of Dynamic Systems, Measurement, and Control SEPTEMBER 2015, Vol"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000686_s10409-008-0184-8-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000686_s10409-008-0184-8-Figure1-1.png",
        "caption": "Fig. 1 a A sketch of the rigid body approximation; b Definition of the state variables",
        "texts": [
            " Similar to [2,3], we make the rigid body approximation: the wingbeat frequency of the insect is assumed to be much higher than that of the natural modes of motion of the insect, and the insect is treated as a rigid body of six degrees of freedom (in the present case of symmetric longitudinal motion, only three degrees of freedom), with the action of the flapping wings represented by the wingbeat-cycle-average forces and moment. In [6] it has been shown that the rigid body assumption is applicable to bumblebees. This model of the bumblebee is sketched in Fig. 1a. Let oxyz be a noninertial coordinate system fixed to the body. The origin o is at the center of mass of the insect and axes are aligned so that the x-axis is horizontal and points forward at equilibrium. The variables that define the motion (see Fig. 1b) are the forward (u) and dorsal-ventral (w) components of velocity along x- and z-axis, respectively, the pitching angular-velocity around the center of mass (q), and the pitch angle between the x-axis and the horizontal (\u03b8). A coordinate system oExE yEzE is fixed on the earth; xE-axis is horizontal and points forward. Let c be the vector of control inputs. In the study of bumblebees in hovering and forward flight, Dudley and Ellington [7] observed that freely-flying bumblebees control the longitudinal motion mainly by changes in geometrical angles of attack of the wing and changes in the fore/aft extent of the flapping motion of the wing (see [3] for a description of the flapping motion of the wing)",
            " The general morphological data: m = 175 mg; wing length R = 13.2 mm; c = 4.01 mm, r2 = 0.554R; area of one wing (S) is 53 mm2; body length (lb) is 1.41R; distance from anterior tip of body to center of mass divided by body length is 0.48lb; Iy = 0.213 \u00d7 10\u22128 kg m2. Kinematic data on wing-motion and body orientation at five flight speeds are listed in Table 1 [these data include: , n, \u03c6\u0304, \u03b1d, \u03b1u stroke plane angle \u03b2 and body angle \u03c7 (stroke plane angle is the angle between the stroke plane and the horizontal, see Fig. 1a; body angle is the angle between the longitudinal axis of the body and the horizontal, see Fig. 1b)]. As described in [3], , n, \u03b2 and \u03c7 were provided by Dudley and Ellington [7], while \u03c6\u0304, \u03b1d and \u03b1u were determined by using the equilibrium conditions, i.e. Eqs. (6)\u2013(8). Using the flight data, U ,m+, I + y and g+ have been computed as U = 4.4 m/s, m+ = 91.45, I + y = 9.224, g+ = 0.01485. 2.3 The control derivatives The equilibrium flight and the stability derivatives have been determined in our previous work on stability study [3] (i.e. Eqs. (6)\u2013(8) are satisfied and A matrix in Eq. (18) is determined)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003142_1.3616922-Figure19-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003142_1.3616922-Figure19-1.png",
        "caption": "Fig. 19 Diagram showing geometrical relationships used in determining the point of load concentration K and its projection to the mean section K' with respect to the pitch line through point P",
        "texts": [
            " This differs from the previous assumption that the area was a symmetrical ellipse. Note that the center of pressure is toward the heel (outer end of the tooth). Although there is a distributed load, for simplicity a point is chosen to represent a concentrated load replacing the distributed load. Based on the etching tests and other similar studies by the present author it was determined that point K will lie along the line of contact at a distance j toward the heel of the fc tooth from the center of the line of contact 0'. See Fig. 19- Note that point K does not necessarily lie in the mean section Y Y , but will usually lie toward the heel. Since calculations are normally made in the mean section, point K will project to point K'. If it is now assumed that point K ' represents the critical point on the tooth, then the distance Zo in the mean normal section measured in the surface of action from the pitch line to the critical point K', Fig. 19, will be: where Z = length of action in transverse plane V = PNHIO ipb = base spiral angle F = face width k = A po = NP N 0 3.2Arg + 4.0iVP N0 - NP (22a) (22 b) distance along line of action in mean normal section from gear addendum circle (outside circle) to pitch point number of pinion teeth number of gear teeth \u00b1 FZr^ _ Apf l (22) On curved-tooth bevel and hypoid gears use the plus ( + ) sign for the concave side of the pinion tooth (convex side of the gear tooth) and the minus ( \u2014 ) sign for the convex side of the pinion tooth (concave side of the gear tooth) in equation (22)",
            " Recent work [6, 7] indicates that this formula may need revision to include not only the velocity and surface finish, but also the load and the lubricant viscosity as variables, since these will affect the film thickness and therefore the shear rate in the lubricant film. Lubricant testing is being pursued at the Gleason Works in order to expedite this revision. Scoring Formula By making substitutions in equation (20) it is possible to solve for the value of A7' at any point designated by the distance/, Fig. 19, from the center of the surface of action. Since the critical point on the tooth will be that point at which A T is a maximum, it is now necessary to carry out the foregoing calculations for several values of / . It is then possible to plot a graph similar to that shown in Fig. 23. On a computer it is a relatively simple task to write a program which will determine the maximum value of AT. 1 2 2 / A P R I L 1 9 6 7 Transactions of the AS ME Downloaded From: http://tribology.asmedigitalcollection"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003292_e2013-01943-7-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003292_e2013-01943-7-Figure2-1.png",
        "caption": "Fig. 2. Actuated truss structure.",
        "texts": [
            " This approach has been successfully employed in several applications, ranging from underwater robotic vehicles [26,27] and electro-hydraulic actuated systems [28] to chaos control in a nonlinear pendulum [29,30]. In this work, the proposed control problem is to ensure that, even in the presence of modeling inaccuracies and external disturbances, the state vector x = [x, y] will be stabilized in a desired state xd = [xd, yd], i.e. the error vector x\u0303 = [x\u0303, y\u0303] = [x\u2212 xd, y \u2212 yd]\u2192 0 as \u03c4 \u2192\u221e. In order to ensure the stabilization, a linear actuator is supposed to be installed vertically at the junction between the two bars, as illustrated in Fig. 2. The combination of linear actuators with shape memory elements enables the development of variable geometry trusses that also have the ability of self-attenuate their vibration levels. This kind of adaptive structure could be very useful, for example, in aerospace applications. On this basis, the related control variable u must be added to the equation of motion (6), which for control purposes could be simply rewritten as x\u2032 = y y\u2032 = f + d+ u (7) where u is the control action, d = \u03b3 sin(\u03a9\u03c4) is an external disturbance assumed to be unknown, and f = \u2212\u03bey+x{\u2212[(\u03b8\u22121)\u22123\u03b12+5\u03b13]+ [(\u03b8\u22121)\u2212\u03b12+\u03b13](x2+ b2)\u22121/2\u2212 [3\u03b12 \u2212 10\u03b13](x2 + b2)1/2 + [\u2212\u03b12 + 10\u03b13](x2 + b2) + 5\u03b13(x2 + b2)3/2 \u2212 \u03b13(x2 + b2)2}"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001703_1.47212-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001703_1.47212-Figure2-1.png",
        "caption": "Fig. 2 Banked configuration.",
        "texts": [
            " According to this assumption, CNwB 2 Sw Z b a fCN w0 f y CN w2 f y 3 CN w4 f y 5 gD dy (9) Since the horizontal panel-normal force is an odd function of the angle of attack, only odd powers of the effective angle of attack are used, making Eq. (9) applicable to negative angles of attack. It is implicitly assumed that the wing panels have no camber, so the normal force at a negative angle of attack is the negative of that at the same positive angle. If the configuration is now banked in the positive direction an angle \u2019, as depicted in Fig. 2, the unperturbed angle of attack seen by the strips on panel 2 is cos\u2019, and the normal-force coefficient generated at low angles of attack by panel 2, projected in the crossflow direction indicated by , is CNwBh2i cos\u2019 Sw Z b a CN w0 cos\u2019 f y g y sin\u2019 D dy (10) where the effective angle of attack seen by the strips on panel 2 is cos\u2019 f y g y sin\u2019 . The pattern of interference in Eq. (10) was taken from the SBT [2]. As before, Eq. (10) can be particularized for the SBT, and g y can be approximated by g y sbt, where g y sbt 4 2 p y4 a4 3=2 ARy5CN w0sbtD F 1; k F 2; k 0 (11) F is the elliptic integral of the first kind, and cos 1 b y2 a2 y b2 a2 ; cos 2 b y2 a2 y b2 a2 k2 y 2 a2 2 2 y4 a4 ; k02 y 2 a2 2 2 y4 a4 The extension of Eq",
            " If we expand those functions as power series of p and T , we can write fEij fij0 fij2 pD V 2 fij4 pD V 4 fOij fij1 pD V fij3 pD V 3 fij5 pD V 6 lOij lij10 pD V lij30 pD V 3 lij01 T lij03 3T lij21 pD V 2 T lij12 pD V 2T lEij lij00 lij20 pD V 2 lij02 2T lij04 4T lij11 pD V T lij22 pD V 2 2T (B4) In the preceding coefficients, the subscripts indicate the powers of , , pD=V, and T , respectively. The force coefficients fij are assumed to be independent of T . In the force expansion of Eq. (B2), the roll-dependent terms can be regarded as Magnus terms. The remaining terms, which correspond to coefficients fij0, yield the following expansion: F f100 f210 2 f030 2e i4\u2019 f500 4ei4\u2019 f320 4 f140 4e i4\u2019 f610 6ei4\u2019 f430 6 f250 6e i4\u2019 f070 6e i8\u2019 (B5) Introducing the angle of attack i (Fig. 2), the complex forceF can be broken down into the normal and side forces: Fz= N= f100 f210 2 f320 4 f430 6 cos 4\u2019ff030 2 f140 f500 4 f250 f610 6g cos 8\u2019ff070 6g (B6) Fy= sin 4\u2019ff030 2 f140 f500 4 f250 f610 6g sin 8\u2019ff070 6g (B7) In a similar manner, the Magnus terms can be separated into normal and side contributions. If quadratic and higher powers of roll rate are disregarded, the following expansion is obtained: Fyp pD V f101 f211 2 f321 4 f431 6 cos 4\u2019ff031 2 f141 f501 4 f251 f611 6g cos 8\u2019ff071 6g (B8) Np pD V sin 4\u2019ff031 2 f141 f501 4 f251 f611 6g sin 8\u2019ff071 6g (B9) For the roll moment, by combining the roll expressions of Eqs"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001512_s11668-010-9398-8-Figure7-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001512_s11668-010-9398-8-Figure7-1.png",
        "caption": "Fig. 7 Schematic view of the gear test rig [23]. 1,2: gears box; 3: load coupling; 4: DC motor shaft; 5,6: pinion and gears (test gears)",
        "texts": [
            " The plastic gears were produced by using a hobbing machine. Cooling water was used for all of the operations to limit the temperature developed while the plastic gear was manufactured. The tooth surface roughness value for both of the plastic gears and the AISI 8620 was between Ra 0.6\u20130.8 lm. A Klingenberg PFS-600 Gear Lead/Profile test apparatus was used to check the contact errors for all of the test specimens. Gear Test Procedure A test of the load-carrying capacity of the gear pairs was performed on a FZG test machine (Fig. 7). The cover on the gear box was opened. The test conditions are shown in Table 3. Initially, the gears were run for 10 min at a pinion rotation speed of 200 rpm and a low contact load (2 N/mm) to smooth the original machine-finished teeth surfaces. The experiment was not stopped unless a tooth was broken or sudden thermal damage occurred because of a rapid decrease in the flank temperature once the gear stopped. The temperature of the tooth surface was measured using a non-contact-type temperature sensor described in a previous study (see Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000407_s00170-008-1708-x-Figure4-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000407_s00170-008-1708-x-Figure4-1.png",
        "caption": "Fig. 4 Schematic diagram of tube bulging using a solid bulging medium",
        "texts": [
            " In order to simulate bulge forming of tubes using a urethane rod, a finite element model was built using commercial finite element software ABAQUS/Explicit. To be able to analyse the bulge forming process, simulation was carried out using an explicit integration scheme. The motivation for this approach is the extremely simple nature of the explicit formulation over conventional implicit formulations, allowing very inexpensive time steps and greatly simplified contact treatment. For many forming processes, implicit solvers often break down due to convergence or contact algorithm failure. Figure 4 shows a schematic diagram of the setup and the shape at the last stage of the conical bulging using a solid bulging medium. A urethane rod of Shore Hardness 95A and diameter 32 mm was used to bulge an annealed copper tube of diameter 42 mm and wall thickness of 1.2 mm. Due to the axisymmetry of the process, the finite element meshes were generated on the half cross-section of die, tube and flexible medium. We used CAX4R elements for meshing the parts. CAX4R is a four-node bilinear axisymmetric quadrilateral, reduced integration solid element"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003713_978-1-4614-2419-2-Figure6.1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003713_978-1-4614-2419-2-Figure6.1-1.png",
        "caption": "Fig. 6.1 3-DOF typical aeroelastic section with control surface (Adapted from Tang and Dowell [11])",
        "texts": [
            " This work presents the investigation of an electromechanically coupled airfoil section with a control surface. A base configuration (a nonlinear electromechanically coupled 2-DOF airfoil) was experimentally verified in Sousa et al. [7]. The electromechanical coupling is introduced in the plunge DOF. The investigation of the system behavior with the variation of the uncoupled natural frequencies (of the degrees of freedom considered in this work) shows that an airflowexcited energy harvester can be optimized for low-speed envelope or for maximum power output. Figure 6.1 shows the schematic of a 3-DOF aeroelastic typical section. The plunge, pitch and control surface displacement variables are denoted by h, a and b, respectively. When subscript, these variables denote a DOF. The plunge displacement is measured at the elastic axis (positive in the downward direction) and the pitch angle is measured at the elastic axis (positive in the clockwise direction). The control surface angle is measured from the typical section chord line (positive in the clockwise direction)",
            " The rigidities per unit span in the plunge, pitch and control surface DOFs are denoted by kh, ka and kb, respectively. Dissipative effects are also considered in the system. The damping coefficients per unit span of the plunge, pitch and control surface DOFs are denoted by dh, da and db, respectively. The coupled equations of motion, (6.1a)\u2013(6.1d), are presented in terms of dimensionless parameters. The dimensional equations can be obtained from the Lagrange\u2019s equation applied to the electromechanically coupled lumped parameter model shown in Fig. 6.1 [7, 9]. The dimensionless time is defined by t \u00bc oht, where oh is the uncoupled natural frequency of vibration in the plunge DOF and t is the dimensional time. The prime (0) denotes derivative with respect to t. The dimensionless plunge displacement is defined by h \u00bc h b= . 54 V.C. de Sousa et al. r2aa 00 \u00fe r2b \u00fe \u00f0c a\u00dexb b00 \u00fe xa h 00 \u00fe xaa 0 \u00fe fa\u00f0a\u00de a as\u00f0 \u00de \u00bc Ma (6.1b) r2b \u00fe \u00f0c a\u00dexb a00 \u00fe r2bb 00 \u00fe xb h 00 \u00fe xbb 0 \u00fe fb\u00f0b\u00de b bs\u00f0 \u00de \u00bc Mb (6.1c) c v0 \u00fe 1 l v\u00fe w h0 \u00bc 0 (6.1d) The dimensionless radii of gyration are ra \u00bc \u00f0Ia mb2 \u00de1=2 and rb \u00bc \u00f0Ib mb2 \u00de1=2, where Ia and Ib are the airfoil and control surface moments of inertia, respectively"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001637_s11044-010-9215-x-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001637_s11044-010-9215-x-Figure2-1.png",
        "caption": "Fig. 2 Parameterization of curved track",
        "texts": [
            "2 Parameterization of curved track In existing contact geometry analysis procedures, a tangent track model is used and contact points obtained using this assumption is used for curved negotiation as well [6]. This assumption can be valid when the radius of curve is large enough as compared to the track gauge. In order to relax this assumption, a curved track model is developed in this investigation for accurately predicting the location of contact during curved negotiations. To this end, as shown in Fig. 2, the profile coordinate system of rail r is defined for right and left rails as XrkY rkZrk . Using this profile coordinate system, the longitudinal and lateral surface parameters given by srk 1 and srk 2 are introduced as shown in Fig. 2. With these two parameters, the cross-sectional shape of rail can be defined by f k(srk 1 , srk 2 ). If the shape of the rail can be assumed to remain constant along the arc-length of the rail, one can have f k = f k(srk 2 ). Assuming that the rail is rigidly fixed to the global coordinate system (ground), the global position vector of the contact point on the rail can be defined as rrk = Rrk + Arku\u0304rk k = 1,2,3 (5) where Rrk defines the location of the origin of the rail profile coordinate system with respect to the global coordinate system and is given as function of the longitudinal surface parameter srk 1 ; Ark defines the orientation of the profile coordinate system that is also function of the longitudinal surface parameter srk 1 ; and u\u0304rk is the location of the contact point defined with respect to the profile coordinate system"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001246_j.triboint.2009.05.001-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001246_j.triboint.2009.05.001-Figure1-1.png",
        "caption": "Fig. 1. Description of a radial piston hydraul",
        "texts": [
            " Although these studies have helped in getting an insight into the mechanisms of scuffing/seizure yet these do not explain the behaviour of the complete tribological system. The present work therefore focuses on investigating the seizure of hydraulic motors with a view to establishing the seizure limits and understanding the mechanism of seizure. In all, 24 motors were tested under extreme operating conditions of speed, pressure and lubrication with low viscosity fluids. A radial piston hydraulic motor is described in Fig. 1. As can be seen, there are two sliding interfaces and one rolling interface involved in converting the energy from the pressurized oil flow into the rotating motion of the motor shaft. These interfaces are: Interface 1: Sliding interface between cylinder and piston. Interface 2: Sliding interface between roller and piston. Interface 3: Rolling interface between roller and cam. A port plate distributes the oil to the cylinders and the timing is made so that the cylinder will have high pressure on one side of the cam top and low pressure on the other"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003520_j.euromechsol.2011.11.003-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003520_j.euromechsol.2011.11.003-Figure1-1.png",
        "caption": "Fig. 1. T2R1-type parallel manipulator with decoupled and bifurcated planar-spatial motion of the moving platform: constraint singularity (a), branch with planar motion (b) and branch with spatial motion (c); limb topology PkRkRkRtR-PtRkRtRPkRkRS.",
        "texts": [
            " In the two distinct branches, the parallel mechanism is characterized by M\u00bc SF \u00bc 3 and (RF)\u00bc(v1, v2,ua) or (RF)\u00bc(v1, v2,ud). In the constraint singularity, both connectivity of moving platform and mobility of parallel manipulator increase their instantaneous values. More details about the behaviour of parallel manipulators in a constraint singularity and the associated structural parameters can be found in Gogu (2009c). To simplify the notations of the links eGj (j \u00bc 1, 2, 3 and e \u00bc 1,., n) by avoiding the double index in Fig. 1, we have denoted by eA the links belonging to the limb G1\u00f0eAheG1 \u00de, by eB, and eC the links of the limbs G2\u00f0eBheG2 \u00de and G3\u00f0eCheG3 \u00de: The moving platform 6 of the parallel manipulators F ) G1 G2 G3 with decoupled motions in Figs. 1e4 is connected to the fixed base 1 by three simple limbs actuated by three linear motors mounted on the fixed base. No closed loops exist inside a simple limb, that is irl \u00bc rl \u00bc 0. Just revolute R, prismatic P and spherical S joints are used in these solutions, in which two consecutive revolute and prismatic joints have parallel (k) or perpendicular \u00f0t\u00de axes/directions"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000161_iccas.2007.4406826-Figure3-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000161_iccas.2007.4406826-Figure3-1.png",
        "caption": "Fig. 3 Inverted pendulum model to describe humanoid center of mass.",
        "texts": [
            " After all, in order to have exact CoM estimates by adopting the discrete Kalman filter, it is prerequisite to construct an equation of motion to address the change of CoM position when the humanoid behaves and to utilize available sensors whose measurements can be converted to the value of CoM. To describe the walking motion of humanoids simply, the inverted pendulum model with variable length is popularly used. If we assume that the total mass of a humanoid ( ) is concentrated in the CoM and the vertical position is fixed, we have the following relationship between ZMP and CoM: (1) where the positions of ZMP ! and CoM \" # $ % $ ! are defined with respect to the global coordinates system shown in Fig. 3. Under the assumption of zero-order-holding samplings, above continuous equations can be discretized as& (' ) *! + (' ) *! -, &/. 0 132 5476 8 9: ; 13< = 2 5476 8 476 13< = 2 5476 8 . 0 132 5476 8 , & (' + (' >,) & *? . 0 132 5476 8 @476 13< = 2 5476 8 -, (' (2) for A -axis, where 4B6C ED FG and the same form is given for H -axis. While, using the torque outputs JI! 3IK L from F/T sensors at ankles and a suitable coordinate conversion procedure, the position of ZMP on walking plane can be determined at discrete time ' as (' M N IK O (' O (' M IK (' QP (3) In fact, this inverted pendulum model is useful to generate the CoM trajectory according to the given ZMP patterns"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003954_213548-Figure9-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003954_213548-Figure9-1.png",
        "caption": "Figure 9: The figure is showing the temperature field in the bearings in different angular positions with different inlet diameters.",
        "texts": [
            "More of the inlet channel is nearer to the \u201cwarm zone\u201d (the minimum thickness zone in the convergent-divergent channel) while more of the maximum temperature has decreased. On the other hand, the average temperature increases because the capability of the inlet channel walls is progressively decreasing. In fact, the average temperature, Figure 8(b), increases when the inclination of the channel is higher than 20\u2218 and when the channel is near the warm zone. While \ud835\udf03 \ud835\udc60 is increasing, the average temperature has aminor decrease, due to the inflow from the outlet, of fresh fluid, which is set to the mixing temperature via UDF (Figure 9). The fact that the optimal diameter rests on the edge of the research spacemeans that themultidimensional objective function has its global minimum on the boundary of this same space. If the research space could be increased, another global minimum would be found. This fact confirms the theory on which this optimization process is designed: only an objective able to take into account several decision variables and fuzzy limits would be able to solve the problem of by guest on June 23, 2016ade.sagepub"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000846_iros.2008.4650783-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000846_iros.2008.4650783-Figure1-1.png",
        "caption": "Fig. 1. Model of planar underactuated biped robot with semicircular feet and torso",
        "texts": [
            " However, it has not been investigated how the bifurcation changes the gait efficiency. Our simulation results show that period-doubling bifurcation also occurs in accordance with the method this paper proposes. We then discuss whether the gait efficiency improves or worsen with respect to the bifurcation. This section describes the basic definitions and analyzes the effect of the upper body incorporating the BHM as a counterweight. This paper deals with a planar underactuated biped model with semicircular feet and a torso as shown in Fig. 1. A bisecting hip mechanism (BHM) [2], which is a mechanism to bisect the relative hip-joint angle with respect to the torso passively, is used to connect the torso with the biped so as not to destroy the robot\u2019s natural dynamics. By the synergistic effect of BHM, we can generate a dynamic bipedal gait efficiently without having to maintain the torso\u2019s posture 978-1-4244-2058-2/08/$25.00 \u00a92008 IEEE. 2934 actively [3]. Semicircular feet are also very effective for generating an efficient dynamic gait [4][5]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001061_09544062jmes1748-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001061_09544062jmes1748-Figure1-1.png",
        "caption": "Fig. 1 Film profile and plug flow",
        "texts": [
            " Isothermal and thermal analyses of grease-lubricated concentrated point contacts are shown in this article. The formulations reported here are generic. For grease lubrication, a modified form of Reynolds equation should be derived to take into account the flow behaviour of pseudo-plastics and the resulting plug formation and flow. The direction of entraining motion is assumed to coincide with the x coordinate, while the z-coordinate remains normal to the bounding contacting surfaces with the origin of the coordinate centre being coincident with the centre of the oil film (see Fig. 1). Thus, the y-direction represents the side leakage. As in the usual Reynolds assumption, inertial and body forces are neglected. The shear stress \u03c4 is considered to be a function of the z-direction variations only, meaning that the oil film is nearly parallel. The shear stress and plastic viscosity only alter with pressure and in the x- and y-directions. Thus, the force balance on an element of fluid yields \u2202\u03c4zx \u2202z = \u2202p \u2202x (1) and \u2202\u03c4zy \u2202z = \u2202p \u2202y (2) Proc. IMechE Vol. 224 Part C: J. Mechanical Engineering Science JMES1748 at UNIVERSITE DE MONTREAL on June 14, 2015pic",
            "comDownloaded from If pressure variation across the thin film is ignored, then integrating the above equations gives \u03c4zx = z \u2202p \u2202x (3) and \u03c4zy = z \u2202p \u2202y (4) The Herschel\u2013Bulkley equation is one of the more realistic constitutive models for grease, described by a three-parameter rheological model as [13] \u03c4 = \u03c40 + \u03c6|D|n (5) When the local shear stress is below the yield stress \u03c40, the Herschel\u2013Bulkley fluids are deemed to behave as rigid solids. Once the yield stress is exceeded, however, the Herschel\u2013Bulkley fluids flow according to a non-linear constitutive relationship, either as a shearthickening fluid or as a shear-thinning one. According to the model a central flow region occurs in which \u03c4 < \u03c40, which is enclosed by regions of shear flow (if \u03c4 > \u03c40). If the plug flow region is of thickness hp as shown in Fig. 1, then \u03c4zx = hp 2 \u2202p \u2202x (6) \u03c4zy = hp 2 \u2202p \u2202y (7) Rearranging the Herschel\u2013Bulkley equation (5) gives \u03c6 ( du dz )n = ( z \u2212 hp 2 ) \u2202p \u2202x (8) and substituting for (1/n) = m and (\u2202p/\u2202x) = (2\u03c40/hp) gives du dz = ( z \u2212 hp 2 )m ( 2\u03c40 \u03c6hp )m (9) Since pressure, and therefore the yield stress and plastic viscosity, are assumed to be a function of the x- and y-coordinates only, these equations can be integrated directly to yield the velocity gradients u = ub + ( 2\u03c40 \u03c6hp )m 1 (m + 1) \u00d7 [( z \u2212 hp 2 )m+1 \u2212 ( h 2 \u2212 hp 2 )m+1 ] (10) where for plug flow z = hp 2 It is convenient and reasonable to consider grease as a Bingham fluid; thus m = n = 1"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001246_j.triboint.2009.05.001-Figure13-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001246_j.triboint.2009.05.001-Figure13-1.png",
        "caption": "Fig. 13. Illustration of frictional heat generation. Area marked 1 will be heated from interface 1 whereas area marked 2 will be heated from interface 2. The wear pattern is normally located at \u2018a\u2019. Friction heat from interface 2 will shift the contact pattern to \u2018b\u2019 or even to \u2018c\u2019 if the friction heat from interface 2 is extremely high. N is the normal force from the cam ring.",
        "texts": [
            " During stage 1 there is only an increase in friction without scuffing damage. In stage 2, scuffing initiates and during stage 3 scuffing propagates to several pistons. Stages 2 and 3 are associated with the initiation of scuffing. However, stage 1 is mainly associated with the system level of understanding described in the Introduction. Since this is the trigging event in the seizure process, it will be discussed in more detail. Observations from stage 1 are: (1) Misaligned wear pattern between piston and cylinder. Normal contact is marked \u2018a\u2019 in Fig. 13, the misaligned wear pattern is marked \u2018b\u2019. (2) Increased leakage. (3) Indication of boundary lubrication between roller and piston. Observations from stage 2 are: (1) When scuffing initiates at interface 1, it originates from the same area as the misaligned wear pattern in stage 1. (2) When there have been high friction power at interface 2, for example if scuffing has initiated there, the area of contact pattern at interface 1 is even further down, marked \u2018c\u2019 in Fig. 13. Pistons that are close to scuffing or have initiation of scuffing seem to have a misaligned wear pattern between piston and cylinder and the distance from where it should normally be is longer when the friction heat from the roller-piston contact is higher. The friction heat in the piston comes from interfaces 1 and 2, Section 2.1. According to Fig. 13, interface 1 heats up region 1 and interface 2 heats up both regions 1 and 2. Thermal expansion of region 1 will only change the shape of the piston slightly but thermal expansion in region 2 may change the shape of the piston and the contact area will move downwards in Fig. 13. This could explain the occurrence of the misaligned wear pattern. Also, it was verified through piston rig tests that increase of friction and leakage, which is associated with stage 1, can occur without interface 1. The initiation of stage 1 in the seizure process seems to originate from the piston\u2013roller interface. Two possible explanations for the increase of friction at interface 2 are: The hydrodynamic oil film is ruptured between roller and piston due to lower viscosity when losses reach a certain level"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001016_icems.2009.5382792-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001016_icems.2009.5382792-Figure2-1.png",
        "caption": "Fig. 2 shows the analyzed model of an IPM motor, which is 1/48 of the whole region because of the symmetry. The length of axial direction in the analyzed region is the same as half the height of the permanent magnet. The material of the rotor and stator core is 50A470. The electric conductivity of the permanent magnet is 694,444 S/m.",
        "texts": [
            "0 0 60 120 180 240 300 360 sinusoidal 9 pulse -1.0 -0.5 0.0 0.5 1.0 0 60 120 180 240 300 360 sinusoidal 27 pulse -1.0 -0.5 0.0 0.5 1.0 0 60 120 180 240 300 360 sinusoidal 27 pulse 0.0 0.5 1.0 0 90 180 270 360 carrier1 carrier2 command value P W M(9 pulse) 0.0 0.5 1.0 0 90 180 270 360 carrier command value P W M(9 pulse) III. ANALYZED MODEL AND CONDITIONS The IPM motor driven by the 2- or 3-level PWM inverter with the several carrier frequencies is computed and compared with that by the sinusoidal waveform using the 3-D FEM. Fig. 2. Analyzed model. Fig. 3. 3-D finite element mesh. 2- and 3-level PWM inverter [6]. Fig. 4(i) shows the phase voltage of 9 pulse modulation. Fig. 4(ii) shows the line voltage of Fig. 4(i). The line voltage of 27 pulse modulation is also shown in Fig. 4(iii). The sinusoidal waveform, whose RMS value is the same as the PWM waveforms, is also drawn in the same figure. TABLE I ANALYSIS CONDITION. Rotation speed (rpm) 935 Frequency of the power supply (Hz) 46.75 Number of coil turn 7 Coil resistance (p"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001349_aim.2009.5229824-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001349_aim.2009.5229824-Figure1-1.png",
        "caption": "Figure 1. The general spatial manipulator as the typical 6-6 hexapod",
        "texts": [],
        "surrounding_texts": [
            "Index Terms\u2014Parallel robot, Gough platform, spatial parallel manipulator, forward kinematics, position based model, genetic algorithms, mutation-based operators, elitist selection, real root isolation.\nI. INTRODUCTION\nThe truly parallel manipulator is defined as an hexapod constituted of a fixed base, a mobile plateform where the end-effector is mounted and six kinematics chains, fig. (1). The platforms are attached through universal or ball joints,\nrespectively with 2 or 3 DOFs. Each kinematics chain contains one 1 DOF prismatic joint.\nRonga, Lazard and Mourrain have proven that the general 6- 6 hexapod FKP has 40 complex solutions using respectively Gro\u0308bner bases, Chern classes of vector bundles and explicit elimination techniques, [1]\u2013[3]. From en engineering pointof-view, the significant issue is the one of real solutions since they correspond to effective manipulator postures. The number of real solutions is always equal or less than the number of complex ones. Fast numerical approaches usually implement Newton\u2019s method, however it is sometimes plagued by Jacobian inversion problems and numerical instabilities. Resultant or dialytic elimination might add spurious solutions, [4]. Homothopy methods are prone to miss some solutions, [4]. In the majority of parallel manipulator cases, the FKP is a difficult problem, [13]. Therefore, this justifies the implementation of another method which could find solutions numerically and rapidly.\nThe genetic algorithms introduced by Holland have evolved significantly in order to suit real-world optimization challenges faced by engineers, [5]. Evolutionary algorithms have been applied for solving the FKP of simple parallel manipulators. However, the later is a classical problem of finding all solutions to a non-linear equation system, whereas GAs solve optimization problems. Hence, there exists a difficulty to derive an optimization problem from a root finding one. In the first, we are only interested by one maximum in the later, we calculate all solutions. A real-coded genetic algorithm (RCGA) was proposed, [7] integrating crossover and mutation operators inspired by operators used in binary GA. It was reported that the GA method is more time consuming than NewtonRaphson\u2019s method. However, it was shown that the domain in which the GA will converge to a solution is larger allowing process launch with a more distant initial guess. Recently, genetic algorithms have succeeded to solve the 3-RRR, [8], and the SSM Gough platform, [9], [10]. However, in [9], one binary coded genetic algorithm is implemented to find one solution of the SSM manipulator. Solving the FKP for the general Gough platform has never been yet attempted.\nIn this work, the RCGA uses Wright\u2019s heuristic crossover operator with different mutation operators, [11]. The Pivot Mutation operator is introduced. The RCGA uses roulette\n978-1-4244-2853-3/09/$25.00 \u00a92009 IEEE 1637",
            "wheel selection combined with the elitist strategy in order to avoid oscillation during the search. Furthermore, as a first, the general Gough platform FKP results obtained by the RCGA are verified with the exact ones obtained from a proven Gro\u0308bner based method implemented on computer algebra [12].\nThis article is presented as follows: Section I addresses the issues involved in kinematics modeling of the general Gough platform and the equation solving problem conversion to an optimization problem. Section II reviews prominent and recently proposed real-coded genetic operators and furthermore gives details of the proposed Pivot Mutation operator along with other operators used in this study. Section 3 presents the results obtained on a 6-6 hexapod with 16 real solutions with conclusion and directions for future work.\nAny manipulator is characterized by its mechanical configuration parameters and the posture variables. The configuration parameters are thus OAO, the base attachment point coordinates in O (the base reference frame), and CB|C , the mobile platform attachment point coordinates in C (the mobile platform reference frame). The kinematics model variables are the joint coordinates and end-effector generalized coordinates. The joint variables are described as li, the prismatic joint or linear actuator positions. The generalized coordinates are expressed as \u2212\u2192 X , the end-effector position and orientation.\nThe kinematics model is an implicit relation between the configuration parameters and the posture variables, F ( \u2212\u2192\nX , L,OA|O ,CB|C ) = 0 where L = {l1, . . . , l6}. This article shall only investigate the forward kinematics problem (FKP ), Fig. 2. Usually the inverse kinematics problem is required to model the FKP and is defined as: given the generalized coordinates of the manipulator end-effector, find the joint positions.\nAccordingly, the forward kinematics problem is defined as: given the joint positions, find the generalized coordinates of the manipulator end-effector.\nB. Vectorial formulation of the basic kinematics model\nContaining as many equations as variables, vectorial formulation constructs an equation system, [14], as a closed vector cycle between the following points: Ai and Bi, kinematics chain attachment points, O the fixed base reference frame and C the mobile platform reference frame. For each kinematics chain, an implicit function \u2212\u2192\nAiBi = U1(X) can be written\nbetween joint positions Ai and Bi. Each vector \u2212\u2192 AiBi is expressed knowing the joint coordinates \u2212\u2192\nli and X giving\nfunction U2(X, \u2212\u2192 L ). The following equality has to be solved: U1(X) = U2(X, \u2212\u2192 L ). The distance between Ai and Bi is set to Li. Thus, the end-effector position X or C can be derived by one platform displacement \u2212\u2192\nOC |O and then one platform\ngeneral rotation expressed by the rotation matrix R. Vectorial formulation 2 evolves as a displacement based equation system using the following relation :\n\u2212\u2192 AiBi|O = \u2212\u2192 OC |O + R \u2212\u2192 CBi|C \u2212 \u2212\u2192 OAi|O (1)\nFor each distinct platform point \u2212\u2192\nOBi|O with i = 1, ..., 6,\neach kinematics chain can be expressed using the distance norm constraint, [15]:\nL2 i = ||AiBi|| 2 (2)\nIn 3D space, any rigid boby can be positioned by three distinct points. Every variable have then the same units and their range is equivalent leading to same weight. Hence, the rotation impact is included into the point parameters and made equivalent to the translation impact. The main disadvantage is the unknown number exceeding the end-effector DOF number, [16].\nThe three platform distinct points are usually selected as the three joint centers B1, B2, B3. The nine variables are set as : \u2212\u2192 OBi|O = [xi, yi, zi] for i = 1 . . . 3. To simplify calculations, one reference frame Rb1 is precisely located on B1. The unit vectors u1, u2 and u3 represent the new frame axes and are defined by:\nu1 =\n\u2212\u2192\nCB1CB2|O\n|| \u2212\u2192 CB1CB2|O || , u2 =\n\u2212\u2192\nCB1CB3|O\n|| \u2212\u2192 CB1CB3|O || , u3 = u1 \u2227 u2\n(3)\nKnowing that the platform is supposed infinitely rigid, any platform point M can be expressed\n\u2212\u2192\nB1M = aMu1 + bMu2 + cMu3 (4)\nwhere aM , bM , cM are constants in terms of these three points. Hence, in the case of the IKP , the constants are noted aBi , bBi , cBi , i = i . . . 6 and can explicitly be deduced from CBC by solving the following linear system of equations :\n\u2212\u2192\nB1Bi|Rb1 = aBi u1 + bBi u2 + cBi u3 , i = 1 . . . 6. (5)\nUsing the relations, equ. (5), the distance constraint equa-\ntions l2i = || \u2212\u2192 AiBiO|| 2 , i = 1 . . . 6 can be expressed Thus, for"
        ]
    },
    {
        "image_filename": "designv11_12_0003211_amr.505.154-Figure3-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003211_amr.505.154-Figure3-1.png",
        "caption": "Fig. 3 Finite element models of friction clutch with load types (a) A type & (b) B type.",
        "texts": [],
        "surrounding_texts": [
            "In this paper, an educational software called Heatflux for computing the amount of heat generated on the flywheel, clutch (flywheel side and pressure plate side) and pressure plate is used. In this software, any function of rotational sliding speed and torque of friction clutch can be inserted to obtain a function for heat generated with radius and time. In the second part of this work the finite element method has been applied to study the influence of non-dimensional radius (R) on the maximum temperature, average temperature and temperature distribution of a friction material. The conclusions obtain from the present analysis can be summarized as follows: 1. The ratio of inner to outer radius of friction surface (R), which is considered the single most important factor affecting the design parameters and thermal behaviour of friction clutch. 2. In this work the analysis was based on the uniform pressure and uniform wear theories. To obtain results with high accuracy, one must know the proper functions of pressure with radius and rotational sliding speed with time. 3. The friction material of clutch should have perfect thermal properties and higher wear resistance for thermal stabilities. 4. The amount of heat generated on the friction clutch side (flywheel side or pressure plate side) is less than 5% from the total heat generated between surfaces in one side. 5. The maximum effect of thermal load (the highest temperatures) at the friction interface for all cases occur approximately at half sliding time (0.2 sec), and the highest average temperatures occur in 0.3 sec (0.75% from the slipping time)."
        ]
    },
    {
        "image_filename": "designv11_12_0002129_s0263574709005426-Figure10-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002129_s0263574709005426-Figure10-1.png",
        "caption": "Fig. 10. Closure error represented by a virtual chain in a four bar mechanism.",
        "texts": [
            " In the example, the twist R$j , associated with the joint j, will have a positive sign in the circuit a equation and a negative sign in the circuit b equation. An integration algorithm is necessary to integrate the differential kinematics equation to obtain the joint positions.19 The algorithm proposed in this paper has two steps. The first step is to introduce a virtual chain to represent the closure error resulting from the integration error as shown in Fig. 2. For the same example of a four bar planar mechanism, the resulting closed chain is shown in Fig. 10. The constraint equation of this closed-loop chain results in Npq\u0307p + Nsq\u0307s + Neq\u0307e = 0, (18) where Np and Ns are the primary and secondary network matrices obtained by integration, q\u0307p and q\u0307s are the primary and secondary magnitude vectors, respectively, Ne is the error network matrix and q\u0307e is the error magnitude vector. The second step is to replace Eq. (3) by q\u0307s = \u2212N\u22121 s Npq\u0307p + N\u22121 s NeKeqe, (19) where the gain matrix Ke is chosen to be positive definite and qe is the position error vector"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001247_0954406jmes321-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001247_0954406jmes321-Figure1-1.png",
        "caption": "Fig. 1 Transverse tooth profiles of an involute cylindrical shaper",
        "texts": [
            " As a result, this method, for which computer-aided design software for tooth surface geometry has been developed, is suitable for face gears made by die, including forging or metal mould injection processes. In addition, as the numerical examples illustrate, it is highly flexible. The mathematical model for an involute helical gear can also be applied to simulation of the tooth surface of the shaping cutter used for standard face gears. Assuming that the tooth profiles in the transverse plane of the involute helical gear are represented in the coordinate system Ss(xs, ys, zs), as shown in Fig. 1, the variables \u03b8ks (k = \u03b7, \u03b3 ) are the tooth profile parameters. The constant \u03b8os determines half the space width on the base cylinder, and the transverse tooth profile of a standard involute helical gear is represented by equations (1) to (3) [10] \u03b8os = \u03c0 2Ns \u2212 inv\u03b1s inv\u03b1s = tan \u03b1s \u2212 \u03b1s (1) \u03b1s = tan\u22121 ( tan \u03b1n cos \u03b2 ) (2) where \u03b2 is the helical angle, and parameters \u03b1s and \u03b1n are the transverse pressure angle andthe normal pressure angle, respectively. xs = \u00b1rbs[sin(\u03b8ks + \u03b8os) \u2212 \u03b8ks cos(\u03b8ks + \u03b8os)] ys = \u2212rbs[cos(\u03b8ks + \u03b8os) + \u03b8ks sin(\u03b8ks + \u03b8os)] zs = 0 (3) The upper and lower signs in equation (3) correspond to the right- and left-side profiles, respectively, while rbs is the radius of the base circle"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000501_s00227-007-0891-x-Figure3-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000501_s00227-007-0891-x-Figure3-1.png",
        "caption": "Fig. 3 Orthogonal right-hand coordinate systems: fixed in space O1XYZ and fixed in the body Oxyz. The origins of the moving coordinate systems coincide with the center of mass of the larva-particle. The unit vectors of the fixed in the space coordinate system are denoted as jX, jY and jZ, respectively. The rotation of the fixed in the body coordinate system Oxyz as a whole is described by three Euler\u2019s angles (/, w, h), which can be chosen in different ways. We chose here a coordinate system with the line of nodes defined as the line of intersection of the planes x0Oy0 and yOz. The angle of yaw / represents the rotation around Oz0 axis, the rotation about the line of nodes is described by the angle pitch w, and the angle of roll h corresponds to the rotation around the Ox axis. The relation between the coordinate systems Ox0y0z0 and Oxyz is given by",
        "texts": [
            " In the course of their motion larvae can change their shape and volume. In our investigation, however, we do not take all these factors into account and consider only the first approximation of the form of a lava: a rigid smooth sphere. Coordinate systems Three right-hand orthogonal coordinate systems are used here: fixed in space O1XYZ, fixed in the body of a larvaparticle Oxyz, and an auxiliary coordinate system Ox0y0z0, which moves with the larva-particle while its axes remain parallel to the axes of the fixed in the space coordinate system (Fig. 3). The orientation of the axes of the attached to the body coordinate system is determined by three Euler angles and the corresponding matrix of the cosines of directions (Fig. 3). The motion of a larva-particle can be represented as a translation described in the fixed in space coordinate system O1XYZ by the vector of linear velocity V(VX, VY, VZ) of the origin O and the rotation with the angular velocity x\u00f0xx;xy;xz\u00de: The origin of the moving coordinate systems in the inertial coordinate system is determined by its radius vector rO (XO, YO, ZO) and its velocity can be calculated as the derivative of the radius vector with respect to time t drO dt \u00bc V: \u00f06\u00de The relations between the Euler angles and the corresponding components of the angular velocity x\u00f0xx;xy;xz\u00de on the axes of the fixed in the body coordinate system are given as follows (Goldstein et al"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001637_s11044-010-9215-x-Figure4-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001637_s11044-010-9215-x-Figure4-1.png",
        "caption": "Fig. 4 Rail and wheel profiles",
        "texts": [
            "1 Test track and truck specifications In order to validate the location of contact points predicted by the numerical procedure developed in this investigation, contact points on tight radius curved track are compared with those obtained using experiments. A test track of the University of Tokyo, Institute of Industrial Science, is used for this purpose. The curve radius is 48.3 m; track gauge is 1435 mm; 20 mm slack is given on the curved section; and the cant is 0 mm. The rail profile is shown in Fig. 4(a), which is a Japanese industrial standards (JIS) 50 kg N rail used in commuter trains in Japan. The truck used in this experiment is a bolster truck. The wheelbase, wheel diameter and weight of the truck are, respectively, 2100 mm, 762 mm and 4965 kg. The wheel profile is as shown in Fig. 4(b), which is a 1:20 conical profile. The difference in rolling radii and contact angles obtained using these profiles are shown in Fig. 5 when the angle of attack is assumed to be zero. It can be observed from this figure that the flange contact occurs for the lateral displacement of approximately 19 mm when the angle of attack is equal to zero. Since 20 mm slack (gauge widening) is given due to very small radius of curve, large lateral displacement is required to have flange contact. The dynamic curving simulation of the two-axle truck used in the experiment is performed to estimate the angle of attack during the curved negotiation on the test truck",
            " This result is consistent with Hertz\u2019s contact theory, in which a ratio of the longitudinal and lateral semi-axes of the contact ellipse (a/b ratio) varies in the same way as a ratio of the following longitudinal and lateral curvatures (B/A) of surfaces in contact: A = 1 2 ( 1 Rw 1 + 1 Rr 1 ) , B = 1 2 ( 1 Rw 2 + 1 Rr 2 ) (17) where R1 indicates the radius of longitudinal curvature at contact point, while R2 indicates the radius of lateral curvature. In the case of flange contact, the lateral curvature B becomes large since the wheel flange comes into contact with rail at the gauge corner where the radius of curvature is very small (Rr 2 = 0.013 m as shown in Fig. 4(a)), while the longitudinal curvature A becomes small since the rolling radius of wheel become larger at the flange point. As a result, the longitudinally long contact ellipse is obtained as in the measurement result shown in Fig. 10. On the other hand, in the case of tread contact at the neutral position, since the lateral curvature on the top of the rail is Rr 2 = 0.3 m as shown in Fig. 4(a) and the nominal rolling radius is Rw 1 = 0.381 m, the contact ellipse becomes a circular shape with a little large longitudinal semi-axis. In order to compare the location of contact point, the lateral and longitudinal positions of the contact ellipse are, respectively, shown in Figs. 11 and 12. In these figures, the longitudinal and lateral widths of the contact ellipse are shown by the bar. That is, the contact point predicted by solving either (13) or (15) has to be inside this range. Note also that tread and flange two contacts are given in the outer rail of Experiment #3"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003777_20110828-6-it-1002.01580-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003777_20110828-6-it-1002.01580-Figure1-1.png",
        "caption": "Fig. 1. Frames and actuators",
        "texts": [
            " The next section describes the AUV non-linear model and its linearization for controller synthesis. In section 3 the design of a pitch angle controller using H\u221e framework 978-3-902661-93-7/11/$20.00 \u00a9 2011 IFAC 14729 10.3182/20110828-6-IT-1002.01580 is presented. Then an altitude controller scheduling w.r.t the sampling interval is built using the LPV/H\u221e control design. The last section presents some simulation results and some future perspectives are drawn. The vehicle considered here is the Asterx AUV designed and operated by Ifremer (Figure 1). The model is adapted from Fossen (1994), and described in more details in Roche et al. (2010). For the description of the vehicle behavior, we consider a 12 dimensional state vector : X = [\u03b7(6) \u03bd(6)] T . \u03b7(6) is the position, in the inertial referential R0, describing the linear position \u03b71 and the angular position \u03b72: \u03b7 = [\u03b71 \u03b72] T with \u03b71 = [x y z] T and \u03b7(2) = [\u03c6 \u03b8 \u03c8] T where x, y and z are the positions of the vehicle , and \u03c6, \u03b8 and \u03c8 are respectively the roll, pitch and yaw angles. \u03bd(6) represents the velocity vector, in the local referential R (linked to the vehicle) describing the linear and angular velocities (first derivative of the position, considering the referential transform) : \u03bd = [\u03bd1 \u03bd2] T with \u03bd1 = [u v w] T and \u03bd2 = [p q r] T The AUV is actuated using 6 inputs: \u2022 a forward force Qc for the axial propeller \u2022 2 horizontal fins in the front part of the vehicle (controller with angle \u03b21 and \u03b2\u2032 1) \u2022 3 fins at the tail , one vertical (angle \u03b4) and 2 tilted with angle \u00b1\u03c0/3 (controlled by angle \u03b22 and \u03b2\u2032 2) The nonlinear model includes 12 state variables and 6 control inputs"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003950_rnc.3288-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003950_rnc.3288-Figure2-1.png",
        "caption": "Figure 2. Inverted pendulum system.",
        "texts": [
            " Remark 4 From (20), it is obvious that with the norm bound of the estimation error of the initial condition, only the estimated state at the current updating instant is required to determine the next updating Copyright \u00a9 2014 John Wiley & Sons, Ltd. Int. J. Robust Nonlinear Control (2014) DOI: 10.1002/rnc instant. That is, at the current updating instant, which means at each updating instant, it is already known when the next event is going to happen. Consider the cart with an inverted pendulum system as shown in Figure 2, where M is the mass of the cart, m is the mass of the pendulum, l is the length to the pendulum center of the mass, x is the cart position coordinate, is the pendulum angle from the vertical, and F is the input force. The linearized system model can be represented in state-space form as 2 664 Px Rx P R 3 775 D 2 6664 0 1 0 0 0 .ICml2/b I.MCm/CMml2 m2gl2 I.MCm/CMml2 0 0 0 0 1 0 mlb I.MCm/CMml2 mgl.MCm/ I.MCm/CMml2 0 3 7775 2 664 x Px P 3 775C 2 6664 0 ICml2 I.MCm/CMml2 0 ml I.MCm/CMml2 3 7775u; y D 1 0 0 0 0 0 1 0 2664 x Px P 3 775 ; where b is the friction of the cart and I is the inertia of the pendulum"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001936_iembs.2010.5627660-Figure6-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001936_iembs.2010.5627660-Figure6-1.png",
        "caption": "Fig. 6. The scanned calibration phantom. (a) The reference image of the calibration phantom. (b) The phantom at one of the robot positions used for calibration.",
        "texts": [
            " When the head is tracked via the MRI reference image with the laserscanner, the robot can move the coil to the desired target position at the real head. Therefore, we use the transform from coil to endeffector that has been calculated in the coil registration step. The needed transforms for the head navigation with the laserscanner are illustrated in Figure 5. Our first experiments have shown, that a calibration with a mean translational error < 2 mm can be achieved with the presented setup and methods. The rotational error was < 0.75\u25e6 with a scaling error < 0.001. We used n = 10 different robot positions for performing the calibration. Figure 6 presents the reference image of the calibration phantom and a laserscan of the phantom at one of the robot positions that are used for calibration. For head tracking, the first experiments have shown that the mean error using ICP for the calculation of reference image to scanned image was < 0.2 mm. Figure 7 shows the head phantom with the reference image generated from an MRI-scan and the image obtained by the laserscanner. The computation time for ICP using a low resolution laser scan image and a high resolution MRI reference image was in the range of 30 s"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002949_etep.1642-Figure11-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002949_etep.1642-Figure11-1.png",
        "caption": "Figure 11. The serial number of the stator tooth and rotor tooth.",
        "texts": [
            "1002/etep the stator winding would in effect create a rotating constant flux that sweeps the rotor bars to generate the rotor e.m.f.. As a result, the e.m.f. and the induced current are identical in all the rotor bars around the rotor periphery. At standstill, however, the rotor bars are fixed in position with respect to the stator and are thus subjected to local variations in the stator m.m.f. [35]. The current amplitude of the healthy rotor varies with position around the rotor periphery and is not equal at standstill, which is different from the rated load. Figure 11 indicates the broken bar positions within the rotor geometry that is used throughout the paper, and Figure 12 shows the rotor current distributions in the three cases. It is easily understood that the currents would increase in the neighboring bars next to the broken ones. When bar 3 is broken, the currents of bars 2 and 4 are seen to rise, and in the case (b), the amplitude of the current in bar 4 presents a staggering increase. Similarly, bars 2 and 5 increase dramatically when bars 3 and 4 are broken in the case (c)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002241_978-1-4471-4141-9_70-Figure70.2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002241_978-1-4471-4141-9_70-Figure70.2-1.png",
        "caption": "Fig. 70.2 The Bricard mechanism that can be deployed onto \u2018\u2018Y\u2019\u2019 shape profile. a Deployed configuration, b folded configuration",
        "texts": [
            "1, for the general line-symmetric case, the geometric constraints are given as X1 \u00bc X2; X3 \u00bc X4; X5 \u00bc X6 a12 \u00bc a45; a23 \u00bc a56; a34 \u00bc a61 R1 \u00bc R4; R2 \u00bc R5; R3 \u00bc R6 8 >< >: \u00f070:1\u00de where Xi is the common perpendicular of zi and zi\u00fe1; aij is the rotation angle of axes zi and zi\u00fe1 around Xi; Ri is the distance of link i 1 and link i along zi, refers as the offset of joint i: Ref. [6] has presented the possibility of using a trihedral case Bricard mechanism to design deployable mechanism that can be deployed onto \u2018Y\u2019 shape profile. As shown in Fig. 70.2, it can be folded onto a bundle compact form as shown in Fig. 70.2b. For this mechanism, one can see that the link pairs a; f\u00f0 \u00de; b; c\u00f0 \u00de; d; e\u00f0 \u00de of the mechanism can be designed with arbitrary length without changing the mobility of the mechanism, in the deployed configuration, three angles \\AOB; \\BOC and \\COA are identical. The GLSBL can also be designed to be deployable onto such \u2018\u2018Y\u2019\u2019 shape deployed profile by using the joint axis position determination method as presented in [4]. Suppose that ABC is the required triangular deployed profile, for this case, however, the three angles \\AOB; \\BOC and \\COA are no longer identical because the D\u2013H lengths of the six links are not identical, so that the mechanism can be designed with non-equilateral triangular profiles, but the physical link of the six links for GLSBL can be designed with identical lengths"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000578_009-Figure9-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000578_009-Figure9-1.png",
        "caption": "Figure 9. TA inside the confinement vessel.",
        "texts": [
            " The small viewing port is prepared as an alternative for viewing and lighting. The TA has a higher Li leak risk than other parts since the bayonet TA employs a mechanically replaceable back plate and a dissimilar weld joint as a connection between the F82H TA and the SS 304 pipe of the main loop at the upstream of the inlet nozzle. As a countermeasure for Li leak from the TA, the TA is held in an air-tight steel container called a confinement vessel to provide a two-fold boundary structure. The structure of the confinement vessel is illustrated in figure 9. The vessel, whose diameter is approximately 2.4 m and whose height is approximately 5.3 m, is connected to an Ar gas system of the loop and is equipped with an air blower to replace the inner atmosphere with Ar gas or air. With this equipment, this confinement vessel is filled with Ar gas at a slightly positive pressure during Li loop operation to prevent Li fire in case of accidental Li leak. The confinement vessel has a measurement hatch which is connected to the large viewing port in the target chamber"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001015_j.talanta.2007.08.036-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001015_j.talanta.2007.08.036-Figure2-1.png",
        "caption": "Fig. 2. Scheme of the tyrosinase biosensor used: (a) external structure of the electrode; (b) internal solution (phosphate buffer 0.067 mol l\u22121 at pH 6.6 and KCl 0.1 mol l\u22121); (c) reference inner electrode: Ag/AgCl (anode); (d) Pt electrode (cathode); (e) Teflon cap; (f) gas-permeable membrane; (g) Teflon O-ring; (h) d (",
        "texts": [
            " The ecorded signal was proportional to the variation of the parial vapor pressure of oxygen dissolved in solution, which was onsumed in the course of the enzymatic reaction (1): henol + O2 tyrosinase\u2212\u2192 o-quinone + H2O (1) or this purpose, a tyrosinase biosensor made of polytetrafluroethylene (PTFE) capable of measuring the decrease in L. Campanella et al. / Talanta d n ( e r ( t t h i s C t d T w s a w F 3 w u e w b b s o t t m c i W s i m a a r o p w t a C w r r t a 3 t p i t d ialysis membrane; (i) tyrosinase immobilized in Kappa-Carrageenan gel; and l) glass insulator. issolved oxygen concentration due to the oxidation of a pheolic compound to a o-quinone operating in n-hexane was used this kind of biosensor is usually denoted as an organic phase nzyme electrode, OPEE) (Fig. 2). The variation of the signal ecorded is proportional to the species reduced at the cathode O2), the amount of which depends on the concentration of he oxidized phenols in the sample. The decision to carry out he determination of the polyphenols in an organic solvent (nexane) was based on the very high solubility of the oil matrix n this organic solvent and to its very low solubility in aqueous olution [8,15]. The tyrosinase enzyme was immobilized in a Kappaarrageenan gel [16,17] prepared as follows: 3"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002759_apec.2010.5433395-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002759_apec.2010.5433395-Figure1-1.png",
        "caption": "Fig. 1 Voltage source inverter fed DTP-PMSM",
        "texts": [
            " The controlled power can be divided among more inverter legs to reduce the single static switches current stress without the need for parallel techniques or multilevel converter. The multi-phase AC motor drives also show several advantages over the conventional three-phase ones such as reducing the amplitude of torque pulsation, lowering the dc link current harmonics and higher reliability [1-3]. The DTP-PMSM has two three-phase windings spatially shifted by 30 electrical degrees with isolated neutrals. The stator windings are fed by a current controlled PWM six-phase voltage source inverter. A DTP-PMSM and the inverter arrangement are illustrated in Fig.1. The one three-phase system is composed of stator windings A, C and E. Another one is B, D and F. The two sets of windings have spatially shifted by 30 electrical degrees with isolated neutrals. During the last years, many authors have presented several SVPWM techniques [4-7]. The aim of the space vector PWM techniques is to reduce the stator current harmonics which have been observed in the machine stator current spectrum. The main goal of this paper is to provide a comprehensive comparison between the performances of these modulation techniques based on several criteria: current harmonic minimization, utilization of DC bus and implementation complexity with low cost fixed-point DSP platforms"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001196_0094-114x(72)90004-3-Figure6-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001196_0094-114x(72)90004-3-Figure6-1.png",
        "caption": "Figure 6. The p lane RP*R*P* mechan i sm.",
        "texts": [
            " , as ~ 1sin = a , COS al sin as tsln= a , cos a, J sin a4f K t ] as ~ t s in = a , COS a J \" (61) In addition, joints 2 and 4 of the chain have pass ive rotat ions in one phase (construction) of the mechan ism if equat ion (38) is satisfied. N o t e that any set of twist angles which satisfy equat ion (38) also satisfy equat ion (43). Thus any linkage of which the paramete rs satisfy equat ion (38) and one of the preceeding seventeen sets of conditions for s3 to be constant is an RCP-*C in one phase and an R P *\u2022 *P * in its o ther phase. A linkage with four ninety-degree twist angles is an R P * R * P * mechanism in both phases. As shown in Fig. 6, this is a planar mechan i sm in that its relative mot ions occur in planes which are perpendicular to the two revolute axes. 8. Linkages of the PCCC Chain To include linkages which only have cyl indrical and prismatic joints, the PCCC chain is investigated. Due to the similarit ies between the equations describing this chain and the RCCC chain, the same five mechanism groups exist. I t can also be said that any linkage arising from this chain which has a revolute jo int has already been considered in the study of the RCCC chain",
            " Thus the RCCR linkage can be constructed as an RCRC linkage without changing anything else in addition to kinematic inversion where the entire set of joint and requirement subscripts must be indexed. To show the general form of physical realizations of these linkages, Figs. 9-12 show models of linkages of the first four mechanism groups. The fifth group is omitted since it only contains the plane four-bar linkage. In addition to the plane four-revolute mechanism, there are three more plane four-link mechanisms in the table. The RP*R*P* mechanism produced by equation (45) is a plane mechanism as previously noted and shown in Fig. 6. The RP*P*R* mechanism of group 2 is also a plane mechanism with two revolute joints and two prisms or sliders. Figure 13 illustrates this mechanism. And finally, the RRR*P* mechanism of group 4 is the familiar crank-slider linkage. It has been brought to the attention of the author that a similar but not identical linkage summary was simultaneously obtained[10]. Whereas the present study is an extension of the German literature[6-8] the similar study is an extension of the Russian literature[5, 9, 14]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001306_icelmach.2008.4800129-Figure4-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001306_icelmach.2008.4800129-Figure4-1.png",
        "caption": "Fig. 4. Magnetic flux density and flux lines within synchronous machine as determined by the FEM analysis.",
        "texts": [
            " Further operations are then applied to determine the magnetic and electric fields within the model. To determine the eddy-current losses in the rotor and stator sleeves of the machine, a 2D transient model was created. The stator and rotor sleeves were defined as coils within the model and were short circuited. This technique neglects end effects of the sleeve and will result in a higher than expected value for the power loss [8]. An image of the magnetic flux density within the machine is shown in Fig 4. The resulting eddy-current losses in the rotor sleeve were Pr = 24.8 kW, and in the stator sleeve were Ps = 384 kW. CONCLUSIONS As expected, the losses obtained from the FEM calculations are higher than those obtained from analytical calculations (384 kW versus 316 kW for the stator sleeve and 24.8 versus 11.02 kW (2D field) and 7.8 kW (simple analytical equation) for the rotor sleeve. This is, amongst others, due to taking into account the actual distribution of the magnetic field in the FEM instead of only certain harmonics of the air gap magnetic flux density in analytical calculations"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001866_bf00251592-Figure9-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001866_bf00251592-Figure9-1.png",
        "caption": "Fig. 9. Cusp locations in Stvphcnson-I symmetrical six-bar curve",
        "texts": [
            " As in the Watt-1 mechanism, therefore, we confine our analysis Six-Bar Motion. II 47 to the symmetrical six-bar curve for which the joints of the triangular links are collinear. To any cusp position, then, there also corresponds a mirror-image cusp position (reflection about A E). The special case of a cusp position for a symmetrical Stephenson-1 curve is shown in Fig.8. The cusp configurations may occur with CG and GF parallel or antiparallel to A C and FE, respectively. Any cusp, therefore, lies at one of the intersections of the following four circles (see Fig. 9): Circle C1 : Center A, radius rl = s + / ; Circle C2 : Center A, radius I s - I I ; Circle C3 : Center E, radius s' + l'; Circle C4: Center E, radius Is'-l ' l . Since there are at most eight real intersections of these four circles, the maximum number of real cusps for the symmetrical six-bar curve cannot exceed eight. To find these, we apply the cosine law to triangles GBD and GAE of Fig.8: cos2= (s+l)2+(s'+l')Z-m2 (s+l-r)2+(s '+l ' -r ' )Z-n2 - ( 6 ) 2(s+t)(s'+r) 2(s + l-r)(s ' + l ' -r ' ) Equation (6) is the condition for the existence of a cusp in the position shown. This equation can also be written with 1 and l' replaced by ( - / ) and/or ( - l ' ) , resulting in a total of four equations for the four cusp positions 1, 2, 3, 4 of Fig. 9 (and their mirror images). The equations are necessary, but not sufficient to insure the existence of eight cusps. For each equation must not only be satisfied, but the value of cos 2 determined from the equation must lie within ( - 1, 1). Following the procedure used in the Watt-1 curve and earlier [5], we put equation (6) and those derived from it into the form ql I l' +q2 l\" +q3 l + q 4 = 0 where ql, q2, q3, qr are invariant with respect to changes in sign of 1 and/or l '; that is, we have four equations of this type",
            " Letting r, = s + l , rE=s-I or l - s , ra=s'+l', r4=s ' - l ' or l ' - s ' , where all r i are non-negative, we are led by this condition, after some algebra, to the following cubic equation in m2: m 6 _ ( r 2 + 2 2 2 r 2 + r 3 + r 4 + r l r2 + r 3 r4) m 4 + + [ ( r x r2 -- r3 r4) 2 + r l r2 (r 2 + r 2) + r3 r4( r 2 + r 2) + (r 2 + r 2 ) ( r 2 + r2 ) ] m 2 - - - ( r , - - r,,) [ r , + r# + r , ' 9 - - r , , ( q + + r, ,)] = O. (11) Equation (11) determines a value of m as a function of s, l, s', I'. If these four values are given and m is determined from equation (11), the remaining proportions, obtained by rearrangement of equations (7)-(10), will be r=(s' 2 + 3 s2--1 ' 2+ lZ m2)/2S r'=(s2 + 3s' 2 +l' 2--12--m2)/2s ' nE=r2-Fr' 2 - - 4 s 2 - - 4 s ' 2 + 3 m E . (12) (13) (14) For real cusps, it is necessary that n 2 and m 2, obtained as above, be positive and that m lies within certain limits determined from Fig. 9, so that the circles C1,..., C4 have eight real and distinct intersections. Curves S 1-101, S 1-114, S 1-201, obtained in this fashion, show symmetrical Stephenson-1 six-bar curves with eight cusps. For fewer cusps it is possible to devise simple geometrical constructions to determine the mechanism proportions. 8. Circuits Following the procedure used for the Watt-1 curve, we shall find the equation between w=(CF) 2 and tp = L EAB, in the absence of the floating links CG and GF. This equation, as before, is called the range equation, or the (q~, w)-curve, since its roots when w = (l_+/,)2 determine the number of circuits of the six-bar curve, as well as the extreme positions of link A B C in each of these circuits"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000104_j.triboint.2006.10.004-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000104_j.triboint.2006.10.004-Figure1-1.png",
        "caption": "Fig. 1. Tribopair with liquid\u2013solid lubrication.",
        "texts": [
            " In this paper, a mixed liquid\u2013solid lubrication model is presented which combined Reynolds equation, the load carrying equations of spherical particle, and the contacting equation of asperities. Then, utilizing mathematical simulations, this paper compared the results, which obtained under different conditions. Furthermore, the influence of the particles\u2019 and tribopairs\u2019 properties on the loadcarrying capacity and the temperature is discussed both qualitatively and quantitatively. The physical model of the research shown in Fig. 1 illustrates the configuration of a mixed liquid\u2013solid contact between two surfaces. As is shown in Fig. 1, the particle size is of the same order of magnitude with that of the clearance and surface asperities. In the contact zone, both particles\u2019 and asperities\u2019 deformation are significant. To analyze the performance of the mixed liquid\u2013solid lubrication, all of the liquid, the spherical solid particles, and the tribopairs\u2019 surface asperities must be taken into account. The profile of upper surface, which is chosen according to traditional piston ring profile design, is (3 10 6\u20131.33x2) where the x is between 0 and B"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002569_1350650112451218-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002569_1350650112451218-Figure2-1.png",
        "caption": "Figure 2. Representation of the studied geometry.",
        "texts": [
            " Friction coefficient The friction torque is obtained by integration of shear stresses at the surface of the shaft or the bushing9 C \u00bc R2L Z2 0 XY\u00f0 y \u00bc h\u00ded \u00f06\u00de with XY \u00bc R! h \u00fe 1 2 dp dx \u00f02y h\u00de \u00f07\u00de Then, the friction coefficients ( f ) are calculated with the following formula f \u00bc Torque RB:W \u00f08\u00de The term RB is the radius of the plain bearing. at NANYANG TECH UNIV LIBRARY on June 4, 2015pij.sagepub.comDownloaded from Numerical model is carried out by the resolution of the deformation relationship using finite element method, with ANSYS CFX code. Pressure field in oil film is given by the resolution of Reynolds equation with the finite difference method. Figure 2 shows the model used in the analysis. The mesh of bearing is composed by tetrahedral and hexahedral elements, of which the number of node is 74163 and 30433 elements. The numerical resolution strategy of an ANSYS CFX code is the following: . modeling of the structure, meshing; . modeling of the load; . solving Reynolds equation by the finite difference method; . calculation of the pressure field; . solving the interpolation equations and calculation of overall structure stiffness matrix; . calculation of displacement, strain; and "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001637_s11044-010-9215-x-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001637_s11044-010-9215-x-Figure1-1.png",
        "caption": "Fig. 1 Parameterization of wheelset",
        "texts": [
            " 2, the wheel and rail coordinate systems are introduced and the formulation of wheel and rail contact is briefly summarized. The governing equations to be solved for one and two-point contact scenarios are presented in Sect. 3. In Sect. 4, specifications of test track and two-axle truck are explained and measurement systems developed for this investigation are presented. Numerical results and comparison with the experimental results are presented in Sect. 5. Summary and conclusions drawn from this study are given in Sect. 6. 2.1 Parameterization of wheelset As shown in Fig. 1, the global position vector of an arbitrary point of contact on wheelset w can be defined as [9] rwk = Rw + Awu\u0304wk k = 1,2,3 (1) where superscript k denotes the contact number (k = 1 for the right wheel tread, k = 2 for the left wheel tread, and k = 3 for the wheel flange); the position vector Rw = [Rw X Rw Y Rw Z ]T defines the global position of the origin of the wheelset coordinate system; Aw is the orientation matrix defined using the three Euler angles \u03b8w = [\u03c8w \u03c6w \u03b8w]T (yaw angle \u03c8w , roll angle \u03c6w and pitch angle \u03b8w); and u\u0304wk is the local position vector that defines the location of contact point k with respect to the wheelset body coordinate system. The profile coordinate system XwkY wkZwk (k = 1,2) is defined in order to parameterize the wheel geometry as shown in Fig. 1. Using the wheel profile coordinate system, the lateral and circumferential surface parameters given by swk 1 and swk 2 are introduced as shown in Fig. 1. With these two parameters, the rolling radius of each wheel is given by gk(swk 1 , swk 2 ). If the rolling radius can be assumed to be constant in the circumferential direction of wheel, one can have gk = gk(swk 1 ). Accordingly, the locations of arbitrary points of contact on the right and left wheels can be defined with respect to the body coordinate system as follows: u\u0304w1 = \u23a1 \u23a3 g1(sw1 1 ) sin sw1 2 \u2212Lw1 + sw1 1 \u2212g1(sw1 1 ) cos sw1 2 \u23a4 \u23a6 , u\u0304w2 = \u23a1 \u23a3 g2(sw2 1 ) sin sw2 2 Lw2 + sw2 1 \u2212g2(sw2 1 ) cos sw2 2 \u23a4 \u23a6 (2) where Lwk is the distance between the origins of the body and profile coordinate systems along the Y w axis as shown in Fig. 1. Using the expressions given by (2), the two tangents at contact point k along the lateral (swk 1 ) and circumferential (swk 2 ) directions as well as the normal can be given as t\u0304wk 1 = \u2202u\u0304wk \u2202swk 1 , t\u0304wk 2 = \u2202u\u0304wk \u2202swk 2 , n\u0304wk = t\u0304wk 1 \u00d7 t\u0304wk 2 (3) The preceding equations are written with respect to the global coordinate systems as twk 1 = Aw t\u0304wk 1 , twk 2 = Aw t\u0304wk 2 , nwk = twk 1 \u00d7 twk 2 (4) 2.2 Parameterization of curved track In existing contact geometry analysis procedures, a tangent track model is used and contact points obtained using this assumption is used for curved negotiation as well [6]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000094_jp073212s-Figure6-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000094_jp073212s-Figure6-1.png",
        "caption": "Figure 6. UV-vis spectra of Hb in pH 7.4 PBS (a) and on CaP (b) and CaP-PDDA (c) modified ITO glass and hemin in pH 9.0 PBS (d).",
        "texts": [
            " Here we immobilized Hb on the CaPPDDA-modified GCE which was prepared by electrodeposit in the electrolyte-containing 1% PDDA for 10 min and was used as a model to search the new applicable approach for the fabrication of biosensors. UV-vis Absorption Spectroscopic Characterization. The variation of the conformational structures of proteins can be demonstrated from the changes of the shape and position of the Soret absorption bands since it is sensitive to the variation of the microenvironment around the heme site.49 Figure 6 shows the spectra of Hb in PBS solution (curve a), assembled on a CaP film (curve b) and a CaP-PDDA film (curve c) on the surface of ITO glass. When Hb was in pH 7.4 PBS alone, the Soret band appeared at 408 nm, while it shifted to 411 nm both in CaP and CaP-PDDA films. The slight shift may be due to the interaction between the films and protein. For further comparison, we tested the UV-vis spectra of hemin, a free heme group, the Soret band of which was observed at 388 nm in solution (curve d)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000501_s00227-007-0891-x-Figure21-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000501_s00227-007-0891-x-Figure21-1.png",
        "caption": "Fig. 21 Approximation of a boundary layer velocity profile by a linear velocity profile. Circles denote larvae in a linear shear flow and in a boundary layer of a body of general geometry. Since both flows are vortical, the kinematic behavior of larvae in a boundary layer and in a linear shear flow are expected to be similar",
        "texts": [],
        "surrounding_texts": [
            "We made an attempt to explain larval contact in laminar flows as a result of the combined action of the flow vorticity and of the self-propulsion of a larva. We studied larval contact in three types of laminar flows: in a two- dimensional linear shear flow, in the two-dimensional Poiseuille channel, and in the three-dimensional Poiseuille tube. According to the presented here mathematical model the probability of contact is a decaying function of the ratio of the characteristic flow velocity and swimming velocity of larva. The simulated values of the contact probability for a tube flow are in satisfactory qualitative agreement with experimental measurements of the attachment probability. Most experimental observations confirm the decay of attachment rate as the flow velocity increases. The common explanation of this experimental fact is well known. It is believed that shear stress on the body of a larva after it made contact is strongly associated with the wall shear stress. Because in a tube flow the wall shear stress is proportional to the corresponding wall rate of shear, Crisp (1955) suggested that the probability of larval settlement is related to the rate of shear. However, in tube flows the rate of shear, the shear stress and the vorticity are proportional to the same gradient of the axial velocity calculated in the radial direction. This similarity can easily obscure the actual physical relation between the contact and attachment phenomena. In our interpretation, they are related\u2014but governed by two different hydrodynamic characteristics of the flow\u2014one by the flow vorticity, determining contact, and the other one by the wall shear stress that is responsible for the ultimate attachment. The mathematical model is formulated here for laminar flows. It is likely, however, that it can be applied to certain (not all) natural flows. The corresponding examples are given in Figs. 19\u201321. According to the main assumption of our work, to attach, at least in laminar flows, larvae must use their selfthrust. In such a case a question arises how propagules without motility, e.g., many of almost spherical aglal spores can colonize vertical substrates and, in particular, rocky shores (Denny and Shibata 1989). The most likely explanation is that turbulent mixing will cause larval transport. Our model is not intended to describe such processes. Another interesting question is how larvae attach in still water, a case that can be reproduced only under laboratory conditions. Certain types of larvae even in still water are able to reach a substrate by employing gravitational sinking or self-propulsion. However, larval behavior in still water is drastically different from that in flow (Butman and Grassle 1988; Pasternak et al. 2004) and, obviously, our mathematical model is not intended to describe such process either. Our mathematical model is restricted by larval forms with a small degree of non-sphericity and with hydrodynamically smooth surfaces. Real larvae are non-spherical and non-smooth. As long as we remain within the hydrodynamics of low Reynolds numbers, the slenderness of larvae and small variations of their form can still be included into the mathematical model (Happel and Brenner 1983). However, it is likely that analysis of the dynamic of a slender, buoyant and self-propelled larva of an arbitrary form which moves in a vortical laminar along a curvilinear trajectory can be tackled only by solving numerically the full nonlinear Navier\u2013Stokes equations. The analysis presented here is carried out for laminar flows. For turbulent flow with high turbulent intensity and low mean flow velocity the model proposed here can not be applied directly. However, as it was already discussed above many relevant natural flows, where larvae settle, can be considered as partially laminar or can be characterized by a low level of turbulence. In this respect the basic assumption of our model are valid although obviously restricted. Turbulent wake Turbulent wake Flow Fig. 20 Bodies with a well-defined flow separation area which is indicated by a thin black arrow. The thick white arrow on the left shows the direction of the ambient flow. The wakes of the bodies are turbulent; whereas the boundary layer on their front upstream part may be laminar. Examples. The critical Reynolds number for a smooth sphere when the flow on its downstream rear part becomes turbulent is about 3 9 105, which corresponds to the sphere diameter of an order of 5 m and an ambient fluid velocity of an order of 6 cm/s. For a slender smooth body like an ellipsoid with an aspect ratio lager than six the critical Reynolds number is of order of 106, which corresponds to the length of a body of about 16.5 m and to the fluid velocity of about 6 cm/s. Even on relatively large bodies the flow may be fully laminar or laminar on its upstream front parts. Of course, roughness of the body, non-steadiness of the flow and external turbulence may change these estimates but, nevertheless, the order of magnitude of the above numbers is generally accepted (Schlichting 1979) Z X U 1O Y Boundary layer pD O x z y T F Fig. 19 Larva-particle in the laminar boundary layer of a thin plate. The velocity profile in the boundary layer is a nonlinear function of the coordinate normal to the plate. It can be approximated by a linear dependence, yielding a linear shear layer, or by a parabola relating to the velocity profile in Poiseuille flow. The boundary layer in the vicinity of the leading edge is laminar although at the rest of it the flow may be turbulent. The critical Reynolds number when the laminar flow on a plate turns into a turbulent flow is of an order of 3 9 105 which is considered as a lower bound of the critical Reynolds number. For flows with very small external perturbations the critical Reynolds number can be higher by an order of magnitude. Examples. For the fluid velocity of an order of 3 cm/s the length of the laminar region on a flat plate is of an order of 10 m. The thickness of the boundary layer on such a plate after the first 11 cm from the leading edge is more than 1 cm. This is by two orders of magnitude higher than the size of a larva with a typical length of 100 lm. The laminar boundary layer of a flat plate is a very good approximation of the boundary layer of a slender body In our work, we separated biotic and abiotic factors, although it is known that some larva may respond to light and gravity by correcting their body orientation. In this case the effect of vorticity on the body rotation and its trajectory may be less significant. It must be stressed that we did not solve the problem of contact phenomenon in its entire complicity. We did not intend to do it. We have chosen only one important aspect of the problem and studied it by using a simplified hydrodynamic model of the contact phenomenon. More theoretical and experimental works are needed in this direction. In conclusion we wish to mention another important implication of our work, which was suggested by one of the anonymous reviewer of the manuscript. Examples of laminar flows, in which self-propelled organisms are small compared to the geometrical scale of the shear flow, include bacteria and protists in biofilms, and bacteria or motile parasites in internal vessels of other organisms. Acknowledgments The Italian Ministry of Land and Environments (grant 2004/2006) supported this study. The authors are grateful to A. Abelson for constructive discussions and his valuable comments, to L. Shemer for his useful advice on the Monte-Carlo simulations and to N. Paz for her editorial assistance. The manuscript of this work was read by four anonymous reviewers whose comments are greatly appreciated by the authors. Appendix I In this Appendix, we analyze the contact probability for a linear shear flow Eq. 29 bounded by a fixed wall with a coordinate Z = 0. The flow can be unbounded from above, it can represent a flow between two walls, when one of them moves or it can be viewed as linear boundary layer. In all the three cases the characteristic linear scale of the problem is h and the characteristic velocity is Ua. These two parameters allow us to define the non-dimensional time as s = Ua t/h and the non-dimensional coordinates of the larva n = XO/h and f = ZO/h, correspondingly. In nondimensional coordinates the boundaries of the Couette flow are represented by lines f = 0 (lower boundary) and f = 1 (upper boundary). The dimensionless equations of motion Eqs. 30\u201332 can be expressed in a form of a parametric dependence as f \u00bc 2 k C sin2 w 2 ; \u00f0I:1\u00de where C is a constant of integration, which is determined by initial conditions (f0, w0). Expression I.1 represent an infinite number of curves because the number of constants of integration is also infinite. If we plot all such curves f = f(w,C) together (Fig. 22) we obtain a diagram which can be used for analyzing the contact probability. For each particular constant Ci, we have a particular curve fi = f(w,Ci). The correspondence between the constant and the curve (or the curve and the constant) is one-to-one. Each curve of the diagram in Fig. 22 represents a coordinate of a particle f as a function of its angle of pitch, which varies from -p to p. Consider, for instance, a particular value of the constant of integration, say, Ci \\ 1 and the curve fi = f (w,Ci), which pertains to that constant. Assume that at the moment of time t0 a larva has initial coordinates (f0, w0) which correspond to a certain point of the curve. For anther moment of time t1 [ t0 the angle of turn w grows and another pair of coordinates (f1, w1) gives another point of the same curve. If we continue this process, all such points will draw a curve which corresponds to the constant Ci. This process of drawing the curve by a moving point is indicated in Fig. 22 by arrows. The point of intersection of a curve of the diagram with a boundary (boundaries) determines the coordinates of the particle in the moment of contact. For any curve crossing a bounding line any of its points (f0, w0) can be considered as an initial one. Then, a larva-particle starting the motion with this initial condition will eventually reach the boundary (boundaries). The coordinate of the intersection point of the curve f = f (w, C) with the boundaries are f = 0 or f = 1. We are interested neither in these coordinates, nor in the moment of time when contact takes place, but in the fact of crossing of the boundary (boundaries) by curves f = f(w,C), i.e., in the fact of contact. The further analysis is carried out, by using the diagram in Fig. 22 for two separate cases, k [ 2 (Fig. 22a) and k B 2 (Fig. 22b), correspondingly. In Fig. 22 the line f(w, C = 1) (separatrix) divides the domain (0 \\ f\\ 1, -p \\ w \\ p) into two subdomains. All curves which are located below the separatrix correspond to constants C \\ 1. All curves which are located above the separatrix correspond to constants C [ 1. All curves corresponding to C \\ 1 cross the lower boundary. Any point which belong to the area bounded by the curve f(w,C = 1) and line f = 0 crosses the boundary because it belongs to a certain curve f(w,C \\ 1). All such points are depicted as black circles. Points which are depicted as white circles belong to the domain C [ 1; they do not cross the boundary f = 0. If the line f = 1 represents the wall of the Couette channel, the gray points should be counted as those, which contact the upper wall. However, if the line f = 1 represents the boundary of the linear boundary layer the gray points should be counted as those, which do not contact the boundary. For the further calculations it is assumed that larvae are distributed randomly and uniformly within the range (0, 1) and that the initial angles of turn are also random numbers distributed uniformly in the range (-a, a), where a B p. If we take a sufficiently large number of points (an infinite in a limiting case) they will fill corresponding areas. The area, which is filled by particles, crossing the boundary (boundaries) divided by the area filled by all particles can be defined as the probability of contact. It is not difficult to calculate the contact probability for all possible cases, considered above. However, for the sake of brevity we give here formulas for contact probability only for the linear boundary layer. For k [ 2 (Fig. 22a) the probability of contact P can be approximated as the ratio of the area filled by black points and the total area S0 = 2a. Simple calculations of the corresponding areas yield the contact probability in the following form: P \u00bc 1 k 1\u00fe sin a a : \u00f0I:2\u00de For k\\ 2 the calculations of the contact probability are somewhat more complicated because it is necessary to take into account the fact that the line f = 1 crosses the curve f(w,C = 1) at the points b \u00bc 2 arcsin ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 1 k=2 p (Fig. 22b) and the result depends on the sign of the difference a-b. Similarly to the previous case, the contact probability can be calculated as the ratio of two areas, one with black points and the other one including all particles. A closer scrutiny of Fig. 22b and simple calculations of the corresponding areas give the contact probability in the following form: P \u00bc 1 b a; 1 a b\u00fe 1 k \u00f0a b\u00fe sin a sin b\u00de b\\ a: \u00f0I:3\u00de If we assume that all larvae enter the vortical flow with the same zero initial angle of turn (a?0) then formulae I.2 and I.3 can be greatly simplified, yielding Eq. 51. Expressions I.2\u2013I.3 are plotted in Fig. 16. Appendix II In this Appendix, we derive a closed form solution for the probability of contact of larvae moving in a plane Poiseuille channel. It many details the corresponding analysis is similar to that given in Appendix I for a linear shear flow. Therefore, for the sake of brevity, we give here only the principal details of the calculations. Introduce first the non-dimensional variables n = XO/R, f = ZO/R and s = Ua t/R. In non-dimensional coordinates the boundaries of the channel are represented by lines f = \u00b11. It is assumed that the initial coordinates of larvae f0 and the initial angles of turn w0 are represented by random numbers which are distributed uniformly in the rectangle |f0 | \\ 1 and | w0 | \\ a B p. Although the analytic solution of Eqs. 39\u201341 is unknown, the contact probability can still be expressed in a closed form. For the case of a neutrally buoyant larva Eqs. 39\u201341 can be easily reduced to the differential equation of a pendulum d2w ds2 \u00bc 2 sin w k ; \u00f0II:1\u00de which can be transformed into a so-called phase portrait of a pendulum dw ds \u00bc 2 ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi C cos w k r : \u00f0II:2\u00de Here C is a constant of integration, which is determined by the angle of turn and its derivative with respect to nondimensional time (Jordan and Smith 1987). Using Eq. 41 the phase portrait can be transformed into a parametric dependence, which is more convenient for our purposes: f\u00f0w;C\u00de \u00bc ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi C cos w k r ; \u00f0II:3\u00de where the constant of integration is now determined by a pair of numbers (f0, w0). Relation Eq. II.3 is plotted in Fig. 23 and represents closed and open curves with a separatrix f(w,C = 1). Denoting the curve which touches the wall f = \u00b11 at the points w = \u00b1p as fl (w) = \u00b1f (w, Cl), the value of the corresponding constant of integration Cl can be obtained as a solution of an algebraic equation \u00b1 f(\u00b1p,Cl) = \u00b11 yielding Cl = k-1 and fl\u00f0w\u00de \u00bc ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi k 1 cos w k r : \u00f0II:4\u00de Analysis of Fig. 23a shows that for k[ 2 the probability of contact can be calculated as the ratio of the area with black points and the total area filled by all points. Calculating the ratio of the two areas yields P \u00bc 1 1 a Za 0 ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi k 1 cos w k r dw: \u00f0II:5\u00de For k B 2, the limiting curves fl cross the axis f = 0 at angles b = \u00b1 arccos (k - 1) (Fig. 23b). In Fig. 23b, we have three types of points. The white particles belong to closed curves which do not cross the boundary. Once again, the probability of initial contact can be calculated as the ratio of the area with black and gray points and the total area of the corresponding rectangle 2a, which includes all symbols. Analytic calculations of the corresponding areas show that for k B 2 the probability of initial contact is given by P \u00bc 1; b a; 1 1 a Ra b ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi k 1 cos w k q dw; b\\a: 8< : \u00f0II:6\u00de If we assume that all larvae enter the vortical flow with the same zero initial angle of turn (a?0) formulae (II.5) and II.6 become very simple yielding 52. Relations II.5 and II.6 incorporate integrals, which are computed here numerically with a relative error of less than 0.1%. The results of computations for different values of the parameters involved are plotted in Fig. 16. Appendix III The coordinates of N particles randomly and uniformly distributed inside a circle of radius R are calculated as follows (Sobol 1994): XOi \u00bc R ffiffiffiffiffi c1i p cos 2pc2i and ZOi \u00bc R ffiffiffiffiffi c1i p sin 2pc2i \u00f0i \u00bc 1; 2; N\u00de; \u00f0III:1\u00de where c1,2 are distinct random numbers distributed uniformly between 0 and 1. We also assume that the initial direction of the self-propulsion vector is represented by random angles distributed uniformly in the range (-a, a) as r = (2cr-1)a, where r is one of the Euler angles |a| \\ p and cr is a random number distributed uniformly between zero and one. Appendix IV In the right-hand orthogonal coordinate system O1XYZ with vector units of the axes jX, jy and jZ correspondingly, the mathematical operation curl U is given by a determinant curl U jX jY jZ o oX o oY o oZ UX UY UZ \u00bc jX oUZ oY oUY oZ jY oUZ oX oUX oZ \u00fe jZ oUY oX oUX oY : In the case of a two-dimensional motion when the fluid velocity vector is located in the plain O1XZ, the vector of the vorticity has only one component curl2U \u00bc oUX oZ oUZ oX jY ; which is directed perpendicularly to the plane O1XZ. In the case of a linear shear flow when the vector of fluid velocity is parallel to the axis O1X and depends linearly on the coordinate Z the vorticity vector is equal to the constant gradient of the fluid velocity. In this particular case the absolute value of the vorticity vector and of the value of the rate of shear are expressed by the same gradient of the velocity in the direction perpendicular to its vector, i.e., by oUX=oZ:"
        ]
    },
    {
        "image_filename": "designv11_12_0001939_s12283-009-0033-4-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001939_s12283-009-0033-4-Figure1-1.png",
        "caption": "Fig. 1 Representation of instrumented skate blade holder (top), and cross-sectional views of gauge locations (bottom). Gauges AML and PML are oriented vertically along each post, while gauge V is oriented longitudinally along the beam element of the blade holder",
        "texts": [
            "1 Apparatus The instrumentation system consisted of three main components: a hockey skate with strain gauges bonded to the blade holder, a portable data acquisition system, and postprocessing software to convert microstrain signals to force estimates. To determine the vertical and medial\u2013lateral forces exerted on the ice hockey skate (right side only; Nike-Bauer Supreme One95), 0.3175 cm strain gauges (350 X, Vishay, Malvern, PA, USA) were adhered to the skate blade holder at specific locations (Fig. 1). The recorded signals indicated the compressive or tensile deformation of the skate blade holder between the skate boot and skate blade. Temperature compensated half-active Wheatstone bridge circuits were used to convert the resistances provided by the strain gauges into voltage signals, which corresponded to microstrains. One gauge was used to measure the vertical strain (V), and was oriented along the longitudinal axis (transverse plane) of the blade holder\u2019s beam element. The V gauge was referenced to a static gauge, making the V gauge a quarter-active Wheatstone bridge"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003702_cjme.2013.04.801-Figure19-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003702_cjme.2013.04.801-Figure19-1.png",
        "caption": "Fig. 19. 3D model of support phase using crab-type tripod gait to pass over 34.8\u00b0 slope",
        "texts": [
            " and knee joint Based on the results of the analysis in section 3.2 and section 3.3, the pose of the robot is obtained when maximum static torques exist for the hip joint and knee joint. According to the pose of the robot mentioned above and correlative 3D model of the supported phase, a static simulation analysis is performed by the ADAMS software. Meanwhile, the results of a static simulation analysis related to the static torques for the hip joint and knee joint are obtained. The 3D model of the support phase is shown in Fig. 19, and the simulation curves of the static articulated torques are shown in Fig. 20. Fig. 20 shows that the maximum static torque is 992.2 N \u2022 m for the hip joint when the heavy-duty six-legged robot traverses a 34.8\u00b0 slope using the crab-type tripod gait. In Fig. 20, it is also found that the maximum static torque is 471.4 N \u2022 m for the knee joint. In addition, the rearmost leg, which is called leg 2 in Fig. 5, includes the maximum articulated torques on the hip joint and knee joint. Based on the theoretical calculation values for the maximum static torques for the hip joint and knee joint, it is deduced that the theoretical calculation value of the torque of the hip joint is 1"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000460_0022-4898(76)90005-7-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000460_0022-4898(76)90005-7-Figure1-1.png",
        "caption": "FIG. 1. Tire characteristics.",
        "texts": [
            " The different types must be characterized when attempts are made to unders tand the way of application of the efforts and how some effects are induced into the upper layer of the soil. Research tries to answer to such problems. Experiments are conducted in both statics and in dynamics upon plates and on different soils. They will be verified on soils in place. *Research of the Agricultural and Forest Mechanization Working Group subsidized by the Belgian Ministry of Agriculture, Declercq, D., I. Ag. Lv., Assistant. tD6partment de G6nie Rural, Facult6 des Sciences Agronomiques, Universit6 Catholique de Louvain, Belgique. 183 The parametersJor the casing (carcass) are (Fig. 1): H0 tire height without load, in mm; B0 tire width without load, in mm; C external circumference measured following the equator on the ribs or cleats, in mm ; E casing thickness for the side walls (Es) and for the tread (Eb) (between ribs), in mm; PR ply rating or ply equivalent number, mentioned by the manufacturer; Pi inflation pressure, in kPa. W load, in daN. ~B angle for the cleats. Commercially, a tire is defined only by information such as: tire construction type, mounting data, PR and some other dimensions. We must then proceed by testing to approach some more characteristics in order to give sufficient informat ion abou t a tire. Nevertheless one fundamenta l fact, the nature of the gums used and their influence on the deflection phenomena for a tire, remains unknown. Observations and measurements are made on: (1) Tire deflection under load, including (Fig. 1): Bw Width of the max imum enlargement of the tire for a given load, in mm, Hb level where the max imum enlargement occurs, measured f rom a rigid support , in mm, Hw height of the tire under load, measured between the rim and the rigid support , in ram. (2) Contac t trace on a rigid plate and on different soils: F Form of the trace : rectangle, ellipse, / max imum width of the contact area, in mm, L max imum length of the contact area, in mm, b dimensions of the ribs or cleats with their characteristics (reference: equator line of the tire), SVOT total area of the trace, m m 2, SEFF effective area of the trace, representative for the area actually support ing the load, m m ~"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000501_s00227-007-0891-x-Figure6-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000501_s00227-007-0891-x-Figure6-1.png",
        "caption": "Fig. 6 Larva in the Poiseuille flow; D is the diameter of a tube or the width of a channel. A rigid particle, which is located below the axis of the tube rotates with a positive angular velocity x",
        "texts": [
            " Using Eq. 29 to calculate the vorticity gives the following system of differential equations of motion of a larva in a linear shear flow: dn ds \u00bc 2f\u00fe 1 k cos w; \u00f030\u00de df ds \u00bc 1 k sin w Vt; \u00f031\u00de dw ds \u00bc 1 e sin w; \u00f032\u00de where n = XO /h, f = ZO /h, s = Ua t/h e \u00bc gbC 6m h Ua ; \u00f033\u00de k \u00bc Ua VS ; \u00f034\u00de Vt \u00bc Vt Ua ; \u00f035\u00de and Vt \u00bc \u00f0qp=q 1\u00deD2 p 18m g: \u00f036\u00de Consider now a laminar flow in a tube of radius R with an axis directed along the axis O1X of the fixed in space coordinate system (Poiseuille flow) (Fig. 6). The vector of the fluid velocity in the tube is represented as follows: U \u00bc 2Ua 1 Y2 \u00fe Z2 R2 jx: \u00f037\u00de The flow in the plane of symmetry of such a tube is identical to the flow in the two-dimensional Poiseuille channel of width D = 2R: U \u00bc 2Ua 1 Z2 R2 jx: \u00f038\u00de For a two-dimensional Poiseuille flow the non-dimensional equations of motion of a larva are the following: dn ds \u00bc 2\u00f01 f2\u00de \u00fe 1 k cos w; \u00f039\u00de df ds \u00bc 1 k sin w Vt; \u00f040\u00de dw ds \u00bc 2f e sin w; \u00f041\u00de where e \u00bc gbC 6m R Ua \u00f042\u00de n = XO/R, f = ZO/R and s = Uat/R"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000556_icma.2007.4303844-Figure4-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000556_icma.2007.4303844-Figure4-1.png",
        "caption": "Fig. 4 Serious link mechanism: Mechanism-A",
        "texts": [
            " Three mechanisms Mechanism-A, Mechanism-B and Mechanism-C which refer to this simplified musculoskeletal system are proposed Essential difference between these mechanisms is compared by computer simulation with a simulation model as shown in Fig. 3. In this model, each actuator which is equivalent to the muscle Ke and Ae, can be arranged on the joint K and A. The joint K and A are connected by a attachable wire which is equivalent to the muscle Bi. The position of center of gravity G is can be adjusted on the point B, C or K. Fig. 4 shows the Mechanism-A. Two actuators are arranged on the joint K and A independently. This mechanism is a robot like serious link mechanism equipped with the mono-articular muscles only, without the bi-articular muscle. In the Mechanism-A, each torque Tgk and Tga are given torque (actuatot torque) by the actuator to the joint K and A. The joint torque Tk and Ta are as the following, naturally. Tk = Tgk , Ta = Tga (1) In order for the ground reaction force to turn to the direction of center of gravity G, it is necessary to satisfy the following"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003713_978-1-4614-2419-2-Figure15.3-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003713_978-1-4614-2419-2-Figure15.3-1.png",
        "caption": "Fig. 15.3 Resulting strains at the inner-side strain gauges (positioned at radius of 208 mm from the wheel center) at 0 \u2218 , 90 \u2218 , 180 \u2218 , 270 \u2218 (left) and 45 \u2218 , 135 \u2218 , 225 \u2218 , 315 \u2218 (right) due to a constant contact pressure with slow rotation of the wheel",
        "texts": [
            " For the Q-bridges, the relation between the supply voltage vs and the output voltage (signal) vG can be computed as vG \u00bc 2\u00fe Ge3 4\u00fe G\u00f0e2 \u00fe e3\u00de 2\u00fe Ge4 4\u00fe G\u00f0e1 \u00fe e4\u00de vs (15.2) where e1 \u00bc e0 in \u00fe e0 out, e2 \u00bc e90 in \u00fe e90 out, e3 \u00bc e180 in \u00fe e180 out, e4 \u00bc e270 in \u00fe e270 out for the Q-bridge (a) and e1 \u00bc e45 in \u00fe e45 out, e2 \u00bc e135 in \u00fe e135 out, e3 \u00bc e225 in \u00fe e225 out, e4 \u00bc e315 in \u00fe e315 out for the Q-bridge (b) The combined output signal is defined as the sum of absolutes vout \u00bcj vG\u00f0a\u00de j \u00fe j vG\u00f0b\u00de j (15.3) For illustration, Fig. 15.3 shows the inner-side strains for the case of constant contact force for a slowly rotating wheel (20 km/h). The corresponding voltage signals vG\u00f0a\u00de , vG\u00f0b\u00de and vout are shown in Fig. 15.4. 15 Identification of Wheel-Rail Contact Forces Based on Strain Measurement. . . 171 The output signal obtained by applying a constant contact force p(t) \u00bc pconstant for a slowly rotating wheel vout constant p(t) is used to compute a calibration vector A(t) as A\u00f0t\u00de \u00bc pconstant vconstant p out \u00f0t\u00de (15.4) This calibration vector can then be multiplied by the output signal vout generated from any other contact force to estimate the applied force pstatic calibrated\u00f0t\u00de \u00bc A\u00f0t\u00devout\u00f0t\u00de (15"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000597_978-3-540-73129-0-Figure6.121-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000597_978-3-540-73129-0-Figure6.121-1.png",
        "caption": "Fig. 6.121. Micropump. After [324]",
        "texts": [
            " The dimensions of the base plate are 31 \u00d7 32mm2. Miniaturized lubricating systems are under development at the IMIT. The first application will be for improving the \u2018wick\u2019 lubricating process. In this process, a film of oil is carried by capillary action from a container to the part to be lubricated, typically a rotating part. Problems can sometimes arise when undesired excess lubrication and strongly varying oil consumption result from varying rotating speeds. A microsystem consisting of the dosing pump (as presented in Fig. 6.121), a microbuffer (volume< 5 mm3) and an oil sensor offers a viable solution [324]. The oil sensor measures the level in the buffer and the pump provides the lubrication as needed. Dosing of the smallest quantities of liquid on the order of nanoliters and microliters was the goal in the cooperation between the Research Centre of Rossendorf and the GeSiM company of Dresden in the development of a microdrop injector [325]. The unit consists of a micro-injection pump (MEP) and a microsieve functioning as a diode for liquids (Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002733_0954410012464002-Figure3-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002733_0954410012464002-Figure3-1.png",
        "caption": "Figure 3. Helical turn.",
        "texts": [
            "35\u201339 For maneuvering flight, the aircraft linear velocities are (Figure 1) u v w T \u00bc VA cos\u00f0 F\u00decos\u00f0 F\u00de VA sin\u00f0 F\u00de VA sin\u00f0 F\u00decos\u00f0 F\u00de T \u00f02\u00de where fuselage angle of attack, F, and sideslip, F, are given by F \u00bc tan 1 w=u\u00f0 \u00de; F \u00bc sin 1 v=VA\u00f0 \u00de \u00f03\u00de Level banked turn is a maneuver in which the helicopter banks towards the center of the turning circle. For helicopters, the fuselage roll angle, A, is in general slightly different than the bank angle, B. For coordinated banked turn A \u00bc B. A picture describing these angles for a particular case ( A \u00bc 0) is given in Figure 2, where Fresultant is the sum of the gravitational force (W) and the centrifugal force (Fcf). Helical turn is a maneuver in which the helicopter moves along a helix with constant speed (Figure 3). In a helical turn, the flight path angle is different than zero being given by sin\u00f0 FP\u00de \u00bc sin\u00f0 A\u00de cos\u00f0 F\u00de cos\u00f0 F\u00de sin\u00f0 A\u00de cos\u00f0 A\u00de sin\u00f0 F\u00de cos\u00f0 A\u00de cos\u00f0 A\u00de sin\u00f0 F\u00de cos\u00f0 F\u00de \u00f04\u00de A picture describing the flight path angle for a particular case, A \u00bc 0, F \u00bc 0, is given in Figure 4. Note that _ A 4 0 is a clockwise turn and _ A 5 0 is a counterclockwise turn (viewed from the top) while FP 4 0 is referring to the ascending flight and FP 5 0 is referring to the descending flight. Note: the numerical results reported from here on (i"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001284_s10015-006-0401-0-Figure3-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001284_s10015-006-0401-0-Figure3-1.png",
        "caption": "Fig. 3. Throwing motion of the manipulator",
        "texts": [
            " \u03b8 \u03b8 \u03b8 \u03c0 \u03c0 \u03b8j ij ij j j ijA t TB t T N= + + + \u2212    \u00b1 \u2212\u02d9 cos 1 2 1 1 2 2 \u2206 (7) The working time is divided into K elements, and the angular displacements are divided into N elements. In order to make the programming easy, we have calculated the trajectory under the condition that the angular acceleration is constant in each section. 4.1 Simulation for throwing motion We take the parameters of the system as shown in Table 1. The simulations of the system are done as follows. A response of the manipulator from the initial position (\u03b81 = \u2212\u03c0/4, \u03b82 = \u22123\u03c0/4) to the position of release (\u03b81 = 3\u03c0/8, \u03b82 = \u03c0/8) is shown in Fig. 3, under the conditions that the number for dividing are K = 8, N = 5, the distance from the origin to the point of arrival is x = \u22120.5m, the release angle is \u03c6 = 3\u03c0/4, and the working time is T = 0.6s. It is shown that the locus of every sampling time (0.02s) is like a pendulum movement.6 The angular velocities of the motors on the base are shown in Fig. 4. Because of the condition that the angular acceleration is constant in each section of searching, the angular velocity changes in the line-shape"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003359_s0025654413010020-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003359_s0025654413010020-Figure1-1.png",
        "caption": "Fig. 1.",
        "texts": [
            " The strongest difficulties arise in the control problem if the control resources are restricted and it is still required to stabilize the unstable equilibrium. The problem becomes more complicated if the number of controls in the system is deficient, i.e., if the number of degrees of freedom of the system exceeds the number of controls. Consider a system consisting of a wheel of radius R that can roll slip-free along a horizontal line and an n-link pendulum. The pendulum links and the wheel move in the same vertical plane. The first link of the pendulum is hinged to the wheel center O (Fig. 1). This hinge at point O will be referred to as the first hinge. The second link is hinged to the other end of the first link, and so on. The pendulum links are numbered successively with increasing distance from point O, and so are the hinges. Let \u03b20 denote the angle of counterclockwise rotation of a fixed (marked) wheel radius directed along the horizontal axis X at the beginning of the motion, and let x denote the displacement of the wheel center O along the horizontal line, so that x\u0307 = \u2212R\u03b2\u03070"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000074_j.gaitpost.2006.09.079-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000074_j.gaitpost.2006.09.079-Figure1-1.png",
        "caption": "Fig. 1. Model I: A modified compass gait model including ipsi-, contralateral knee flexion, pelvic rotation, and heel rise is applied at the time instant of minimum CoM height.",
        "texts": [
            " We found no difference between the excursions of the sacral marker and the computed CoM for the controls with full body marker sets (P > 0.60) or between the excursion of the sacral marker of the controls with lower body marker set and that of the CoM in the full body marker set (P > 0.20). The determinants of gait were computed using the methodology of Della Croce et al. [7]. This method computes the effects of the determinants using one model during double support and a simpler model during single support. The modified compass gait model (Model 1) shown in Fig. 1 was applied at the instant of time of minimum CoM height (during double support). This model included the definition of thighs and shanks instead of a rigid lower limb and a segment representing the pelvis. Three additional cylindrical joints describing ipsi- and contra-lateral knee flexion and pelvic rotation facilitated motion of these segments. A linear joint was also added to represent heel rise. For each trial the model geometry was defined using the 3-D position of each joint center determined from a subject\u2019s kinematics at the instant of minimum CoM"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000224_1-4020-2933-0_13-Figure25-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000224_1-4020-2933-0_13-Figure25-1.png",
        "caption": "Figure 25. A simple rigid-deformable multibody system.",
        "texts": [
            " We distinguish these nodes from the free nodes by designating the following superscripts to entities associated with a node: b for boundary; u for unconstrained (free). As an example, the stiffness and mass matrices associated with a deformable body can be partitioned as: [ ],KK K K KK KK K u,b, u, b, uuub bubb =\u23a5 \u23a6 \u23a4 \u23a2 \u23a3 \u23a1 =\u23a5 \u23a6 \u23a4 \u23a2 \u23a3 \u23a1 = [ ]u,b, u b, uuub bubb MM M M MM MM M =\u23a5 \u23a6 \u23a4 \u23a2 \u23a3 \u23a1 =\u23a5 \u23a6 \u23a4 \u23a2 \u23a3 \u23a1 = , . In this section we derive the equations of motion for rigid-deformable multibody systems. Without any loss of generality, we show the equations for a simple system containing one rigid and one deformable body connected by a spherical joint as shown in Figure 25. It is further assumed that arbitrary forces and moments also act on the system. In this Figure, subscript r refers to entities associated with the rigid body. The position constraints for the spherical joint can be described as: 0dsr =\u2212+ bb rr . (83) The velocity and acceleration constraints are: 0dsr =\u2212\u2212 b r b rr ~ , (84) 0sdsr =+\u2212\u2212 b rrr b r b rr ~~~ . (85) Using Eq. (70), the acceleration constraints can also be written as: bbb r b rr \u03b3\u03b4\u03c9\u03c9 =\u2212+\u2212\u2212 b ~ rs~r , (86) where bb rrr b ws~~ +\u2212= \u03c9\u03c9\u03b3 . (87) The equations of motion for the system can be written in a variety of forms"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003296_detc2013-13039-Figure34-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003296_detc2013-13039-Figure34-1.png",
        "caption": "Figure 34, F-35 STOVL Propulsion System",
        "texts": [
            " Lockheed Martin\u2019s entry in the X-35 STOVL (Shot Takeoff & Vertical Landing) aircraft for the US Marines and British Navy used a shaft driven Liftfan just aft of the pilot to provide forward augmented lift. This lift along with the lift of the engine exhaust nozzle and two roll post nozzles in the wings provided a stable \u201c4 post\u201d platform. Each propulsion system component is used as a function of flight condition and aircraft configuration. This is basically a balance between each component. See Fig 34 for the basic F-35 STOVL Propulsion System concept. The Liftfan system is designed around the turbine turning a shaft the drives a vertically mounted Liftfan. This capability is critical to the marine first strike units in battle. The engine driven drive shaft drives a pinion bevel gear that drives 2 larger counter rotating bevel gears that drive the 2 stage Liftfan. The large bevel gears in this actual hardware are the same size as the single bevel gear I used in the first test rig. This second rig was designed with the ability to rotate all 3 bevel gears in their correct orientation and RPM"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001196_0094-114x(72)90004-3-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001196_0094-114x(72)90004-3-Figure1-1.png",
        "caption": "Figure 1. An RCCC linkage in isometric projection.",
        "texts": [
            " The mechanisms that are considered here are single-loop, four-bar linkages, of which the joints are cylindric, revolute and prismatic. By applying this procedure to the PCCC chain as well, a complete list of single-degree-of-freedom linkages with these properties is obtained. Some of the linkages derived from the RCCC chain are reported for the first time in this paper, and all linkages are identified by closed form relations for their chain parameter values. 2. Unique Model Set An RCCC linkage is pictured in Fig. 1. It contains one revolute joint (R) and three cylindrical joints (C). For convenience, the revolute joint is assumed to be at the input joint (joint 1) and the fixed link is looked at as link 4. Each link i of the chain has a length, a~, which is the shortest distance between its two axes and a twist angle, m, which is a measure of the relative directions of its two axes. The rotations, 0i, and the translations, s~, define the relative positions of the links ( i - 1 and i) which are connected at joint i"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003709_b978-0-12-385971-6.00012-9-Figure12.25-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003709_b978-0-12-385971-6.00012-9-Figure12.25-1.png",
        "caption": "FIGURE 12.25 Diagram of a Seagen tidal power generator",
        "texts": [
            " Tidal power where tides are high can deliver 3 W/m2 by making both the incoming and the outgoing tidal flow drive turbines. This is about the same as wind power, but few countries have the coastline to capture much of it. To those with tidal estuaries or those that lie at the mouths of land-locked seas (like the Mediterranean) harnessing tidal power is an option. There are two schemes for harvesting it: tidal-stream and tidal-barrage systems. The Seagen tidal-stream power generator is an underwater wind turbine driven by the flow of the incoming and receding tides (Figure 12.25). One is in service. It has a power of 1.2 MW, a claimed capacity factor of 48%, and a design life of 20 years. A tidal barrage is a hydroelectric plant driven by a reservoir filled by tides rather than by rain. The largest tidal barrage is the 240 MW unit on the river La Rance in France, where the tidal range is 8 m. Tidal barrages have longer lifetimes than tidal-stream generators because the machinery is simpler. The attraction of tidal power is that it is completely predictable. The drawback, as Figures 12"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001344_icca.2007.4376700-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001344_icca.2007.4376700-Figure1-1.png",
        "caption": "Fig. 1. Two wheels model",
        "texts": [
            " The estimator requires measurement of relative yaw angle and yaw rate. Next, a method to design ideal transient trajectories for lateral distance and relative yaw angle is proposed, and then, an adaptive steering controller to achieve ideal trajectories is developed. The proposed controller doesn't require accurate knowledge of all vehicle parameters. Using the proposed controller, robust lane keeping can be assured for varying longitudinal velocity. II. CONTROLLED OBJECT Let's consider the simple two-wheel model shown in Fig. 1. Nomenclature of symbols is shown in Table 1. Assuming small lateral yaw angle r (t) and the situation where longitudinal velocity v. (t) is controlled just as designers hoped, dynamic equations are given as follows [9]. i3(t) =vx(t)cx b[x(t) + qp(t)- bxv(t)p(t) 4c (t) =VX (t)cx q(t + BC (u (t) - vx (t) 1 HpTT q, (t) ) Be = M-1HPTK, HPT = (T- )THP Xt()T =y[p(t), EFr(t)], U(t)T = ['Uf (t), U, (t)] qc (t)T =[V, (t), (t)],qp (t)T = [Vp (t), E (t) M diag[m, i,], K diag[cf, cr] cT= [1, 0], bT = [0, 1] I (1) (2) For the vehicle model (1) and (2), the following assumptions are made"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002776_09544097jrrt343-Figure16-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002776_09544097jrrt343-Figure16-1.png",
        "caption": "Fig. 16 Residual equivalent plastic strain after the fifth wheel passage at the full-scale roller rig with (a) no applied lateral load (Fx = 0) and (b) a lateral load of Fx = +15 kN",
        "texts": [
            " Rail and Rapid Transit JRRT343 at University of Birmingham on June 1, 2015pif.sagepub.comDownloaded from Here \u03b5 pl eq,0 is the initial equivalent plastic strain with \u02d9\u0304\u03b5pl = \u221a 2 3 \u03b5\u0307pl : \u03b5\u0307pl (17) The magnitude of \u03b5 pl eq during the fifth load cycle may be employed to roughly quantify the amount of ratcheting under the present test conditions. Note that, in accordance to the discussion in the introduction to chapter 5, \u03b5 pl eq should not be considered as a measure of the total plastic deformation at the end of the test. Figure 16 shows the residual equivalent plastic strain at the rail surface in the roller rig after the fifth wheel passage. JRRT343 Proc. IMechE Vol. 224 Part F: J. Rail and Rapid Transit at University of Birmingham on June 1, 2015pif.sagepub.comDownloaded from Test results from a full-scale roller rig, a full-scale linear test rig, and a twin-disc machine have been assessed and compared with respect to RCF. Two approaches have been adopted: an engineering criterion that utilizes evaluated elastic contact stresses and a full elasto-plastic modelling"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000246_978-3-540-88518-4_120-Figure6-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000246_978-3-540-88518-4_120-Figure6-1.png",
        "caption": "Fig. 6. The schematic diagram of the baby robot",
        "texts": [
            " As discussed above, the robot will experience complicated terrain under the mine well, so it is necessary to bring forth much more strict requirements for its ability of climbing over obstacles. Therefore, the baby robot is serially connected by multiple joints just like a snake. Otherwise, the baby robot takes communication nodes to improve the performance of wireless communication under the mine well. The baby robot generally includes a head unit, a tail unit, a control unit, a communication node cabinet and a power unit, as shown in Fig. 6. Each unit is connected by active and passive joints and this may enhance the ability of adaptation to the complicated terrain. Each of unit is equipped with motors, retarders and drivers and so on. The cameras are installed in the head unit and the tail unit, thus the image of forth and back of the baby robot can be transmitted back to the remote control center and it is propitious to the commander to make an accurate judgment. The control unit mainly controls the motion of baby robot, deals with the information of sensors and communicates with the mother robot"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001927_nme.1620030405-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001927_nme.1620030405-Figure2-1.png",
        "caption": "Figure 2. Internal and external degrees of freedom",
        "texts": [
            " As has been its meridional curve in dimensionless co-ordinates f and 7 may be expressed as where a, = tan IS, a, = tan p4 + 472\u201d a, = - ( 5 tan pi + 4 tan p,) + $7;- 75 a, = 3(tan pi + tan &) + $($- 7;) and I = cord length d2 I d t 2 r, C O S ~ p7\u201d = = -___ Note that the curve given by equation (1) satisfies the requirements of continuity of slopes and curvatures at the nodal circles. A REFINED CURVED ELEMENT 497 Displacement pattern co-ordinates, Figure 1, is assumed over each discrete element: In this development the following displacement model, expressed in terms of local Cartesian (2) 1 u1= a l + ~ . $ + a 3 . $ 2 + a p p us= a,+a65+Ly,p+a8p where a\u2019s are the generalized co-ordinates. The number of these generalized co-ordinates is equal to the total number of internal and external degrees of freedom of the element. The six degrees of freedom at the nodes i and j , Figure 2, are the external degrees of freedom and the two at the nodes m and n are the internal degrees of freedom. On assembling the elements into a representation of the over-all shell, compatibility must be maintained for all the displacement degrees of freedom occurring at the interzelement nodes, i.e. at i and j in Figure 2. The two displacement degrees of freedom at the internal nodes m and n are not required in the assemblage. Thus they must be removed by a process of static condensation prior to assemblage of the total structural stiffness matrix. 498 E. P. POPOV AND P. SHARIFI The displacement model (2) must be specialized for the case of the central cap, Figure 3. Here, if Wand U are the radial and tangential components of displacements at the apex, because of symmetry, the tangential component of displacement, Ui, and rotation, xi, vanish"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003349_10402004.2013.812758-Figure4-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003349_10402004.2013.812758-Figure4-1.png",
        "caption": "Fig. 4\u2014Schematic view of the cross-sectional positions within a period ( \u03b8 = 5\u25e6).",
        "texts": [
            " It is known that if the tridiagonal matrix A is strictly diagonally dominant, the linear equations can surely be solved by the Thomas method and the equation is ensured to converge. Once the temperature distribution is calculated by MATLAB, the temperature load will be imposed on the interface between the film and the seal face by using ANSYS codes. The interaction will continue until the change in film thickness meets the convergence criterion. The calculation procedure is shown in Fig. 3. DISCUSSION OF THE TEMPERATURE BOUNDARY The cross-sectional positions within a period are illustrated in Fig. 4. The temperature distribution in different cross sections is shown in Fig. 5. In Fig. 5 the temperature boundary condition S4 is equal to the seal cavity temperature T0. One can indicate that the temperatures at the film entrance, which are shown in elliptic wireframes, present discontinuities and that their temperatures are unreasonable. Because the film thickness is on the order of micrometers, the viscous heat generation is large D ow nl oa de d by [ U ni ve rs ity o f K en t] a t 1 6: 09 0 2 D ec em be r 20 14 Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000224_1-4020-2933-0_13-Figure6-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000224_1-4020-2933-0_13-Figure6-1.png",
        "caption": "Figure 6. A force and a moment acting on a body",
        "texts": [
            " The acceleration constraint for a joint can be expressed in the following general form: + =Dv Dv 0 . (10) Acceleration constraints contain quadratic velocity terms that are moved to the right-hand-side of the equations. Table 2 shows the right-hand-side quadratic terms for our four fundamental constraints. In the body-coordinate formulation, Newton-Euler equations provide the simplest description of the equations of motion. For a body with a mass mi , if the sum of forces acting on the body is denoted as if as shown in Figure 6, the Newton equations of motion describing the translation of the mass center are written as: i i i m =r f . (11) The Euler\u2019s equations describing the rotation of the body are normally expressed in the body-attached iii \u03b6\u03b7\u03be \u2212\u2212 reference frame as: i i i i i i = \u2212J n J\u03c9 \u03c9 \u03c9 , where Ji is a 3\u00d73 constant rotational inertia matrix. The Euler equations can also be described in the x\u2013y\u2013z frame as: i i i i i i = \u2212J n J , (12) where T i i i i =J A AJ is no longer a constant matrix. In order to be consistent with the constraint equations at velocity and acceleration levels, we choose the Euler equations from Eq"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003699_cjme.2013.02.257-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003699_cjme.2013.02.257-Figure2-1.png",
        "caption": "Fig. 2. Planar parallel mechanism",
        "texts": [
            " The mobility of all the 13 DELASSUS linkages is researched by using the Modified Gr\u00fcbler-Kutzbach Criterion here, the similar analyzing process of each DELASSUS linkage is omitted considering the length of the paper, the data are listed in Table 1. This work is a new progress in the mobility research of the classic mechanisms or the single-loop overconstraint linkages using the Modified Gr\u00fcbler-Kutzbach Criterion. The mobility of the modern parallel mechanisms has been well resolved with the Modified Gr\u00fcbler-Kutzbach criterion [7, 9\u201312]. However, a new planar parallel manipulator[22] is found with special nature. The planar parallel manipulator contains three branches, each of them with same structure, shown in Fig. 2. The moving platform is D1D2D3, the fixed platform is the part with joints Aij(i, j=1, 2, 3). All the joints are rotating pairs with parallel axes. According to the Modified Gr\u00fcbler-Kutzbach Criterion, in the mechanism, the number of bodies is n=17; the number of the joints g=21. A reference coordinate system o-xyz being introduced, assuming the z-axis is parallel with the axes of these rotation joints, then T(0 0 1; 0)i i iy x$   ( 1, 2, ,21)i  . So, d=3 is the order of the motion of the manipulator, furthermore, there is no non-common constraints in this mechanism, so 0\u03bd , according to Eq. (16), the mobility is 1 ( 1) 3(17 21 1) 21 6 . g i i M d n g f \u03bd \u03be \u03be \u03be                LI Yanwen, et al: Applicability and Generality of the Modified Gr\u00fcbler-Kutzbach Criterion \u00b7262\u00b7 Here, a new definition should be proposed. Half local freedom: If a part in a mechanism can receive the motions of its fore parts but can not transfer all the motions to its following-up parts, then the motions failed to transfer in the mechanism is called the half passive freedom. In Fig. 2, each sub-platform DiCi1Ci2 (i=1, 2, 3) has a half passive freedom for the rotation of the sub-platform DiCi1Ci2 (i=1, 2, 3) about the axis of joint Di can not be transfered to the platform D1D2D3. So, there are 3 degrees of half local freedom, \u03be =3, M=3. The number of active joints in the mechanism with half passive freedoms equals to the sum of the degrees of freedom of the output part and all the half passive freedoms. So, six active joints have been chosen, seen in Fig. 2. In other words, the screws of the 6 active joints are dependent, only three of them are independent. Thus, the number of the independent active joints equals to the degrees of freedom of the output part. (1) The general validity of the Modified Gr\u00fcblerKutzbach Criterion for mobility is elaborated or demonstrated in both theory and practice. (2) Some relative new terms, such as the half passive freedom, the non-common constraint space and the common motion space of a mechanism are proposed. (3) The mobility of the classical over-constrained singleloop DELASSUS mechanisms classified into thirteen types is determined using the Modified Gr\u00fcbler-Kutzbach Criterion"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000597_978-3-540-73129-0-Figure6.124-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000597_978-3-540-73129-0-Figure6.124-1.png",
        "caption": "Fig. 6.124. Piezoactuator with compliant gear",
        "texts": [
            " An integrated microsystem, designed at the Technical University of Ilmenau in Germany, is able to handle the tight mechanical tolerances of mono-mode waveguide couplings by a controlled adjustment. It basically contains a twoaxis microactuator for moving a fibre or a microlens, an optical sensor for position detection and a control circuit. The piezoelectric drive has a bimorph cantilever movable normal to the wafer surface [327]. Its second direction, the in-plane movement, employs a compliant mechanism in order to enlarge the very small strains of a piezoelectric monomorph. It contains a set of elastic hinges arranged as two-stage gear. Figure 6.124 shows the structure and the kinematic principle of the compliant gear. Electrostatic micromotors built in silicon based surface micromachining technology were first presented by Berkeley University in 1989 [328]. A typical example is shown in Fig. 6.125. The rotor built out of polycrystalline silicone has a diameter of about 100\u00b5m and includes a number of radial teeth. It is surrounded on its periphery by electrodes that can be addressed individually. Their number is larger than the number of teeth (e"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000178_isic.2007.4450908-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000178_isic.2007.4450908-Figure1-1.png",
        "caption": "Fig. 1. The relationship between COG, ZMP and reaction force",
        "texts": [
            " ZMP constraints are introduced to keep the robot from falling 1-4244-0441-X/07/$20.00 \u00a92007 IEEE. 339 down. 3. Robot 3D kinematic constraints are identified to give accurate angular trajectories of every DOFs. To use IPM as a model of humanoid robot, the following ideal assumptions are made: 1) A concentrated mass of robot resides at hip 2) Robot motions in sagittal and frontal planes are inde- pendent 3) The COG Position of IPM is constant in z-axis 4) The position of ZMP is linear to COG The relationship between ZMP, COG and reaction force is depicted in Fig.1. The position of COG is (X, H), the position of ZMP is (aX+b), and the normal vector of ground force is parallel to the vector (cX + d,H). The dynamics of the IPM can be given by Fx : Fz = x\u0308 : (z\u0308 + g) = (cx + d) : H (1) Because z = H , we can get x\u0308 = g(cx + d)/H (2) A. IPM for Humanoid Motion in the Sagittal Plane One step of humanoid walking consists of two phases, single support and double support phase. The COG motion during these two phases in the sagittal plane is shown in Fig.2. The IPM motion in the frontal plane can be obtained in a similar way"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003625_pime_proc_1966_181_022_02-Figure33-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003625_pime_proc_1966_181_022_02-Figure33-1.png",
        "caption": "Fig. 33. Arrangement of O-ring in groove",
        "texts": [
            " G Shear modulus, lbf/in2. h H -- r h* r H* _. Val 181 Pt I No 9 at WEST VIRGINA UNIV on June 4, 2016pme.sagepub.comDownloaded from C. J. HOOKE, D. J. LINES, AND J. P. O\u2019DONOGHUE Film thickness a t any point, in. Film thickness at point where - = 0. dP dx 1:. J P d+. Pressure at any point, lbf/in2. Sealed pressure. P 2G PS 2G Leakage past the seal, in3js. Net leakage over the sealing cycle. Seal section radius. Seal velocity, in/s. Viscosity, lbf s/in2. Coefficient of friction. Angle around the seal, defined in Fig. 33, degrees. Angular position of oil inlet, degrees. Angular position of oil cavitation, degrees. -. -. Glossary of terms Motoring or reverse stroke The stroke of the piston and seal away from the sealed fluid. Nip The difference between the outside diameter of the seal on its groove and the cylinder diameter. O-ring An elastomeric seal with a circular section. Pumping or furzcard stroke The stroke of the piston and seal towards the sealed fluid. The pressure of the sealed fluid. The difference between the maximum undersea l pressure and the sealed pressure. Sealed pressure Sealing pressure The first problem consisted in determining the pressure distribution around a seal due to confinement between the walls of its housing groove and the cylinder wall and to fluid pressure which was cxerted on the fluid side of the seal and in the space between the seal and cylinder wall when the seal was in motion (Fig. 33). The authors were delighted to find that this was a standard problem in elasticity and that a full solution is presented in most textbooks. The particular solution used by the authors is given in (14). The method of using this solution was to assume a pressure distribution at the confining walls of the seal housing, calculate the film shape, and modify the pressure where necessary so that the seal shape conformed to that of the housing. Proc Insm Mech Engrs 1966-67 The static-pressure distribution around the seal was found by finding an equilibrium position of the seal",
            " The pressure distribution was unaffected therefore by values of F up to The film shapes derived for one value of F, can be extrapolated to other cases, therefore, by use of the factor F1\u20192/F,1\u20192. An additional advantage from the calculator\u2019s point of view is that only one cycle of the elastohydrodynamic reiterative procedure Proc Instn Mech Engrs 1966-67 is required because the pressure distribution remains constant. A typical pressure distribution around the seal, together with the distorted seal shape, is shown in Fig. 33 for p8 = 1. For ease of presentation of the full results, the back-pressure distribution and the distribution of pressure at the seal-cylinder interface are shown separately in Figs 34 and 35. Fig. 34 shows how the extent of seal contact at the back of the seal groove is increased by the increase of sealed pressure. The limiting value of sealed pressure was determined when the seal contact at the back of the groove caused extension into the clearance region between the top of the groove and the cylinder wall",
            " In view of the above remarks it remains to be seen whether the performance characteristics depicted in Figs 36-40 may be deemed reliable in that their basis, i.e. the distributions calculated for dry-contact pressure (cf. Fig. 35), would be dependable enough. The next step, verifying the permissibility of the assumption about the complete identity of this distribution with that of the fluid pressures in the film, will not here be carried out. For how to set up such a verification the reader may consult (20). When the analytical formulation of the problem of assessing the three distributions of dry-contact pressure depicted in Fig. 33 is considered, in conjunction with the theory of elasticity as worked out in (14), it would appear that in the boundary conditions needed one may individually specify, for all these distributions, the contact pressures occurring at all of the six edges. But, at least for the distribution of the dry-contact pressures between the O-ring and the cylinder, the authors have added the boundary condition that the pressure gradient, dPldx, should vanish at the two edges with which we are here concerned (cf"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003571_s0263574711000324-Figure4-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003571_s0263574711000324-Figure4-1.png",
        "caption": "Fig. 4. Possible positions of the tip of elastic line with infinite number of modes.",
        "texts": [
            " By superimposing the particular solutions (28), any transversal oscillation can be presented in the following form: y\u0302to1(x\u03021,j , t) = \u221e\u2211 j=1 X\u03021,j (x\u03021,j ) \u00b7 T\u0302to1,j (t). (29) Bernoulli wrote Eq. (29) based on \u201cvision.\u201d Euler and Bernoulli did not define the mathematical model of a link with an infinite number of modes, but Bernoulli defined the motion solution (shape of an elastic line) of such a link, which is presented in Eq. (29). Euler and Bernoulli left the task of a link modeling with an infinite number of modes to their successors (see ref. [5, 6]). The equation of Bernoulli (29) (see Fig. 4) defines a geometrical position of any spot on the elastic body line y\u0302to1 in direction y1-axis, and in a direction of x1-axis it would be a x\u0302to1 coordinate that is also a geometrical size and it can be presented in an analog way as well as the size y\u0302to1. The position of a tip of a presented body with indefinite number of modes is defined by coordinates xto1, yto1 in x1, y1 level. It is supposed that all movements are made in x1 \u2212 y1 level, and a coordinate is z1 = 0 in this case. Equation (29) is actually the solution of dynamics of the presented body\u2019s movement during the time. However, in order to calculate the coordinates x\u0302to1, y\u0302to1 in some specific moment of time (as is seen from Fig. 4), it is necessary (except from angles \u03c91,1, \u03c91,2, \u03c91,3 . . . \u03c91,j ) to know sizes of elastic deformations of all modes yto1,1, yto1,2, yto1,3 . . . yto1,j . . . defined in a space of local coordination system xi,j , yi,j , zi,j . Generally, coordinates xto1, yto1 are the total of elastic deformations, but precisely, in geometrical terms, it is the total of projected elastic deformations on axes x1 and y1, respectively. Equation (29) has a significance as elastic deformation for each mode for Meirovitch (ref"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001866_bf00251592-Figure11-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001866_bf00251592-Figure11-1.png",
        "caption": "Fig. 11. St 9 mechanism with sliding joint at G; obtained by letting s', d', l ' (and poin. F) go to infinity",
        "texts": [
            " Six-bar coupler curves for the inversion of the Stephenson chain in which one triangular link is held fixed, have been considered by WUNDERLICH [23], who has also extended this investigation to 2n-bars derivable by means of this mechanism. In this case, the floating hinge pivots, associated with a binary link, have Six-Bar Motion. II 55 four-bar motion and circular motion, respectively, although the motion of a general point attached to the link describes a curve of order 18. RoBm~xs' theorem can be applied in this particular case [17]. 12. Stephenson-1 Mechanisms with Sliding Joint We consider the case shown in Fig. 11. The only change in the locus derivation of Section 2, involves equation (2), in which A' becomes ( X - m) e i ' ' and P ' becomes 2e, i.e. P no longer has an X Y term and is now a constant. The result is that the with effis'--I\" and r' cos ~\" \u2022s ' - -d\" order of the six-bar curve is reduced from 14 to 12 (with the 12 th degree terms coming from the terms .4.4 [.4 AD 2 + (.4 C - . 4 C~) 2 ] in the determinant of equation (5)), but the genus remains at 7. This is analogous to what happens in the Watt-1 mechanism (Part I, Section 12)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003246_1.4004116-Figure5-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003246_1.4004116-Figure5-1.png",
        "caption": "Fig. 5 Pitch curves of deformed limacon gear possessing several cycles",
        "texts": [
            " Supposing the number of its cycle is n1, the deformed pitch curve of the first gear on the ith cycle can be expressed as follows rdi1 \u00bc R\u00fe e cos u1 (20) udi1 \u00bc m1u1\u00fe 2\u00f0i 1\u00dep n1 \u00f01 m1\u00de u1 2 2\u00f0i 1\u00dep n1 ; 2\u00f0i 0:5\u00dep n1 m2u1\u00fe 2ip n1 \u00f01 m2\u00de u1 2 2\u00f0i 0:5\u00dep n1 ;2ip n1 8>< >: (21) where u1 is the polar angle of the first gear before deformation. If the period number of the conjugate gear is n2 and its pitch curve corresponding to that of the first gear is on the jth cycle, the equations of the pitch curve can be obtained as rdj2 \u00bc a R e cos u1 (22) udj2 \u00bc m1u2\u00fe 2\u00f0j 1\u00dep n2 \u00f01 m1\u00de u1 2 2\u00f0j 1\u00dep n2 ; 2\u00f0j 0:5\u00dep n2 m2u2\u00fe 2jp n2 \u00f01 m2\u00de u1 2 2\u00f0j 0:5\u00dep n2 ; 2jp n2 8>< >>: (23) where u2 \u00bc \u00d0 u1 0 i21 du1. The pitch curves of the deformed gear pair with n1 \u00bc 1 and n2 \u00bc 2 are compared with that of the initial pair in Fig. 5. The solid lines represent the initial gear pair and the dotted lines denote the deformed gear pair. 3.2 Tooth Profile of Deformed Limacon Gear. In theory, the tooth profile of noncircular gears is enveloped by a shaper when its pitch circle rolls on the pitch curves of noncircular gears without sliding. Many efforts are devoted to build the mathematical model for tooth profile of noncircular gears. In the sections below, the tooth profile formulas of deformed limacon gear pair are deduced on the basis of the mathematical model in literature [20]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001909_icsmc.2010.5641981-FigureI-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001909_icsmc.2010.5641981-FigureI-1.png",
        "caption": "Figure I. Simplified helicopter model with two inputs and two outputs.",
        "texts": [
            " The model predictive neural control of a helicopter is proposed in [15]. Several fuzzy logic control and neural network control strategies are compared in [16]. The effectiveness of these controllers is verified in simulations only. In this paper we proposed fuzzy controllers for the helicopter model and investigate their behaviors in both simulation and experimental modes. II. HELICOPTER MODEL DESCRIPTION In this section a mathematical model by considering the force balances is presented for a helicopter model Humusoft CE ISO (Fig. I). The helicopter system contains the helicopter body, the DC motors with permanent stator magnets, power amplifiers (PWMs), encoders as sensors and axel gear (represents load). In the following subsections the mentioned components will be described. A. Helicopter Model Assuming that the helicopter model is a rigid body with two degrees of freedom, the following output and control vectors are adopted: Y = [1f/',qJf U = [upu2f (1) (2) where: If/' is an elevation angle (pitch angle), qJ is an azimuth angle (yaw angle), u, and U2 are voltage of main and tail motors, respectively"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003520_j.euromechsol.2011.11.003-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003520_j.euromechsol.2011.11.003-Figure2-1.png",
        "caption": "Fig. 2. T2R1-type parallel manipulator with decoupled and bifurcated planar-spatial motion of the moving platform: constraint singularity (a), branch with planar motion (b) and branch with spatial motion (c); limb topology PkRtPtkRtRPtRkRtR-PkRtPS.",
        "texts": [],
        "surrounding_texts": [
            "The term of constraint singularity (CS) has been recently coined (Zlatanov et al., 2002) to characterize the configuration of lower mobility parallel manipulators in which both the connectivity of the moving platform and the mobility of the parallel mechanism increase their instantaneous values. From a constraint singularity, the mechanism can get out with or without branching (Gogu, 2008b). When branching occurs in a constraint singularity, the mechanism can reach different configurations, called branches, and have different independent motions of the moving platform. In this case, the constraint singularity is also called branching or bifurcation singularity. A branch refers to the free-of-singularity configurations of the mechanism inwhich each structural parameter keeps its value. For this reason, this value is called global of full-cycle value for a branch. Two types of branching in constraint singularity (BCS) have been defined in Gogu (2008b). Branching of type BCS1 occurs when a parallel mechanism F ) G1 . Gj . Gk gets out from a constraint singularity in different branches characterized by the same mobility and the same connectivity of the moving platform but with different bases of the vector space of relative velocities between the moving and fixed platforms. Branching of type BCS2 occurs when a parallel mechanism F ) G1 . Gj . Gk gets out from a constraint singularity in different branches characterized by distinct values of mobility and connectivity of the moving platform. The parallel manipulators presented in this paper have constraint singularities with branching of type BCS1. Branching of type BCS2 occurs in kinematotropic mechanisms (Gogu, 2009c). The finite displacements and the velocities in the actuated joints are denoted by qi and _qi, the linear velocities of the characteristic point H of the moving platform, by v1 \u00bc _x; v2 \u00bc _y and the angular velocity of the moving platform by u \u00bc ua \u00bc _a or u \u00bc ud \u00bc _d: In both branches, the moving platform undergoes two planar translations and one rotation but the rotation axis is different in the two branches. In the first case, the rotation velocity u \u00bc ua \u00bc _a is parallel to the x-axis and the moving platform undergoes a spatial motion. In the second case, the rotation velocity u \u00bc ud \u00bc _d is parallel to the z-axis and the moving platform undergoes a planar motion. With these notations, the linear mapping between the actuated joint velocity space and the end-effector velocity space for a branch the T2R1-type PMs is defined by: 2 4 v1 v2 u 3 5 \u00bc \u00bdJ 2 4 _q1 _q2 _q3 3 5 (8) where J is the Jacobian matrix. In the constraint singularity, the T2R1-type parallel manipulators with bifurcated motions has instantaneously iM \u00bc iSF \u00bc 4 and (iRF)\u00bc(v1, v2, ua, ud). The bifurcation in this constraint singularity can be used to change motion type of the moving platform. In the two distinct branches, the parallel mechanism is characterized by M\u00bc SF \u00bc 3 and (RF)\u00bc(v1, v2,ua) or (RF)\u00bc(v1, v2,ud). In the constraint singularity, both connectivity of moving platform and mobility of parallel manipulator increase their instantaneous values. More details about the behaviour of parallel manipulators in a constraint singularity and the associated structural parameters can be found in Gogu (2009c)."
        ]
    },
    {
        "image_filename": "designv11_12_0000027_j.optlastec.2007.10.010-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000027_j.optlastec.2007.10.010-Figure2-1.png",
        "caption": "Fig. 2. Schematic arrangement of transverse rupture strength testing.",
        "texts": [
            " Transverse rupture strength (TRS) is an ultimate strength index, which is used as standard in carbide industry. TRS is a combination of shear strength, compressive strength and tensile strength and generally used as a measure of the toughness of the cemented carbide material. Three-point flexural testing is employed to compare the transverse rupture strength of sintered cemented carbide and ALFa fabricated WC\u2013Co samples, as per ASTM B406-96 (2005). The schematic arrangement of the testing setup is presented in Fig. 2. Universal testing machine (Instron: 2401) is used to carry out this test. The strain rate +0.1mm/min during the test. In order to understand the effect of processing parameters, a number of clad tracks were fabricated at different process parameters. The visual appearance of the tracks and cross-sections of laid tracks were examined and correlated with the process parameters. Similar to previous observations, it was seen that the process parameters for overlapped pulsed material deposition could be grouped into four parameters: average energy per unit area (Ea), peak power density (Ip), spot overlap (SO) and track overlap index (i) [26]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001939_s12283-009-0033-4-Figure5-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001939_s12283-009-0033-4-Figure5-1.png",
        "caption": "Fig. 5 Local forces with respect to local axis. Vertical forces are shown as bold arrows",
        "texts": [
            " The V GRF pattern closely followed the total GRF pattern, while the ML GRF pattern deviated from this mould. The ML GRF was relatively low during initial contact through contralateral push-off as the ipsilateral skate was gliding in a straight direction to guide the skater, thus not generating a ML force on the blade holder. Subsequently, as the ipsilateral limb externally rotated and extended through push-off, the skate boot collapsed medially, generating a ML force acting on the blade holder. Local forces with respect to the skates\u2019 local axis during push-off can be seen in Fig. 5. From previous research it can be discerned that the skate is oriented approximately 30 from vertical during the push-off [15]. Based on this, interpretation of the action\u2013reaction force dynamics with respect to the skate can be assumed. The V reaction force was oriented along the frontal plane of the skate, approximately 30 from vertical, while the reaction ML force was directed laterally with respect to the skate. The action force vectors can subsequently be described as the opposite of the reaction forces"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003895_ls.171-Figure14-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003895_ls.171-Figure14-1.png",
        "caption": "Figure 14. Steady-state pressure distributions for rigid and compliant gas bearings, (e0 = 0.80,\u039b = 1.07).",
        "texts": [
            "11 Effects of elastic deformations on the steady-state and dynamic gas-film pressures Table IV shows the details of bearing geometry and operating conditions of a sample problem investigated in the present study. The dimensionless parameters calculated from numerical values given in Copyright \u00a9 2012 John Wiley & Sons, Ltd. Lubrication Science 2012; 24:95\u2013128 DOI: 10.1002/ls Table I are as follows: \u039b= 1.07, a = 0.4, R/L= 0.5 and C/R= 2 10 3, which are the compressibility number, the compliance factor and the aspect and clearance ratios of the journal bearing, respectively. Figure 14 depicts the steady-state pressure profiles and contours calculated in the half bearing for a highly loaded journal bearing operating at e0 = 0.8. It is observed that the effect of the bump-foil elasticity leads to a spreading of the pressure distribution in the circumferential direction of the bearing over a greater area and to an important reduction of the peak pressure inducing a reduction of the journal bearing carrying capacity. The increasing of the fluid-film thickness over the whole bearing area explains the pressure drop"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003713_978-1-4614-2419-2-Figure61.11-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003713_978-1-4614-2419-2-Figure61.11-1.png",
        "caption": "Fig. 61.11 Base-upright FEM with master DOF selection",
        "texts": [
            "0 in some cases, depending on the scale of the analytical and experimental vectors along with the actual mass matrix.) 612 L. Thibault et al. A representative helicopter wing structure was used to demonstrate the proposed technique. This laboratory structure is comprised of 2 in. thick aluminum plates attached together with two steel L-brackets. This structure will be referred to as the Base-Upright (BU). An FEM of the BU was developed and consists of over 40,000 DOFs. The base plate of the structure is tied to ground in all 6 DOF at four locations, as shown in Fig. 61.11. An eigensolution is performed on the analytical BU model to determine the mode shapes of the full-space model. These mode shapes are then extracted at the master DOF points shown in Fig. 61.11 and consists of 78 DOF, which are in the x, y, and z directions. This set of 78 DOF will be considered as the n-DOF set for the cases discussed in this paper. The full 40,000 DOF BU model is then reduced down to a master set of 52 DOF, which consists of selected directions at each point, denoted by the arrows in Fig. 61.11. This set of 52 DOF will be considered as the a-DOF set for the cases discussed in this paper. The a-DOF set is also the measurement directions that were used to acquire experimental data in the cases that use VIKING to expand test data. The typical mode shapes for the BU model are shown in Fig. 61.12 for reference. 1 in. wide by 2 in. thick steel rib, as shown in Fig. 61.13. The rib is considered to be a realistic, but significant modeling error, especially in the case of a helicopter wing model"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001643_09544062jmes1340-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001643_09544062jmes1340-Figure2-1.png",
        "caption": "Fig. 2 (a) Meshing of two conical gears with an imaginary common rack and (b) pitch cone model",
        "texts": [
            " The constants A and B are determined from the geometry of surfaces, that is A = 1 4 ( \u03ba 1 + \u03ba 2 \u2212 \u221a g 2 1 + 2g1g2 cos 2\u03b7 + g 2 2 ) (22) B = 1 4 ( \u03ba 1 + \u03ba 2 + \u221a g 2 1 + 2g1g2 cos 2\u03b7 + g 2 2 ) (23) with \u03ba i = \u03baIi + \u03baIIi, i = 1, 2 (24) and gi = \u03baIi \u2212 \u03baIIi, i = 1, 2 (25) The skew gear pair used in this study is composed of a helical and a conical gear. The design parameters of the skew conical\u2013helical gear drive under investigation are (a) assembly parameters: shaft angle and offset d; (b) gearing parameters: tooth number z1 and z2; cone angle \u03b81, helix angle \u03b21 of the conical gear; and helix angle \u03b22 of the helical gear. In general, the meshing of a conical gear pair is regarded the same as that two engaged conical gears with an imaginary working common rack of zero thickness [5, 6] (see Fig. 2(a)). Since the common rack also defines the topology of each conical gear, the spatial arrangement of the gear axes can thus be determined from the relation between the common rack and the gears. The pitch cone and pitch plane model are employed for derivation of the assembly relation (Fig. 2(b)) [5]. The general assembly relations can be derived along with the working gearing parameters (subscript w) from the working common rack as follows [5, 6, 11] cos = cos \u03b8w1 cos \u03b8w2 cos(\u03b2w1 + \u03b2w2) \u2212 sin \u03b8w1 sin \u03b8w2 (26) d = (rCw1 cos \u03b8w2 + rCw2 cos \u03b8w1) sin(\u03b2w1 + \u03b2w2) sin (27) If the working common rack differs from the rackcutter used for gear generation, this is a case of profile-shifted transmission [16, 17]; otherwise it is a case of standard transmission. The assembly relations for a skew conical\u2013helical gear drive in a standard transmission, where \u03b82 = 0, are expressed as cos = cos \u03b81 cos(\u03b21 + \u03b22) (28) d = (rC1 + rC2 cos \u03b81) sin(\u03b21 + \u03b22) sin (29) In general, the line of action between the base cylinders of a skew cylindrical gear pair is unique, as shown in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001148_s11012-008-9178-7-Figure6-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001148_s11012-008-9178-7-Figure6-1.png",
        "caption": "Fig. 6 Bench-scale piston ring reciprocating liner test rig and ring holder with force sensor assembly (from [26])",
        "texts": [
            " A full transient elastohydrodynamic solution of the finite line conjunction of cam to flat-follower, combined with the multi-body dynamics of the valve train system, is reported in [15]. The results show again that the quasi-steady solution underestimates the film thickness. A reduction of the computing time is possible with the multi-grid and multi-level multi-integration techniques [16]. The piston ring-cylinder liner contacts have been investigated experimentally using electrical [17, 18] and optical [19] methods. Laser fluorescence [20] has been used also for measuring film thickness. Friction is measured in real engines [21\u201323] and in simulation rigs [24\u201326], Fig. 6. The importance of considering elastohydrodynamic effects, particularly near the top dead centre, for a correct evaluation of the film thickness has been proven numerically in [27]; squeeze and starvation effects are also important. Cavitation and transition to mixed and boundary lubrication conditions are included in a transient hydrodynamic model in [26]; the calculated values of the friction coefficient are in good agreement with the experimental ones obtained with a piston ring reciprocating liner test rig, as shown in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001287_cdc.2007.4434251-Figure8-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001287_cdc.2007.4434251-Figure8-1.png",
        "caption": "Fig. 8. Illustrating proof sketch of Lemma IV.1. The shaded regions are the capture discs of pk\u22121 and pk .",
        "texts": [
            " For the want of space, we provide the sketch of the proofs to Lemmas III.3, IV.3 and IV.5. We refer the reader to the report [13] for the complete proofs. We say that a daisy-chain with separation sip is closed if for some pk (k 6= 1), there exists a tk \u2264 sip and a solution \u03b7 : [0, tk] \u2192 R 2 of equation (1) satisfying \u03b7(0) = p1, \u03b7\u0307(0) = vp,1, \u03b7(tk) = pk, \u03b7\u0307(tk) = vp,k, such that the evader cannot move between p1 and pk without being captured. We first sketch the proof of Lemma III.3. Proof sketch of Lemma III.3: As shown in Figure 8, we consider the evader\u2019s motion in a reference frame attached to the center O of the circle of radius \u03c1 through pursuers pk\u22121 and pk and rotating with angular speed 1 \u03c1 in the direction of pursuer motion. We show that if the interpursuer separation does not exceed s\u2217ip(\u03b3, \u03c1), then any evader trajectory from arc UW to PQ must enter the capture ball of pursuer pk\u22121 or pk. We compute the analytical expression for s\u2217ip(\u03b3, \u03c1) by determining the optimal trajectory for the evader in polar coordinates with origin at O that maximizes the radial distance covered by the evader for a given angular displacement"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001262_icnsc.2008.4525353-Figure8-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001262_icnsc.2008.4525353-Figure8-1.png",
        "caption": "Fig. 8. Lead and tilt angles",
        "texts": [
            " A pen is attached to the robot arm, and the robot tries to reach the surface of the work piece from the predefined 2D coordinate. If the robot hits the surface, the z position is stored. These 3D positions are known as the cutting points p' (index w= work piece reference coordinate system). In each cutting pointp' we determine the surface inclination in the feed direction xn as n+l _ n n pn+l n PW _-PW the tilt angle (a) is the angle between the surface normal zn and the tool axis zv in a direction perpendicular to the feeding direction Fig. 8 shows the lead and tilt angles. (3) Surface inclination y, in a direction perpendicular to the feed direction is determined by n2 nl n2 nl PW -PW (4) Then the surface normal zn in the cutting point is found as Zn = Xn XYn (5) In each cutting point p the directional cosines X,x , zn forms a surface coordinate system Kn. To collect all parameters a (4 x 4) transformation matrix is created Xn 0 T = (6) Pw 1 The matrix (4) represents the transformation between the cutting point coordinate system Kn and the work piece reference coordinate system KW"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003267_j.jfluidstructs.2012.01.004-Figure14-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003267_j.jfluidstructs.2012.01.004-Figure14-1.png",
        "caption": "Fig. 14. Three-dimensional view of wake field.",
        "texts": [
            " 1 and 2(b)) modified the boundary layer through a similar mechanism, though less dramatically. It is noted that the phenomena depicted on different cross-sectional planes occur simultaneously due to the non-axisymmetric seam pattern, and that the boundary layer and the wake flow field are fully three-dimensional even in the time-averaged sense. Note that side force variations shown in Figs. 3\u20135 at various orientations about three axes can occur simultaneously and manifest themselves as complex three-dimensional vertical and horizontal forces. Fig. 14 schematically illustrates such three-dimensional wake pattern at one ball orientation. The wake is reduced and its pattern is symmetric in the horizontal plane (also see Fig. 6(a)), but in the vertical plane asymmetric seam has deflected the wake upward with an enlarged recirculation region, generating a negative lift. While we have focused our attention to the side force variation that affects the ball trajectory most, it is worth revisiting the drag force. Variation in drag force and that in the wake size are well correlated as reported widely for spheres and sports balls"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003033_s11191-012-9502-4-Figure4-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003033_s11191-012-9502-4-Figure4-1.png",
        "caption": "Fig. 4 Be\u0301lidor\u2019s model of friction where the surfaces are represented as made of many spherical rigid asperities (Architecture Hydraulique, 1737)",
        "texts": [
            " He criticized the authors who treat mechanics by neglecting friction, because As the effect of friction is much greater than it is imagined and one cannot make an exact calculation of any machine without studying it scrupulously \u2026 it is necessary to apply to consider well [this chapter] as one of most essential of this work (p. 70). He considered that friction force equals the force needed to make a body surmount the asperities of the other body in contact with it, and introduced an interesting representation of surface roughness, where the surface was made of spherical rigid asperities (Fig. 4). He considered that the force R necessary to make one layer of spheres surmount another similar layer was independent of the number of the spheres, and therefore from the area, in agreement with the Amontons\u2019 law. So, he considered the case of only one sphere supported by three spheres of the other body (Figs. 4, 5 of Be\u0301lidor here reported in Fig. 4). He found that the ratio between such a force R and the weight of the body is equal to the ratio between the segments GT and GO of figure. By means of elementary geometrical reasoning, he calculated that the square of this ratio equals 3/24 and then the ratio GT/GO equals 173/489, a value which he remarked to be about 1/3, similarly to the Amontons\u2019 measures (in modern notations we can say that his calculation corresponds to a friction coefficient 1/H8 % 0.35). The idea of spherical roughness would be reconsidered in more recent times, in a more complex way (Archard 1957, see Sect"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001289_robot.2008.4543428-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001289_robot.2008.4543428-Figure2-1.png",
        "caption": "Fig. 2. Model of kneed passive walker with feet",
        "texts": [
            " Moreover, many researchers have proposed robots based on passive walking [9]-[15]. However, they have not considered the mechanism of leg-swing motion of passive walking. Therefore, we focus on the flexion and extension of knee joint of swing leg. In this paper, first, an equation of angular acceleration of knee joint is derived from the simplified and linearized model of passive walking. Secondly, we demonstrate the mechanism of the flexion and extension of knee joint of swing leg. Finally, we demonstrate the influence of leg and foot on its mechanism. Figure 2 shows the model of kneed passive walker with feet. The model consists of stance and swing legs. The knee of the stance leg is locked straight. The motion is assumed to be constrained to the saggital plane. For the purpose of simplicity and clarity of analysis, as possible, we give a 978-1-4244-1647-9/08/$25.00 \u00a92008 IEEE. 1588 simplification of the model as follows: M \u226b m, M \u226b m1, M \u226b m2 (1) In addition, we assume that inertia moments of thigh and shank are very small. 1) Motion equation of 3 links (with Knees): Stance leg is assumed to be fixed on the ground with no slippage or take off"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001180_02726340701272154-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001180_02726340701272154-Figure1-1.png",
        "caption": "Figure 1. Cross-section of proposed induction motor.",
        "texts": [
            " The results of applying TSFE method and those obtained by application of WF theory are compared. Mechanical faults in an induction motor (such as fatigue of the ball-bearing) disturbs the conformity of the symmetrical axes of the rotor and stator as well as rotor rotation axe. In the static eccentricity, the rotor rotation axis coincides with its symmetrical axis, but displaces to the stator symmetrical axis. In this case, the air gap surrounding the rotor misses the uniformity but it is time-independent. Figure 1 shows the cross section of the motor with asymmetrical air gap. In the static eccentricity, one point on the inner surface of the stator has the minimum and the opposite point has the maximum air gap length. However, the stator sees a fixed air gap length independent of the rotor rotation. In this D ow nl oa de d by [ K or ea U ni ve rs ity ] at 0 7: 00 2 8 D ec em be r 20 14 case, the air gap can be considered as a cosine function against the rotor angular position relative to a static axis on the stator"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000597_978-3-540-73129-0-Figure6.61-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000597_978-3-540-73129-0-Figure6.61-1.png",
        "caption": "Fig. 6.61. Microstructured NiTi actuator mounted to the silicon surface of the gripper [90]",
        "texts": [
            " SMA Actuated Miniature Silicon Gripper. By using flexure hinges the micro gripper is designed in a compliant mechanism. To provide the gripper with a centering capability a four-bar-linkage mechanism with a transmission of \u22121 between input and output crank is used, where the ends of the cranks represent the gripping jaws [89]. The SMA actuators are connected to one crank, forming a serial differential type actuator. A parallel movement of the gripping jaws can be achieved with two additional linkages (see Fig. 6.61). The micro gripper in Fig. 6.61 consists of a silicon structure with a dimension of approximately 7\u00d7 4 mm2. In the open position the gripping jaws are 0.5mm apart. The flexure hinges have a minimum thickness of 30 \u03bcm. By machining a sputtered NiTi foil the SMA actuator has been realized with a minimum thickness of 30 \u03bcm. The gripping force averages by circa 11mN. High-precision positioning devices are often required for micro system production. Therefore miniaturized robots or fine positioning systems are built up using flexure hinges in order to increase accuracy and resolution"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002717_0954406212454390-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002717_0954406212454390-Figure1-1.png",
        "caption": "Figure 1. Cross sections of two kinds of internal gear pump: (a) involute internal gear pump; (b) conjugated involute internal gear pump.",
        "texts": [
            " Flowrate characteristic, conjugated involute, internal gear pump Date received: 23 February 2012; accepted: 20 June 2012 Gear pumps are the most often employed powersupplying components in hydrostatic systems, allowing low cost, good efficiency and high reliability. Based on structures, gear pumps are classified into two types: external gear pumps and internal gear pumps.1 Internal gear pumps have less flowrate pulsation than external gear pumps, which makes them operate quietly and attract attention in applications where people concern environmental noise.2,3 Involute internal gear pump is the main product type on the market, cross section of which is shown in Figure 1(a). Figure 1(b) presents another kind of internal gear pump, conjugated involute internal gear pump. The only difference between them lies in the tooth profile of internal gear. To describe briefly, we call the pump in Figure 1(a) as \u2018conventional pump\u2019 and the one in Figure 1(b) as \u2018conjugated pump\u2019. It is known that the profile of pinion mainly comprises an involute and a root fillet. In the conjugated pump, the whole profile of internal gear is completely conjugated to that of pinion. While in the conventional pump, there is no curve conjugated to the root fillet in the profile of internal gear. Additionally, its involute is only partly conjugated to that of pinion. Consequently, clearances among meshing teeth are larger in the conventional pump than those in the conjugated pump (Figure 1). These clearances directly decide the volumes of trapped fluid which induces large pressure ripples and high noise levels to hydrostatic systems. Since the conjugated pump has smaller volumes of trapped fluid, it will consequently have better operating performances than the conventional pump. Investigations into gear pumps have been done for years, most of which concern the flowrate performances of external gear pumps. Manring and Kasaragadda3 calculated the instantaneous flowrate State Key Laboratory of Fluid Power Transmission and Control, Zhejiang University, Hangzhou, P"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001903_physreve.79.011705-Figure14-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001903_physreve.79.011705-Figure14-1.png",
        "caption": "FIG. 14. Schematic illustration as to how to solve Eq. 9 .",
        "texts": [
            " A5 and A6 , z / implies the average of the eight corresponding derivatives in the cells containing the vertex i , j ,k \u201ceight\u201d should be replaced by \u201cfour\u201d when k=0 or N . But just for simplicity, in our numerical calculations, we evaluate z / at x= i x, y= j y and =k in the calculation of Eq. A6 . In this appendix we first give a brief explanation as to why 0 =90\u00b0 at =90\u00b0 twist deformation is absent when qA= /160 0.0196 and /16 0.196, while 0 90\u00b0 twist deformation is present for qA=3 /16 0.589. To this end, we illustrate schematically in Fig. 14 how to solve Eq. 9 when =90\u00b0. In Fig. 14, the solution of Eq. 9 corre- sponds to the intersection of the curve representing f\u0303a ( 0 ) with a straight line with a slope \u2212K2 /Lz. Notice from the symmetry of the system that f\u0303a ( 0 =0\u00b0 )= f\u0303a ( 0 =90\u00b0 ) =0. For the cases we consider, f\u0303a ( 0 ) 0 for 0\u00b0 0 90\u00b0, as is evident from Figs. 1\u20133. From Fig. 14, we readily find that when f\u0303a ( 0 =90\u00b0 ) K2 /Lz, with fa ( 0 )=dfa ( 0 ) /d 0 , Eq. 9 has no solution other than 0 =90\u00b0. The above condition implies weak surface anchoring or small cell thickness or small Lz . On the other hand, when f\u0303a ( 0 =90\u00b0 ) K2 /Lz, there exists a solution 0 90\u00b0, which corresponds to the presence of twist deformations in the bulk. We note that 0 90\u00b0 always gives lower total free energy than that for 0 =90\u00b0, because from Eq. 8 the former equals the area of region a in Fig. 14, while the latter equals the sum of the areas of regions a and b . Therefore, so long as a solution 0 90\u00b0 exists, it always minimizes the total free energy; 011705-11 in other words, the existence of a 0 90\u00b0 results in twist deformations in the bulk. From Eq. 3 , we have f\u0303a ( 0 =90\u00b0 )=\u2212KA2q3, in which K is a function of K1, K2, K3, and Ks and has the same dimension as that of those elastic constants 42 . From the discussion above, spontaneous twist deformations in the bulk exist when KA2q3 K2 /Lz"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000685_j.triboint.2007.02.012-Figure5-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000685_j.triboint.2007.02.012-Figure5-1.png",
        "caption": "Fig. 5. Grid geometry for the analysis of stresses.",
        "texts": [
            " And also these results show that the numerical method used in this study can be applied to the analysis of stresses for coated film having different modulus of elasticity from the substrate\u2019s one. The calculation conditions are listed in Table 1. For the analysis of oil film, the analytical area is 4.5px/ap1.5 and 3.0py/ap3.0, respectively. The analytical area is divided into nx \u00bc ny \u00bc 240 in x and y directions. The number of nodal points is 58081. For the analysis of stresses, the analytical area is 4.5px/ap1.5, 3.0py/ap3.0 and 28.0pz/ap0.0 as shown in Fig. 5, respectively. The analytical area is divided into n \u00bc 60,768,000 number of tetrahedron elements. When coated film and substrate have different modulus of elasticity, the value of radius of Hertzian contact area a cannot be calculated by Hertzian contact theory easily, so the value of a, in cases where the coated film and substrate have same modulus of elasticity and Poisson\u2019s ratio is used in our study. Thus, the value of radius of Hertzian contact area a is constant, the value of Ti/a varies by changing the value of coated film thickness Ti"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000939_19346182.2008.9648458-Figure3-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000939_19346182.2008.9648458-Figure3-1.png",
        "caption": "Figure 3. Positions of the center of mass and the center of buoyancy of a typical human body in a horizontal, motionless floating position. Buoyant force acts more cranial to the center of mass, generating a leg-sinking moment around the center of mass of the body.",
        "texts": [
            " The stability of a human body in a horizontal, motionless floating position is determined by how the body is configured to form a certain posture and by the composition of the parts of the body. As many of us may know from our own experiences, our body is not generally stable in a horizontal, motionless floating position: the legs tend to sink to a lower position than the initial horizontal position. Studies confirm that the legs in fact tend to sink. This is due to the buoyant force acting more cranial to the CM of the body (Figure 3), generating the moment around the CM that causes the legs to sink [7,8\u201312]. Some of these studies examined the sex differences in the body\u2019s stability and found that women tend to float more horizontally than men [10,11,13], due primarily to women having a greater amount of body adipose tissue stored around the hips and thighs, causing the CB to be located closer to the CM [11]. The state of breathing was found to affect the stability in a horizontal, motionless floating position [10,12]. When air is inhaled, the lung volume increases and the CB shifts cranially to increase the leg-sinking moment"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003803_j.conbuildmat.2015.07.152-Figure6-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003803_j.conbuildmat.2015.07.152-Figure6-1.png",
        "caption": "Fig. 6. O-Y PSM [3].",
        "texts": [
            " As specified by Federal Specification KK-L-165C, the shoe material used in ASTM D 2047 is a piece of cowhide 76.2 76.2 mm in size and 6.4 mm in thickness that should be uniformly ground with a #400 abrasive paper prior to use. ASTM D 2047 does not mention the Shore A hardness of the test specimen, but we employed a piece of cowhide with Shore A hardness of 80 in our study. ASTM D 2047 prescribes that the test be performed on a cleaned floor. 3.3. Test prescribed in JIS A 1454 This test uses the O-Y PSM (O-Y Pull Slip Meter) shown in Fig. 5 to obtain the slip resistance coefficient. Fig. 6 shows the O-Y PSM, a device that reproduces the contact between the sole of the footwear and the floor and the load applied to the sole as the foot is being lifted. The lifting movement is mentioned here because it was verified that, in terms of slipperiness, there is no noticeable change in the ranking of sample floors between the landing and lifting phases of stepping. As Fig. 7, the O-Y PSM puts a \u2018\u2018slip piece\u2019\u2019, a sample of a sole cut from a piece of footwear, in contact with the floor, loads the weight, and measures the load as the slip piece is pulled diagonally upward"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002218_j.ymssp.2010.12.001-Figure3-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002218_j.ymssp.2010.12.001-Figure3-1.png",
        "caption": "Fig. 3. z1(T) and z2(T) with T (A1=2.0 mm/s, A2=1.5 mm/s, f1=15 Hz, f2 f1=0.92 Hz, and j1=0, j2=0).",
        "texts": [
            " (12) contains z1(T) with frequency of 2o and z2(T) with frequency ofo2 o1, in addition to \u00f0A1=2\u00decosj1. \u00f0A1=2\u00decosj1 is the necessary part to calculate A1 and j1 according to the correlation method described in Part 2. z1(T) and z2(T) are induced by x2\u00f0t\u00de \u00bc A2 sin\u00f0o2t\u00fej2\u00dewhen x(t) is correlated with y1s\u00f0t\u00de \u00bc sino1t. Thus we define z1(T) and z2(T) as additional error, and try to minimize them to zero. The error can be expressed as a function of T and the relationship between them is indicated in Fig. 3. Apparently, zi(T) (i=1,2) tends to minimize to zero with the increase in integral time. Theoretically, the longer the integral time, the smaller the error. However, the longer integral interval requires not only larger data storage space but also longer processing time. The memory requirement and computational speed of modern hardware, especially the PC, is surely not the limitation for the advanced technology. However this research aimed at providing theoretical basis for developing the portable device with Single Chip Microcomputer (SCM) for field balancing with low cost",
            " It can be calculated as the maximum of z1(T) and z2(T): z1max \u00bc A1\u00feA2 2To1 , z2max \u00bc A2 T\u00f0o2 o1\u00de Define a\u00bc z1max z2max \u00bc \u00f0A1\u00feA2\u00de\u00f0o2 o1\u00de 2A2o1 \u00bc 1 2 1\u00fe A1 A2 o2 o1 1 \u00f013\u00de The amplitudes A1 and A2 are close in magnitude. In this way only the vibration can be a beat form (otherwise the vibration with larger amplitude will be referred as noise). Thus the value of 1\u00fe\u00f0A1=A2\u00de is small. \u00f0o2=o1\u00de 1 is about zero sinceo1Eo2. Consequently, a\u00bc \u00f01=2\u00de\u00f01\u00fe\u00f0A1=A2\u00de\u00de\u00f0\u00f0o2=o1\u00de 1\u00de is zero, nearly, namely z1max5z2max. Thus, it is prior to diminishing z2(T) to zero. When TA \u00bc 2p=\u00f0o2 o1\u00de (one cycle of the beat, that is point A in Fig. 3), z2(TA)=0, there will only be a residual error caused by z1(T), ascertaining a pocket-size error. This correlation method with one cycle as integral time is called the Whole-Beat Correlation Method. As exhibited in Part 3.2, the residuary error is only induced by z1(T) when the integral interval is a beat period. Substituting TA \u00bc 2p=\u00f0o2 o1\u00de into z1(T) gives z1\u00f0TA\u00de \u00bc o2 o1 8po2 A1 sin 4po1 o2 o1 \u00fej1 sinj1 \u00feA2 sin 4po1 o2 o1 \u00fej2 sinj2 \u00f014\u00de Obviously, the maximum z1max is z1max \u00bc o2 o1 4po2 \u00f0A1\u00feA2\u00de \u00f015\u00de Introducing Eq",
            " (19), because the uttermost processing appears in the derivation of Au 1max. For instance, when the requirements of sin\u00f0\u00f04po1\u00de=\u00f0o2 o1\u00de\u00fej1\u00de sinj1 \u00bc 2 and sin\u00f0\u00f04po2\u00de=\u00f0o2 o1\u00de\u00fej2\u00de sinj2 \u00bc 2 in Eq. (14) are met at the same time, the z1max can be achieved. In addition, in Eq. (18), if and only ifj1 \u00bc 451, could A1 u reach the Au 1max, remaining other, A1 u being A1 uoAu 1max. Define TB \u00bc n\u00f02p=o2\u00de, TC \u00bc \u00f0n\u00fe1\u00de\u00f02p=o2\u00de, and n\u00f02p=o2\u00deo2p=\u00f0o2 o1\u00deo \u00f0n\u00fe1\u00de\u00f02p=o2\u00de. TB and TC represent points B and C, respectively, in Fig. 3, that is, the whole period of sampling and z1(TB)=0, z1(TC)=0. The interval between TB and TC is small compared to the period of z2(T); thus z2(T) can be noted as approximately linear in this interval. It can be obtained that Z2\u00f0TB\u00de=\u00f0TA TB\u00de Z2\u00f0TC\u00de=\u00f0TA TC\u00de or Z2(TB)(TC TA)+Z2(TC)(TA TB)E0 Define a1 \u00bc TA TB TC TB \u00bc 2p=\u00f0o2 o1\u00de n\u00f02p=o2\u00de 2p=o2 \u00bc no1 \u00f0n 1\u00deo2 o2 o1 a2 \u00bc TC TA TC TB \u00bc \u00f0n\u00fe1\u00de\u00f02p=o2\u00de 2p=\u00f0o2 o1\u00de 2p=o2 \u00bc no2 \u00f0n\u00fe1\u00deo1 o2 o1 When the integral interval is selected as TB and TC, respectively, we can get RxsB \u00bc \u00f0A1=2\u00decosj1\u00fez2\u00f0TB\u00de and RxsC \u00bc \u00f0A1=2\u00decosj1\u00fez2\u00f0TC\u00de"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002342_j.mechmachtheory.2011.06.006-Figure3-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002342_j.mechmachtheory.2011.06.006-Figure3-1.png",
        "caption": "Fig. 3. A multibody chain associated to the cut spatial prismatic joint i.",
        "texts": [
            " After imaginary cutting of the joint, the motion of body (Vi) relative to its preceding adjacent body (Vh) can be described by a vector of the translational displacement u\u2192= \u2192 OhO j i = qn+1 \u03bb \u2192 + qn+2 \u03bc\u2192 + qi v \u2192 \u00f028\u00de and the relative rotation of body (Vi) about point Oi j described, as in Section 3.2, by Bryant's angles qn+3, qn+4, and qn+5. Bodies (Vh) and (Vi) interacting by the forces Pi e \u2192 i and\u2212Pi e \u2192 i, where Pi is the driving force in the ith joint. Now, bodies (Vh) and (Vi) can be replaced by kinematic chain consisting of bodies (Vh), (Vf), (Vn+1), \u2026,(Vn+4), and (Vn+5)\u2261(Vi) interconnected by revolute or prismatic joints (see Fig. 3), where joint e\u2192i is actuated while joints e\u2192n+1;\u2026; e\u2192n+5 are unactuated. Bodies (Vf), (Vn+1), \u2026,(Vn+4) are fictitious. Based on this, it is obtained that \u03f1\u2192f = \u03f1\u2192n+1 = \u2026 = \u03f1\u2192n+4 = 0 \u2192 ; \u03f1\u2192n+5 \u2261 \u03f1\u2192i \u00f029\u00de and for qn+1=\u2026=qn+5=0: e\u2192n+1 0 = e\u2192n+3 0 = \u03bb \u2192 ; e\u2192n+2 0 = e\u2192n+4 0 = \u03bc\u2192 ; e\u2192n+5 0 = \u03bd\u2192 : \u00f030\u00de For the joint type considered, Eqs. (23), (24), and (25) take the following form \u03c9\u2192 p = \u03c9\u2192p + \u0398i; p\u00f0 \u00de \u2211 5 \u03b1=3 q n+\u03b1 e\u2192n+\u03b1; p = i;i + 1;\u2026;n; \u00f031\u00de v\u2192 Cp = v\u2192p + \u03c9\u2192p \u00d7 \u03f1\u2192Cp + \u2211 2 \u03b1=1 q n+\u03b1 e\u2192n+\u03b1 + \u2026\u00f0 \u00de + \u2211 5 \u03b2=3 q n+\u03b2 e\u2192n+\u03b2 "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003882_9781118181249-Figure1.5-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003882_9781118181249-Figure1.5-1.png",
        "caption": "Figure 1.5 Two examples of mechanical signal transduction: (a) static microcantilever defl ection where an SRP coating on the topside of the cantilever causes an increase in surface stress upon analyte binding. Right graph: surface stress plotted as a function of time for two glucose - sensitive coatings. The SRP coating provides a faster and more sensitive response upon glucose binding (10 mM) compared with self - assembled monolayers (SAMs). (b) Dynamic transduction mode using a QCM sensor. The SRP coating on the quartz crystal provides a scaffold for the analyte to bind, which causes a decrease in the resonance frequency. Right graph: frequency plotted as a function of time, showing the frequency response when a streptavidin - coated crystal (S) is exposed to biotinylated BSA (B). After a wash step (W) and exposure to anti - BSA (A), a further drop in resonance frequency is observed. Adapted with permission from references [105] and [119] . (See color insert.)",
        "texts": [
            " Volumetric transformations in SRPs involve stretching and contraction of polymer chains to accommodate the change in the system \u2019 s osmotic pressure. For sensing applications, SRPs are typically immobilized on a solid support, and this surface confi nement restricts the ability of the polymer coating to expand or to contract freely as compared to the bulk polymer, and leads to the generation of large interfacial stresses. These stresses are harnessed for sensing applications using, for example, SRP - coated microcantilevers, where the stress - induced cantilever bending can be amplifi ed and recorded (Fig. 1.5 a). 13,111,112 Optical (optical lever and interferometry) and electrical (piezoresistive and piezoelectric) detection schemes are commonly used to monitor microcantilever defl ection. For example, in optical lever defl ection detection, laser light is refl ected from the back of the microcantilever onto a position - sensitive photodetector, and small cantilever defl ections are thus translated into easily measurable detector voltages. For piezoresistive defl ection detection, the change in resistance of a piezoresistive material embedded in the microcantilever is measured. This approach allows the measurement of large cantilever defl ections and obviates the need for a complex alignment procedure, which is often a serious problem in optical - based detection methods. Another mechanical transduction approach is provided by quartz crystal resonators that translate changes in mass on the crystal surface into measurable frequency changes (Fig. 1.5 b). For these sensing approaches, the SRP coating serves as a matrix for direct analyte binding, which increases the coupled mass, 72,105 \u2013 107 or as a matrix that swells due to interaction with the surrounding solvents, which increases the water retention and thus the overall mass. 113 The transduction approach for quartz crystal microbalance (QCM) relies on the shift of resonance frequency of an oscillating quartz crystal and the change of mass on the QCM crystal surface. 114 The sensitivity of QCM thus relies on the accurate measurement of changes in the crystal \u2019 s resonance frequency",
            "1 Examples of SRP Sensors That Use Mechanical Transduction Principles For mechanical transduction schemes, a reactive and analyte - specifi c SRP coating in contact with the transducer surface is needed to detect chemical or biological molecules. 111,115 For microcantilever sensing, several types of SRP coatings, including hydrogels, 116,117 polymer brushes, 118 \u2013 120 and LBL fi lms, 121,122 have been employed. For example, Hilt et al. 116 coated microcantilevers with cross - linked pAA - PEG copolymer hydrogels to detect pH changes. The Zauscher group 118,119 demonstrated the use of pNIPAM copolymer brushes, grown on one side of a microcantilever, to detect changes in pH, ionic strength, and glucose concentration (Fig. 1.5 a). Similarly, Yan et al. 122 used cantilevers modifi ed with GOx/PEI multilayers to detect glucose. Glucose concentration was also measured by QCM using a glucose - sensitive hydrogel copolymer ( m - acrylamidophenylboronic acid - co - acrylamide) electropolymerized on QCM electrodes. 72 Biotin - functionalized hydrogel coatings for QCM sensors were developed for protein immobilization and antibody assays, and, for example, the specifi city of the coating for bovine serum albumin ( BSA ) antibody was measured in real time using QCM sensors (Fig. 1.5 b). 105 In addition, MIP coatings that are specifi c to nucleotides, 107,123 monosaccharides, 107 and herbicides 106 were developed, and the specifi city of the imprinted molecular template to the target was verifi ed using QCM. QCM also provides a convenient technique to study the conformational change in SRP coatings in real time. 39,113,124 For example, Richter et al. 39 used pVA - co - pAA hydrogels, spin coated on the quartz crystal, to map the hydrogel \u2019 s swelling hysteresis in response to changes in pH"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003082_0022-2569(70)90069-8-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003082_0022-2569(70)90069-8-Figure1-1.png",
        "caption": "Figure 1.",
        "texts": [
            " 2 are related as follows: cos a ' = [sin (X + tx)/cos/3]2+ [sinZX + sin2(X + tx) tan\"/3] - sinZo~ 2[sin (h + a)/cos/3] [sin~-~ + sin2(X + a) tan2/3] ,/2 tan/3 tan/.t = sin X sinB' = sin tz cos (h +o~) COS ,u, tan h' = tan (X + a)\" (2.1) (2.2) (2.3) (2.4) J.M. Vol. $, No. 3--H 396 The angle 8~, which the intersection of the plane P and the plane defined by line 1 and C D makes with L is given by tan 8~ = tan/3 cot A. (2.5) Our first task is to express the variable angle ~5 as a function of the independent variable 0. Using the notation in Fig. 1, it can be shown that tan (~) K ~ / K ' Z - - ( A 2 - l ' Z ) = ( A + i ) (3.1) where and J = a cos/3 cos 0 cos a -- a sin 0 sin/3 -- b cos ,~ cos/3 A = ( l \" + J \" + K 2 - c 2 + c F ) / 2 d = / c o s S + K s i n & From (3.2), (3.3) and (3.4), I - '+J\"+K\"- = (a'- '+b z) - 2 a b cos 0 c o s (A + pc). Thus , (3.2) (3.3) (3.4) (3.5) (3.6) A = (aZ+b2) - (c -\"- an) - 2 a b cos 0cos (,~ + pc) 2d (3.7) F rom (3. I) we see that if 8 is to be real (i.e. if the angle 8 exists for a given input angle 0, s o that the mechanism can be put together in this configuration), then K ~ - (A 2-1- ' ) ~ 0",
            " This means that the R - S - S-R linkage can only be assembled in one particular configuration and is a structure. (d) The ellipse intersects the circle in two points. Then the portion of the ellipse within the circle defines a range of possible input angles. (e) The ellipse intersects the circle in four points. Then there exist in general two distinct ranges for the input angle 0. A limiting case could exist where the ellipse intersects the circle in two points and touches the circumference at one point. Two ellipse diagrams for the cases AB as input (Fig. 1), and CD as input (Fig. 2), completely define the input and output angle ranges of motion. The parameter changes for CD as input have been described in Section 2. 399 Figure 4 shows an ellipse diagram for the input angle 0 of an R - S - S - R mechanism (Fig. 1). For any point on the ellipse, we know that X - - L = M cos 0 or x~ = M cos 0 (4.1) (4.1) therefore gives cos 0 at any instant. Also, f rom (3.17) and (3.13), Yl = N cos 0 - P sin 0. (4 .2 ) 400 From (4.1) and (4.2), we may eliminate cos 0 to give sin 0 in terms ofx l and Yr. sin 0 = { ( N x , / M ) - -Yt} /P . (4.3) From Section 3, A -- d = cL + cM cos 0 = cX. Thus A = c X + d . (4.4) From (3.2) and (4.1), t/ X' -'\u00a21 1 = a sin od .--: I + b s i n h. \\ M / (4.5) From (3.4) and (4.1), (4.3), (x,) K = a --~ sin /3 cos ot + a[ N x t / M - yl] /P",
            " The lines I and K (DC and F E + G H ) are drawn at right angles to one another. At B we draw a line I perpendicular to AB. Then with centre A and radius D H we describe a circle intersecting l in two points. Then the angles 81 and 82 (Fig. 5) are the two possible output angles, depending upon how the mechanism is assembled initially. In both cases, it can be seen f rom the construct ion that A B = DC cos S + F H sin 8 or A = I cos 8 + K sin 8. 401 In analysing an R-S -S -R mechanism with A D as base (Fig. 1), we may draw the two ellipse diagrams including the associated four straight lines, for inputs A B or CD. These diagrams then give limiting positions of the chosen input, and the output angle 8 is then given simply by the construction shown in Fig. 5. Thus, only one ellipse diagram needs be drawn to describe the input-output characteristics of the mechanism. However, for a synthesis, both ellipse diagrams (for 0 and 8) will often be needed. In the next sections, when we speak of the \"ellipse diagram\", it should be understood that the lines A, I and K are included",
            " 10(c), we see that 8 lies in the range .1.150 \u00b0 to -1-t-70 \u00b0 approximately. As a check, we can construct the ellipse for d as input. This is actually shown in Fig. 4, but it is shown again in Fig. 11 for the sake of clarity. From Section 2, we get our new set of parameters with d as input, i.e. - i F f I I ', ', : ', \" ~ Figure lO(a). I\" /',/c p lii ~i i ~ I ,,,~, ,,, J.M. Vol. $. No. 3- [ 413 h' =--7 .4\u00b0 ; c~' = 33.9\u00b0; /3' = 24\"9\u00b0; From which ~ = 36-6\", T = 0.25, S = 0.125 and 8s becomes 106\". A small sketch (Fig. 1 I) gives the relationship between input angle 8 for d as input, output 8 for a as input link, and 8s. From this, it is seen that 8 -- 2\u00a2r - (8' +Ss). I fA and B are the points of intersection of the ellipse with the circle X 2 + Y~ = 1, then from Fig. 11: cos S' (A) = 0.12; c o s S ' ( B ) = - l - 0 0 . After checking the sign of sin (8') at these two points, we get: 8 ' (A) = +850; 8 ' (B) = + 180\". The range expressed in terms of 8 is then given by: 8(A) = 169\"; 8(B) =+74* . These results agree with those found approximately from Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003996_icmeae.2013.22-Figure10-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003996_icmeae.2013.22-Figure10-1.png",
        "caption": "Figure 10. Flow vs Level",
        "texts": [
            " Let, 1 ' 11 hHH (13) 2 ' 22 hHH (14) Where, H\u2019 is the normal operating level, and is a constant, h is a small change about that level. Then, for small variations of h about H, the nonlinear equations (11) and (12) can be approximated by the straight-line tangent at H\u2019 in (8), (9) and (10) , in the same way the inflow Qi consist of a steady component Q\u2019i plus a small change qi. Then the equations (11) and (12) can be rewritten iibbaa qQqQqQ dt dHA '''1 (15) aacc qQqQ dt dHA ''2 (16) The slope of the valve characteristics at the level H\u2019 is given by equation (17), see Fig. 10. h q dH dQ (17) Rewritten equations (15) and (16) using (17), ii b b aa a qQ dH dQhQ dH dQh dH dQhQ dt dhA ''' 1 1 2 2 1 1 1 (18) 2 2 1 1 2 2 2 '' dH dQh dH dQhQ dH dQhQ dt dhA aa a c c (19) When the level is constant, with qi=0, h1=0 y h2=0, the equations (18) and (19) give the steady state relation for flow and level, where iba QQQ ''' (20) ac QQ '' (21) Subtracting equation (20) from (18), and (21) from (19) and defining; 2 4 1 3 2 2 1 1 ,,, dH dQD dH dQD dH dQD dH dQD cbaa Solving h1 from (19), getting time derivative and replacing on (18), and then rearranging gives the linear, second order differential equation for the coupled tanks system"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001012_10402000903097387-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001012_10402000903097387-Figure1-1.png",
        "caption": "Fig. 1\u2014Schematic of journal bearing.",
        "texts": [
            " The extended form of the Reynolds equation is derived in which the slip is allowed to occur on the bearing surface. An infinitesimal perturbation technique (Lund and Thomsen (9)) is utilized to determine the dynamic force coefficients. The effects of the slip parameters on the bearing characteristics are discussed. The pattern of the slip/no-slip regions on the bearing surface will lead to improved bearing stability characteristics. The results of the threshold speed and the critical whirl ratio are provided for L/D = 1. The journal bearing configuration is shown in Fig. 1. The extent of the slip zone (\u03b8s) is defined from the vertical load line, and the region of the slip on the bearing surface extends both in the cavitation and the full-film region. The bearing is stationary while the journal surface moves with a speed uj . The journal surface is a conventional material, so that the no-slip condition applies, while some areas on the bearing surface allow the slip. The methodology used in the derivation of the extended unsteady Reynolds equation considering a slip on the bearing surface is quite similar to that employed in Spikes (4) and is given in Appendix A",
            " Hence, the presence of the slip zone on the bearing surface lowers the magnitude of the steady-state pressure profile within the extent of the slip zone. The bearing load capacity is a function of force components in the radial and tangential directions, and the force components are obtained by integrating the pressure profile, which is influenced by the extent of the slip zone as shown in Eqs. [B1]\u2013[B2]. Using the short-bearing approximation, the stiffness coefficients in the XY coordinate system (Fig. 1) are expressed as Kxx = \u22122 ( L D )2 \u03b8s\u222b 0 [ A2 + (1 + A)2 (1 + 4A)(1 + A) ] sin(\u03b8 + \u03c6) cos(\u03b8 + \u03c6) H3 0 d\u03b8 + 6 ( L D )2 \u03b50 \u03b8s\u222b 0 [ (1 + 2A)3 (1 + 4A)2 (1 + A) ] sin \u03b8 cos2 (\u03b8 + \u03c6) H4 0 d\u03b8 \u2212 2 ( L D )2 \u03c0\u222b \u03b8s sin(\u03b8 + \u03c6) cos(\u03b8 + \u03c6) H3 0 d\u03b8 + 6 ( L D )2 \u03b50 \u03c0\u222b \u03b8s sin \u03b8 cos2 (\u03b8 + \u03c6) H4 0 d\u03b8 [B4] Kyx = \u22122 ( L D )2 \u03b8s\u222b 0 [ A2 + (1 + A)2 (1 + 4A)(1 + A) ] sin2(\u03b8 + \u03c6) H3 0 d\u03b8 + 6 ( L D )2 \u03b50 \u03b8s\u222b 0 [ (1 + 2A)3 (1 + 4A)2(1 + A) ] \u00d7 sin \u03b8 sin(\u03b8 + \u03c6) cos(\u03b8 + \u03c6) H4 0 d\u03b8 \u22122 ( L D )2 \u03c0\u222b \u03b8s sin2(\u03b8 + \u03c6) H3 0 d\u03b8 + 6 ( L D )2 \u03b50 \u03c0\u222b \u03b8s sin \u03b8 sin(\u03b8 + \u03c6) cos(\u03b8 + \u03c6) H4 0 d\u03b8 [B5] Kxy = 2 ( L D )2 \u03b8s\u222b 0 [ A2 + (1 + A)2 (1 + 4A)(1 + A) ] cos2(\u03b8 + \u03c6) H3 0 d\u03b8 + 6 ( L D )2 \u03b50 \u03b8s\u222b 0 [ (1 + 2A)3 (1 + 4A)2(1 + A) ] \u00d7 sin \u03b8 sin(\u03b8 + \u03c6) cos(\u03b8 + \u03c6) H4 0 d\u03b8 +2 ( L D )2 \u03c0\u222b \u03b8s cos2(\u03b8 + \u03c6) H3 0 d\u03b8 + 6 ( L D )2 \u03b50 \u03c0\u222b \u03b8s sin \u03b8 sin(\u03b8 + \u03c6) cos(\u03b8 + \u03c6) H4 0 d\u03b8 [B6] Kyy = 2 ( L D )2 \u03b8s\u222b 0 [ A2 + (1 + A)2 (1 + 4A)(1 + A) ] sin(\u03b8 + \u03c6) cos(\u03b8 + \u03c6) H3 0 d\u03b8 + 6 ( L D )2 \u03b50 \u03b8s\u222b 0 [ (1 + 2A)3 (1 + 4A)2 (1 + A) ] sin \u03b8 sin2(\u03b8 + \u03c6) H4 0 d\u03b8 +2 ( L D )2 \u03c0\u222b \u03b8s sin (\u03b8 + \u03c6) cos (\u03b8 + \u03c6) H3 0 d\u03b8 + 6 ( L D )2 \u03b50 \u03c0\u222b \u03b8s sin \u03b8 sin2(\u03b8 + \u03c6) H4 0 d\u03b8 [B7] Using the short-bearing approximation, the damping coefficients in the XY coordinate system (Fig. 1) are expressed as Bxx = 4 ( L D )2 \u03b8s\u222b 0 (1 + A) (1 + 4A) cos2 (\u03b8 + \u03c6) H3 0 d\u03b8 + 4 ( L D )2 \u03c0\u222b \u03b8s cos2 (\u03b8 + \u03c6) H3 0 d\u03b8 [B8] Byx = Bxy = 4 ( L D )2 \u03b8s\u222b 0 (1 + A) (1 + 4A) sin (\u03b8 + \u03c6) cos (\u03b8 + \u03c6) H3 0 d\u03b8 + 4 ( L D )2 \u03c0\u222b \u03b8s sin(\u03b8 + \u03c6) cos(\u03b8 + \u03c6) H3 0 d\u03b8 [B9] Byy = 4 ( L D )2 \u03b8s\u222b 0 (1 + A) (1 + 4A) sin2(\u03b8 + \u03c6) H3 0 d\u03b8 + 4 ( L D )2 \u03c0\u222b \u03b8s sin2 (\u03b8 + \u03c6) H3 0 d\u03b8 [B10] Similarly, the magnitude of the factors ( A2+(1+A)2 (1+4A)(1+A) ), ( (1+2A)3 (1+4A)2(1+A) ), and (1+A) (1+4A) used in Eqs. [B4\u2013B10] are always less than unity"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000399_s10999-008-9077-z-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000399_s10999-008-9077-z-Figure1-1.png",
        "caption": "Fig. 1 A typical contact element",
        "texts": [
            " Contact elements are formulated and assembled into the original FE code in order to enforce the contact conditions (Bo\u0308hm 1987). The solution is then obtained by solving the resulting set of nonlinear equations. A large number of contact element formulations has appeared in the literature and has been implemented in a number of commercial FE packages over the last two decades (see for example Chaudhary and Bathe 1985). The solution techniques adopted in these formulations were based upon using either function method or Lagrange multipliers in identifying the contact surface and imposing the contact constraints. Figure 1 shows a typical example of such contact elements. This particular element is adopted in several FE packages including ANSYS (ANSYS 1999) and MARC (MARC 1993). In this case, the element is based on two stiffness values. They are the normal contact stiffness KN and the tangential contact stiffness KT. The normal stiffness KN is used to penalize interpenetration between the two bodies, while the tangential stiffness KT is used to approximate the sudden jump in the tangential force, as represented by Coulomb\u2019s friction law when sliding is detected between the two contacting nodes"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002399_s12206-012-0524-2-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002399_s12206-012-0524-2-Figure1-1.png",
        "caption": "Fig. 1. Bending fatigue test apparatus at room temperature.",
        "texts": [
            " Moreover, the fracture surfaces of disconnected wire ropes were observed using scanning electron microscopy (SEM) in order to determine their characteristics. Appropriate wire rope was chosen for the experiment so that we could determine the effect of bending fatigue on fracture strength and wire rope life. Wire rope made of KS D 3559 hard steel SWRH37A 8\u00d7S(19)+SS was selected with two different diameters, \u03a68 mm and \u03a610 mm. Wire rope is made of 8 strands, each of 19wires, the strands being laid around a fiber core. The chemical composition of these wire ropes is shown in Table 1. The wire rope bending fatigue tester used is shown in Fig. 1. The tester rotates a motor shaft to move three sheaves in the *Corresponding author. Tel.: +82 53 810 2401, Fax.: +82 53 810 4621 E-mail address: jdkwon@yu.ac.kr \u2020 This paper was presented at the ICMR2011, Busan, Korea, November 2011. Recommended by Guest Editor Dong-Ho Bae \u00a9 KSME & Springer 2012 horizontal direction. This machine can apply three levels of bending by hanging the wire rope on the sheave, and five wires can be tested simultaneously. During the test, its driving velocity was 20 cycle/min and the tensile load was varied within the range of fracture strength"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002776_09544097jrrt343-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002776_09544097jrrt343-Figure1-1.png",
        "caption": "Fig. 1 Left: Definition of vertical load, lateral load, and semiaxes of the Hertzian contact patch. Right: Definition of Hertzian contact pressure p and interfacial shear under full-slip (dashed) and partial-slip (dotted) conditions",
        "texts": [
            "3 shows, the conformal contact conditions for the full-scale test rigs make it inherently difficult to carry out wheel\u2013rail dynamics simulations to establish these parameters. The adopted approach is to employ the dimensionless fatigue index, FI surf [7]. FI surf is defined as the horizontal distance from the curved line in the shakedown diagram, which indicates when the interfacial shear stress exceeds k (see reference [7] and Fig. 7) FIsurf = f \u2212 2\u03c0abk 3Fn > 0 (1) Surface-initiated RCF is predicted when FI surf >0. Definitions of a, b, and Fn are given in Fig. 1. Definitions of f and k are given below. For line contact (as in twin-disc tests), equation (1) will lose its validity because a \u2192 \u221e. Instead it is expressed in the form FIsurf = f \u2212 k p0 > 0 (2) where for line contact p0 = F \u2032 n/(\u03c0b), where F \u2032 n is the contact load per unit length. Modified shakedown diagrams to better account for line contact are presented in the literature [10]. In the current study, such modifications have not been employed. Thus, to evaluate the fatigue index FI surf , the following parameters need to be evaluated or known: 1"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000615_978-1-4020-8829-2_1-Figure5-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000615_978-1-4020-8829-2_1-Figure5-1.png",
        "caption": "Fig. 5. Musculoskeletal model of the lower limb: (a) muscle units; (b) mechanical model with generalized coordinates",
        "texts": [
            " In this section both approaches to solve the muscle force-sharing problem in biomechanics proposed are applied to the normal and to mechanically disturbed gaits measured in a gait analysis laboratory, Ackermann and Gros [2]. The extended inverse dynamics and the modified static optimization are compared to the static optimization. A 2-D mechanical model of the skeletal system of the right lower limb is adopted here, composed by three rigid bodies, the thigh, the shank and the foot. The motion is performed in the sagittal plane and is described by three generalized coordinates and two rheonomic constraints, refer to Fig. 5b. The generalized coordinates are the angle \u03b1 describing the rotation of the thigh, the angle \u03b2 describing the knee flexion, and the angle \u03b3 for the ankle plantar flexion. The two rheonomic constraints are the horizontal and vertical positions of the hip joint, xhip and zhip, respectively. The pelvis and trunk are assumed to remain in the vertical position throughout the gait cycle, what is reasonable for normal walking. The masses, center of mass locations, and the mass moment of inertia of the three segments in the sagittal plane are computed using the tables in de Leva [14] as functions of the subject\u2019s body mass, stature, thigh length and shank length. The motion and the ground reaction forces were measured in a gait analysis laboratory. The eight muscle groups considered in this analysis are shown in Fig. 5a. The Hill-type muscle model is composed by a contractile element CE and a series elastic element SE, while the force of the parallel elastic element PE is set to zero, Fig. 1. In this model all the structures in parallel to the CE and the SE are represented by total passive moments at the joints, which include the moments generated by all other passive structures crossing the joints, like ligaments, too. The formula for the passive moments at the hip, knee and ankle are functions of \u03b1, \u03b2 and \u03b3 as proposed by Riener and Edrich [22]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002426_mawe.201100898-Figure3-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002426_mawe.201100898-Figure3-1.png",
        "caption": "Fig. 3. Calculation of the (x, y)-coordinates of a point PN on a gear flank",
        "texts": [
            " The calculated gear geometry changes can then be compared with changes of the shape and size of the gear base body in order to determine their influence on the toothing deviations. The nominal position PN of any point on a flank of a gear can be determined in (x, y, z)-coordinates as a function of a few parameters (according to [7\u20139]): \u2013 rolling angle , \u2013 position in z-direction, \u2013 base circle radius rb, \u2013 rotational position of the gear q, \u2013 number of teeth n, \u2013 base helix angle bb The position PN of a point in 2D-coordinates is calculated by a vector addition, Figure 3: ~rN \u00bc~a\u00fe~b \u00f01\u00de In 2-D polar coordinates the vector~a has a length of rb and an angle of and the vector~b is described by its length of rb and an angle of =2. In Cartesian coordinates: ~a \u00bc rb cos sin ;~b \u00bc rb cos\u00f0 =2\u00de sin\u00f0 =2\u00de ~b \u00bc rb sin cos : \u00f02\u00de With Equation 1 follows: ~rN \u00bc rb cos \u00fe sin sin cos \u00f03\u00de Adding a factor s for the flank side (s = \u20131 for left side, s = +1 for right side), the rotational angle q of the involute origin and the axial position zN to Equation 3 leads to a generalised description (bb: base helix angle): ~rN \u00bc rb cos \u00fe sin rb sin cos zN 0 @ 1 A; \u00bc \u00fe s \u00fe zN tan b rb \u00f04\u00de The measured point PI is calculated by applying a vector addition to the corresponding point PN on the nominal gear, Figure 3: ~rI \u00bc~rN \u00fe~d \u00bc~rN \u00fe ~d ~ninv ~ninvj j \u00bc ~rN \u00fe d ~ninv ninv \u00f05\u00de For helical gears the z-coordinate cannot be omitted, as the normal vector ~ninv has a non-zero z-coordinate due to the base helix angle bb. The components of the normal vector~ninv are: ~ninv \u00bc @~r @ @~r @z \u00bc sin cos tan b 0 @ 1 A) ninv \u00bc ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 1\u00fe tan b\u00f0 \u00de2 q \u00f06\u00de The calculation of the normal vector is described in detail in [10]. Using the Equations 4 and 6 in Equation 5 results in: ~rI \u00bc xI yI zI 0 @ 1 A \u00bc rb cos \u00fe sin \u00fe d rb ninv sin rb sin cos d rb ninv cos zN \u00fe d ninv tan b 0 BBBBBB@ 1 CCCCCCA \u00bc rb cos \u00fe \u00fe d rb ninv sin rb sin \u00fe d rb ninv cos zN \u00fe d ninv tan b 0 BBBBBB@ 1 CCCCCCA \u00f07\u00de 122 i 2012 WILEY-VCH Verlag GmbH & Co"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000480_1452001.1452004-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000480_1452001.1452004-Figure2-1.png",
        "caption": "Fig. 2. The pursuer and evader game for target capture game.",
        "texts": [
            " If the ratio of the pursuer speed vp to the the evader speed ve, \u03b1, is larger than 1, then the min-max optimal strategy for the evader and pursuers is given by: \u03b8e(x0) = \u03b3 , \u03b8p(x0) = \u03b3 , (2) where \u03b3 = tan\u22121( yr xr ) and V (x0) = \u221a x2 r + y2 r (\u03b1\u22121)\u2217ve . Equivalently, the pursuer moves toward the evader directly until catching the evader, while, the evader moves in the same direction to prolong the catching time. PROOF. Given the current location of the evader and pursuer, the set of points that the evader can reach before the pursuer is given by the well known Appolonius circle. The min-max optimal strategies for both the pursuer and the evader are to go directly to the boundary point. As shown in Figure 2, the current pursuer location is B and the current evader location is A. For any time interval dt, the maximum distance that the pursuer and evader can move are vp \u2217 dt and ve \u2217 dt. All the possible locations are on the circle around the current pursuer and evader locations. Point B\u2032 is the crosspoint of the circle around the pursuer and line BA. Point A\u2032 is crosspoint of the circle around the evader and the other side of line BA. We claim the point B\u2032 and A\u2032 are the min-max optimal strategy pair for pursuerevader movement during time dt",
            " The evader does not deviate from its min-max equilibrium strategy if and only if sampling period Tsample with respect to the distance d pe between the pursuer and evader satisfies: Tsamp(d pe) < d pe vp . (3) ACM Transactions on Autonomous and Adaptive Systems, Vol. 3, No. 4, Article 14, Publication date: November 2008. In other words, the sampling period should decrease proportionally with decreasing distance between evader and pursuer to guarantee that the evader does not have an incentive to deviate from its strategy. PROOF. Assume the pursuer moves first. For any time interval Tsample, (vp \u2217 Tsample < \u221a (xr )2 + ( yr )2), as shown in Figure 2, the pursuer will move to B\u2032, where B\u2032 is the crosspoint of the circle around the pursuer and line BA. The evader can move to any location in the circle that is centered at A and has the radius ve \u2217 dt. It will choose the location A1 that has maximum B\u2032 A1. By triangle inequality, A\u2032 is the location that maximize the B\u2032 A1. However, if vp \u2217 Tsample \u2265 d pe, the evader can find a better location such that B\u2032 A\u2217 > B\u2032 A\u2032 as shown in Figure 3. 2.3.2 Effect of Message Losses. From the previous sampling rate analysis, to guarantee the optimum evader capture, the information must be updated before the pursuer reaches the previous evader location"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000531_pct.2007.4538314-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000531_pct.2007.4538314-Figure1-1.png",
        "caption": "Fig. 1. CP- curves for different pitch angles 0 4 8 12 16 20",
        "texts": [
            " WIND TURBINE MODELING Mathematical relation for the mechanical power extraction from the wind can be expressed as follows [16]: Where, Pw is the extracted power from the wind, is the air density [kg/m3], R is blade radius [m], Vw is the wind speed [m/s] and Cp is the power coefficient which is a function of both tip speed ratio, , and blade pitch angle, [deg]. In this work, the Cp equation as shown below has been taken from [17]. Where, B is rotational speed [rad/s]. The Cp- curves are shown in Fig. 1 for different values of . In this study, the six-mass drive train model of WTGS shown in Fig. 2 is used. The parameters of the six-mass drive train model are shown in the appendix. The detailed description of the six-mass drive train model is available in [7-8]. Fig. 3 shows a model system used for the simulation analyses. One synchronous generator (SG) and one induction generator (IG) are connected to an infinite bus through transformers and transmission lines respectively. A capacitor bank, C, has been used for reactive power compensation of IG at steady state"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002319_j.jsv.2012.05.009-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002319_j.jsv.2012.05.009-Figure2-1.png",
        "caption": "Fig. 2. Sketch of three points on the free end for describing the flapping motion of the elastic plate.",
        "texts": [
            " On the spanwise boundaries, we consider two types of conditions: One is to specify the uniform velocity, which is referred to as BBC (Bounded Boundary Condition), and the other is to impose the periodic boundary condition, referred to as PBC (Periodic Boundary Condition). The reason why we specify the uniform velocity rather than the zero velocity on the side walls in case of BBC is that we want to reduce the effect of the wall constraint. In order to present a quantitative description of the flapping motion of the plate, the y-coordinates of three points on the plate free end are tracked: one central point defined by yc,and two side points defined by y1 and y2, as shown in Fig. 2. In addition, we define the average y-coordinate of all discrete points on the free end by ya. The flapping amplitudes for yc, y1, y2 and ya are denoted by Ac, A1, A2 and Aa, respectively. The length, width and thickness of the plate are L, W and h, respectively. The reference (or equilibrium) configuration of the plate (X, Y, Z) is aligned with the streamwise direction, and the initial configuration (x0, y0, z0) is declined with the streamwise direction by an angle tan 1d0, i.e.: x0 \u00bc X; y0 \u00bc d0X\u00feY ; z0 \u00bc Z The intensity of the initial disturbance to the plate d0 is set to be 0"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003252_j.triboint.2012.07.002-Figure9-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003252_j.triboint.2012.07.002-Figure9-1.png",
        "caption": "Fig. 9. Finger/ball interface at (a) low load (0",
        "texts": [
            " 2), the wet condition was too moist for the device to give a reading. The device was not capable of measuring the moisture on the mitts, so this was not monitored. Fig. 8 shows force data from tests carried out in dry conditions with a finger pad (note that errors for all friction forces presented in this section are 70.4%). For normal forces up to 7\u20138 N, the friction forces for the more densely pimpled Ball 1 surface are higher indicating better grip. Above this level of force the data merges indicating all three surfaces give similar friction. Fig. 9, shows schematically a finger pad against Ball 1 and Ball 2 surfaces. At low force (Fig. 9a), an adhesion mechanism (where contact area is the main influence) would be expected to dominate. With more dense pimples the contact area would clearly be larger, hence the larger friction force. However, as the normal force increases (see Fig. 9b), the finger would be expected to deform, hence deformation, which would be higher with more dispersed al pattern silicone) Silicone round pimple pattern (white = silicone) itt materials tested. pimples, would contribute to a greater degree. For the results seen here a balance has been reached between the two cases so the deformation and increased contact area for Balls 2 and 4 at higher load has increased the friction force to (or slightly above perhaps) the level of Ball 1, which would have higher values of both area and deformation, but to a lesser degree than Balls 2 and 4"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003083_iecon.2011.6119758-Figure3-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003083_iecon.2011.6119758-Figure3-1.png",
        "caption": "Fig. 3. Vision subsystem.",
        "texts": [
            " As mentioned in Section I, although in commercial systems the control of the cladding process is usually based on temperature measurements, the use of CCD or CMOS cameras is becoming increasingly popular. The main problem associated to the use of cameras as sensing elements comes from the fact that, in current industrial systems, the control algorithm is implemented in a PC. Because of its slow response, only low-resolution images can be used, resulting in low accuracy of the measurements, which negatively impacts the performance of the control system. The cladding system in the Laser Application Center includes a vision subsystem (Fig. 3) in a coaxial arrangement with the laser beam. It consists of a CMOS camera with 8-bit (256 grayscale levels) CameraLink interface and dynamic range up to 120dB, and auxiliary optical components: mirrors, a lateral augmentation telescopic system, and two filters, one to prevent the laser beam to damage the camera, and another to adjust the saturation of the camera by the light generated in the fusing process. In this work, the configurable resolution of the camera has been set to 800x600 pixels"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002148_j.otsr.2009.11.005-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002148_j.otsr.2009.11.005-Figure2-1.png",
        "caption": "Figure 2 Illustration of the custom-made jig for bending testing including the specially constructed deflection sensor.",
        "texts": [
            "9% NaCl solution during the entire test period. he embedded specimens were mounted in specially contructed jigs for compressive, 4-point-bending, or torsional esting in a Material Testing System (MTS [Model 858, MTS orp., Minneapolis, USA]). The order of stiffness testing as randomised and measurements were performed on each pecimen by using compressive, bending (anteroposterior nd mediolateral) and torsional load. The resulting deforation was detected by custom-made compression, torsion Fig. 1) and deflection (Fig. 2) sensors (LVTD and precision otentiometer). For each stiffness testing-procedure, a preonditioning of 10 cycles was conducted before the actual esting in order to assure repeatability. The callus tissue ithin the specimens was loaded during the different types f testing up to 15 Nm for torsional, to 750 N for compresive and to 6.5 Nm for bending load. During testing, load nd deformation were simultaneously recorded in order to etermine stiffness, which is defined as the slope of the Stiffness of callus tissue during distraction osteogenesis 157 F d i l s t b S S t ( t o c s a s t w r l a S d t c o a o R T c 5 igure 4 Comparison of regained stiffness characteristics of istracted tibiae at the 74th postoperative day in relation to ntact tibiae"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002241_978-1-4471-4141-9_70-Figure70.3-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002241_978-1-4471-4141-9_70-Figure70.3-1.png",
        "caption": "Fig. 70.3 Top view of GLSBL in deployed configuration. a Top view of deployed configuration, b bundle compact form",
        "texts": [
            " The GLSBL can also be designed to be deployable onto such \u2018\u2018Y\u2019\u2019 shape deployed profile by using the joint axis position determination method as presented in [4]. Suppose that ABC is the required triangular deployed profile, for this case, however, the three angles \\AOB; \\BOC and \\COA are no longer identical because the D\u2013H lengths of the six links are not identical, so that the mechanism can be designed with non-equilateral triangular profiles, but the physical link of the six links for GLSBL can be designed with identical lengths. This is an important property for the construction of tripod deployable module. As shown in Fig. 70.3, AO; BO and CO are the three joint axes of GLSBL connecting to the adjacent two parallel links. LetAG \u00bc BH \u00bc CI, i.e., all the lengths of the physical links are identical. DG; GF; FI; IE; EH; HD are the 1z 2z 3z 4z 6z 5z 1X 2X 3X 5X 4X 6X 1 2 3 4 5 6 Fig. 70.1 A general D\u2013H model of 6R linkage widths of the six physical links, DG \u00bc HD; GF \u00bc FI; IE \u00bc EH. As all the links are designed with rectangular shape in the top view, then AG?DF; BH?DE; CI?EF and extension lines of AG; BH; CI are concurrent at the point O so that the point O is the center of the inscribed circle of DDEF"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002273_1.3680609-Figure12-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002273_1.3680609-Figure12-1.png",
        "caption": "Fig. 12. A negative Magnus force can arise as shown here if the air flow is laminar on the upper side of the ball and turbulent on the lower side. In this example, the peripheral speed of the ball due to spin is 4.4 m/s, the center of mass speed is 10 m/s, point A translates to the right at 5.6 m/s and point B translates at 14.4 m/s. The air flow near A is laminar, and the flow near B is turbulent.",
        "texts": [
            " A reversal in the direction of the Magnus force has previously been observed in wind tunnel experiments and can be attributed to the fact that the boundary layer can become turbulent on one side of the ball and remain laminar on the opposite side.1 For example, consider the case shown in Fig. 8 where the ball was spinning at 367 rpm with a peripheral speed Rx \u00bc 4:4 m/s, and translating at v\u00bc 10 m/s. The relative speed of the ball and the air was 14.4 m/s on one side of the ball and 5.6 m/s on the opposite side of the ball, as indicated in Fig. 12. The local Reynolds number is 2:2 105 on the high-speed side and 8:5 104 on the low-speed side. A turbulent boundary layer on the high speed-side will separate later than a laminar layer on the low-speed side, deflecting air toward the low-speed side. The air exerts an equal and opposite force on the ball in a direction from the low-speed to the high-speed side, in the opposite direction to the conventional Magnus force. At ball speeds less than about 9 m/s, the ball spin was typically about 100\u2013200 rpm and the Reynolds number was not high enough for the boundary layer to become turbulent"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003226_ac60252a021-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003226_ac60252a021-Figure1-1.png",
        "caption": "Figure 1. Electrollysis cell showing optical alignment",
        "texts": [
            " Any light source so used must be sufficiently intense on emerging to be able to stimulate a photomultiplier; it must be exceedingly parallel to make enough minute reflections throughout its entire track on the surface; it must be very narrow so that a significant fractioa of its cross-sectional area is constantly within about lO4A 896 ANALYTICAL CHEMISTRY of the surface; and finally, it is rather advantageous to have a fully polarizecl light source because by maintaining the plane of polarization parallel to the reflecting surface, interference effects are diminished and the reflectivity is improved. Electrolysis Equipment. The radial symmetry of the electrolysis cell and the smooth surfaces of both electrodes ensured a uniform electric field over the entire silver surface. In most experiments the electrolyte was 0.25M Na2S04, made from analar grade salt and triply distilled water. The solution was usually deoxygenated prior to admission to the cell by bubbling Na gas through the solution and flushing out the cell. As shown 111 Figure 1 the Plexiglas cell included filling the emptying ports so that solutions would be interchanged in situ without disturbing the optical alignment. Both ac and dc electrical power for the electrolysis was used. Alternating voltages were provided by an audio oscillator (Hewlett-Packard, Model 200J) capable of supplying frequencies from 6 to 6300 Hz, whereas 6-volt dry cells were Figure 2. Sketch of apparatus L. C. R. P. S. L.I.A. 0. A . B. D. Laser Electrolysis cell Rotating-sector disc Photomultiplier Photomultiplier power supply Audio oscillator Neon lamp Photocell Millivolt recorder Lock-In Amplifier n used as dc supplies"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003389_1.3649934-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003389_1.3649934-Figure2-1.png",
        "caption": "FIGURE 2. Camber angle and slope angle.",
        "texts": [
            " non-null value of compliance may be used. The same may be done for all other joints. The road-tyre contact forces are computed according the well known Magic . T1 T5 T2 T9 UX UY Z T10 T7T8 FIGURE 1. Motorcycle Model. Formula for motorcycle, and two approaches are available for the force coupling: the Similarity Method and the Loss Functions Method, [23]. The tyre forces are applied at the actual contact point whose position on the carcass is defined by means of the tyre camber and tyre slope , see FIGURE 2. Also, the carcass compliance and damping are accounted for by means of the lateral L, radial R, and tangential deflections. This approach automatically includes the tyre lag, i.e. no additional relaxation equations are necessary, see [23,24]. Summarizing, the tyre model has five additional variables =( , , L, R ) and as many equations ( , , , , ) p v w 0 (3) which are solved together with the differential equations of motion Eq. (1) and algebraic constraints in Eq. (2). 3. COMPARISON WITH ROAD TESTS In this section the road tests are compared with numerical results to prove the reliability of the multi-body model used in the paper"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003159_978-1-4419-9985-6-Figure1.9-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003159_978-1-4419-9985-6-Figure1.9-1.png",
        "caption": "Fig. 1.9 Lower figure shows a schematic top view of the sensor. Placement, size and number of the piezoresistors not as in the real sensor, the simplification used for the clarity of the figures. Upper figure shows the cross section at A-A",
        "texts": [
            "Passivisation of the metallization with PECVD2 silicon nitride. Generation of glue areas, dicing marks and light shield. 8.Lithography and opening of the pads and the DRIE areas by dry etching, 9. Lithography using a thick photo resist (25 \u00b5m), dry etching of the oxide passivation and DRIE of the Si substrate 10. Resist strip and cleaning tors with LPCVD1 silicon oxide, Fig. 1.8 Schematical description of the process flow for the fabrication of the force sensors. The cross-sections correspond to the top view shown in Fig. 1.9 2. When time for mounting the sensor element is lifted from the elastic foil with the pick- and place tool and placed precisely into a fabricated mechanical pick-up. 3. Stud bumps are placed on each gluing area of the sensor element to define the distance between the sensor element and the transducer. 4. Glue is dispensed onto all five gluing areas on the sensor element. 5. The transducer, which has previously been cleaned, is placed onto the sensor and stays there until the glue is fixed and the mounting is ready"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002788_vppc.2011.6043245-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002788_vppc.2011.6043245-Figure1-1.png",
        "caption": "Fig. 1. \u03b4-\u03b3 and d-q reference frames.",
        "texts": [
            " Kf=p\u03a8f where \u03a8f is the rotor permanent magnet flux linkage through the stator windings, and p is the number of pole pairs. The mechanical equation is: ( ) ( )\u03a9\u0393\u2212\u0393=\u03a9 Lqm i dt dJ (2) in which J is the moment of inertia and : ( ) qfq iKi =\u0393m (3) is the torque generated by the motor, and: ( ) ( ) +\u03a9+\u03a9+\u03a9=\u03a9\u0393 2 210 .b.bsgn.bL (4) represents the unknown load torque expression. We assumed that 0>\u03a9\u0393 dd L for all practical speeds. Consider now the case where the rotor position is unknown, i.e. the actual d-q reference frame cannot be localized. In this case, we use another reference frame, called \u03b4\u2212\u03b3 (see Fig. 1), in which the electrical equations can be expressed as follows: \u03b3\u03b3\u03b4\u03b3\u03b3 \u03b4\u03b4\u03b3\u03b4\u03b4 veipLiRi dt dL veipLiRi dt dL ssss ssss +\u2212\u03a9\u2212\u2212= +\u2212\u03a9+\u2212= (5) where v\u03b4, v\u03b3, i\u03b4 and i\u03b3 are the \u03b4\u2212\u03b3 components of the stator voltage and current vectors, and ps \u03d1=\u03a9 is the mechanical angular speed of the \u03b4\u2212\u03b3 frame. e\u03b4 and e\u03b3 are the stator backEMF components on \u03b4\u2212\u03b3 frame defined by: \u03d5= \u03d5= \u03b3 \u03b4 cos.ee sin.ee (6) where \u03d5 = \u03d1 \u2013 \u03b8 and e = Kf \u03a9. Then, the mechanical dynamic can be expressed as: ( ) ( )\u03a9\u0393\u2212\u03d5+\u03d5=\u03a9 \u03b3\u03b4 Lf cosisiniK dt d J (7) In the classical sensorless control of PMSM, \u03d1 is the estimated rotor position calculated by: \u222b \u03d1+\u03a9=\u03d1 0dt"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003509_j.proeng.2013.08.203-Figure8-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003509_j.proeng.2013.08.203-Figure8-1.png",
        "caption": "Fig. 8. not optimized posture (a) and (c), optimized posture (b) and (d) while realizing a square defined by 8 points (corners and midpoints)",
        "texts": [
            " One of the difficulties of aggregation methods is the choice of the weighting assigned to each criterion. To change the relative importance of each criterion according to the need, variable weightings are introduced (Lee and Buss (2006)). The form of the objective function becomes: with (17) with )(qw ii the weighting function which depends on the criteria i , which goes from 0 to 1. Models and resolution methods are computed under Matlab\u00ae. The proposed method allows the optimization of the behavior of the robotic cell within the proposed criteria in speed, torque and stiffness. Fig. 8 represents the posture of the robot during the not optimized and respectively optimized path (without and with the use of the turntable and the functional redundancy). Rvs represents the speed ratio (Eq. 7) for IRB6660 robot while Rvt is the tricept one. First, the robotic cell with IRB6660 robot is up to twice faster than the one with the Tricept robot (Fig. 9) by comparing Rvs and Rvt. This method permits to have minimal expectances concerning the different ratio. A higher weighting for the capacity in speed is given with a minimum expectances of 12 for the ratio"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003520_j.euromechsol.2011.11.003-Figure3-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003520_j.euromechsol.2011.11.003-Figure3-1.png",
        "caption": "Fig. 3. T2R1-type parallel manipulator with decoupled and bifurcated planar-spatial motion of the moving platform: constraint singularity (a), branch with planar motion (b) and branch with spatial motion (c); limb topology PkRkRkRtRPtPttRtR-PkRkRS.",
        "texts": [],
        "surrounding_texts": [
            "The term of constraint singularity (CS) has been recently coined (Zlatanov et al., 2002) to characterize the configuration of lower mobility parallel manipulators in which both the connectivity of the moving platform and the mobility of the parallel mechanism increase their instantaneous values. From a constraint singularity, the mechanism can get out with or without branching (Gogu, 2008b). When branching occurs in a constraint singularity, the mechanism can reach different configurations, called branches, and have different independent motions of the moving platform. In this case, the constraint singularity is also called branching or bifurcation singularity. A branch refers to the free-of-singularity configurations of the mechanism inwhich each structural parameter keeps its value. For this reason, this value is called global of full-cycle value for a branch. Two types of branching in constraint singularity (BCS) have been defined in Gogu (2008b). Branching of type BCS1 occurs when a parallel mechanism F ) G1 . Gj . Gk gets out from a constraint singularity in different branches characterized by the same mobility and the same connectivity of the moving platform but with different bases of the vector space of relative velocities between the moving and fixed platforms. Branching of type BCS2 occurs when a parallel mechanism F ) G1 . Gj . Gk gets out from a constraint singularity in different branches characterized by distinct values of mobility and connectivity of the moving platform. The parallel manipulators presented in this paper have constraint singularities with branching of type BCS1. Branching of type BCS2 occurs in kinematotropic mechanisms (Gogu, 2009c). The finite displacements and the velocities in the actuated joints are denoted by qi and _qi, the linear velocities of the characteristic point H of the moving platform, by v1 \u00bc _x; v2 \u00bc _y and the angular velocity of the moving platform by u \u00bc ua \u00bc _a or u \u00bc ud \u00bc _d: In both branches, the moving platform undergoes two planar translations and one rotation but the rotation axis is different in the two branches. In the first case, the rotation velocity u \u00bc ua \u00bc _a is parallel to the x-axis and the moving platform undergoes a spatial motion. In the second case, the rotation velocity u \u00bc ud \u00bc _d is parallel to the z-axis and the moving platform undergoes a planar motion. With these notations, the linear mapping between the actuated joint velocity space and the end-effector velocity space for a branch the T2R1-type PMs is defined by: 2 4 v1 v2 u 3 5 \u00bc \u00bdJ 2 4 _q1 _q2 _q3 3 5 (8) where J is the Jacobian matrix. In the constraint singularity, the T2R1-type parallel manipulators with bifurcated motions has instantaneously iM \u00bc iSF \u00bc 4 and (iRF)\u00bc(v1, v2, ua, ud). The bifurcation in this constraint singularity can be used to change motion type of the moving platform. In the two distinct branches, the parallel mechanism is characterized by M\u00bc SF \u00bc 3 and (RF)\u00bc(v1, v2,ua) or (RF)\u00bc(v1, v2,ud). In the constraint singularity, both connectivity of moving platform and mobility of parallel manipulator increase their instantaneous values. More details about the behaviour of parallel manipulators in a constraint singularity and the associated structural parameters can be found in Gogu (2009c)."
        ]
    },
    {
        "image_filename": "designv11_12_0003137_j.mechatronics.2010.12.002-Figure7-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003137_j.mechatronics.2010.12.002-Figure7-1.png",
        "caption": "Fig. 7. The comparison of the trajectory fo",
        "texts": [
            " For the proposed method, the maximum deviation from the desired trajectory is about 15 m and the average lateral tracking error is 8 m, which is only half of the traditional PID control method. In order to illustrate the system trajectory following ability, the SUAV was commanded to fly a series of concentric orbits and a rectangle region. Based on the research task, the SUAV cruised a rectangle region with 1000 m 1000 m length and circular region with radius of 400 m. The largest wind disturbance was 10 m/s, which corresponds to approximately 40% of the SUAV airspeed. The results are shown in Fig. 7. The position of the SUAV was measured using the on-board DGPS unit. The bias error associated with the DGPS measurement is approximately 0.3 m and is virtually unchanging during the flight experiments. The trajectory following error is calculated by computing the lateral distance between the desired trajectory and the location of the SUAV. The comparison of the tradition PID and the proposed method is shown in Fig. 7. The solid line, dot line and dash-dot line are the predefined flight line, the flight trajectories generated by the proposed method, and PID method, respectively. Since the wind disturbance varies quickly, the mean result of the repeat flight tests are chosen to show the control performance. Through the proposed control method, system adjusts aileron quickly to realize trajectory following control. Due to wind disturbance, the maximum deviation from the desired trajectory is about 30 m and the mean error is 10 m"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001313_09544100jaero155-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001313_09544100jaero155-Figure1-1.png",
        "caption": "Fig. 1 The twin rotor MIMO system",
        "texts": [
            " The TRMS is a laboratory platform designed for control experiments by Feedback Instruments Ltd [18]. In certain aspects, its behaviour resembles that of a helicopter. For example, it possesses a strong cross-couplings between the collective (main) and tail rotors. The TRMS is characterized by its complex and highly non-linear dynamics. Some of its states and outputs are also inaccessible for measurements. All these typify TRMS as a challenging engineering problem. The control objective is to make the beam of the TRMS tracks a predetermine trajectory. Figure 1 shows the TRMS considered in this investigation. The dynamic model as supplied by the manufacturer has been improved in this study and the DC motors are simulated with respect to the corresponding equations. The TRMS possesses two permanent magnet DC motors; one for the main and the other for the tail propelling. The motors are identical with different mechanical loads. The mathematical model of the main motor, as shown in Fig. 2, is presented in equations (1) to (5). The mathematical model of the remaining parts of the system in vertical plane is described in equations (6) to (8) (see Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002712_0954405412461865-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002712_0954405412461865-Figure1-1.png",
        "caption": "Figure 1. The benchmarking part and inspected features.",
        "texts": [
            " Analysis procedures for various feature characteristics can be defined, including for example flatness surfaces, roundness of holes and cylinders, thickness of walls and other feature aspects. Based on an extensive literature study of previous research and in-house experience, a benchmarking procedure was developed aimed to evaluate the performance capabilities regarding rigorous geometric shape and dimensioning accuracy. In this way it is possible to create functional and non-functional features of real parts and also free-form surfaces, fillets and draft angles can be tested. The developed part and inspected features are showed in the Figure 1. Thirty-eight geometric shapes, 22 dimensional elements were inspected, plus 12 points of the freeform surface. The geometric inspection includes the most relevant features that can be found in parts manufactured with RP processes.2,7,31 For dimensional accuracy, a range of fine, small, medium and large distances, radius and diameters, and taper amplitude were considered. The diagrams of the controlled elements are showed in Figure 2. First, the experimental study was conducted to demonstrate the use of the benchmark for the performance evaluation of the TDP (ZPrinter) process"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002573_s11249-011-9820-8-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002573_s11249-011-9820-8-Figure1-1.png",
        "caption": "Fig. 1 Schematic diagram of the slide/roll rig for the measurement of film thickness by optical interferometry",
        "texts": [
            " Despite these findings, the combined means of the fluorescence technique and the optical interferometry have been rarely used in lubrication studies up to the present. In this study, the optical interferometer and the fluorescent microscope are used to observe the dynamic process of a typical oil lubricant in the contact region at high slide/roll ratios of [1.9. The reasons for the lubrication failure under these conditions are discussed in detail. 2 Experimental Section 2.1 Test Rig The optical test rig used in the experiments is illustrated in Fig. 1. A highly polished steel ball with a diameter of inch was loaded against the lower surface of a glass disc coated with a thin semi-reflective chromium layer. The load W of 5.2 N was applied by the weight cells, and the corresponding maximal contact pressure and Hertzian radius a were 0.44 GPa and 0.085 mm, respectively. The tension\u2013compression sensor was added to the outer shaft connected the steel ball and used to measure the traction force FT between the ball and the disc. Hence, the traction coefficient TC was defined as TC \u00bc FT W \u00f01\u00de The ball and the disc were driven by two independent motors, ensuring that the slide/roll tests could be well controlled",
            " In order to obtain clear interferometric images, we accelerated the speeds of the ball and the disc to their pre-specified values allowing them to touch each other. In the slide/roll tests, the interference orders of interferometric images were determined from the pure rolling test under the same testing conditions. Detailed descriptions of the test rig in pure rolling were obtained from the literature [13\u201317], and the entrainment speed ranged from 0 to 1.360 m s-1. 2.2 Fluorescent Rig The light microscope could be switched to the fluorescent mode (model SZX16; Olympus, Tokyo, Japan; Fig. 1). The excitation was provided using a diode-pumped pulsed laser generated by a long-life metal halide light bulb of the X-Cite 120Q system (EXFO, Quebec, QC, Canada). The new test rig was used to observe the fluorescent images in the following tests. A magnification of 29 was found to be convenient for observing the lubrication region of both the Hertzian contact and the inlet/outlet. These sequencer files were recorded via an Evolution QEi cooled charge coupled device camera (Qimaging, Surrey, BC, Canada)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000901_tpwrd.2007.905551-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000901_tpwrd.2007.905551-Figure1-1.png",
        "caption": "Fig. 1. Corona pulse test configuration (1) transformer; (2) insulator string; (3) high voltage line; (4) tower; (5) ground line; (6) Rogowski coil; and (7) oscilloscope.",
        "texts": [
            " If there are some faulty insulators in one string, the voltage across normal insulators will increase because the faulty insulator hardly undertakes voltage. Hence, corona discharges from normal insulators will appear or be enhanced [14]. To prove this fact, some experiments were carried out in laboratory. In china, on a 110 kV transmission line with its phase-toground voltage being 66 kV (considering a 5% voltage improvement) seven porcelain insulators are often used. In laboratory, an insulator string consists of seven XP-7 porcelain insulators was tested with 66 kV ac voltage, the test configuration is shown in Fig. 1. The ground line goes through a Rogowski coil with 3-dB cutoff frequency from 200 kHz to 3 MHz and sensitivity of 20 mV/mA to detect the corona pulse current. The pulse current signals are sampled by a DL1540L digital oscilloscope (200 MSamples/s sample rate). When every insulator in a string is good, few corona pulses are detected. If there is one faulty insulator (25 ) in a string, some corona pulses can be found. If there are two faulty insulators (25 and 5 , respectively) in a string, many corona pulses can be detected [15]",
            " In order to know the background noise, the signals in the ground line on several 110 kV transmission lines were measured with current sensor developed by us. It was found that there are some narrow-frequency band interferences in the ground line, such as insulator leakage current, harmonic current and carrier current. These interferences usually have very high amplitude and they can overlap corona currents. At laboratory, the corona pulses from many insulator strings were measured by using a circuit resembling to Fig. 1, in which a resistance of 50 was connected in series with the ground line and the voltage across this resistance was recorded by the oscilloscope. The typical waveform of the corona signal from the insulator string and its frequency spectrum are shown in Fig. 3. These corona pulses have a rise-time about 50 ns and a pulse width between 100 ns to 400 ns. The main energy of corona pulses distribute in 20 MHz [14]. The most obvious difference between narrow-frequency band noise and insulator corona signal is frequency band"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003713_978-1-4614-2419-2-Figure52.2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003713_978-1-4614-2419-2-Figure52.2-1.png",
        "caption": "Fig. 52.2 (a) Load cell; (b) Connection between the load cell and the hand-grip; (c) Measured force",
        "texts": [
            " Forces applied by different athletes during the competition were analysed in order to understand the different climbing techniques and to correlate the measures with the performance achieved by the athletes. The final goal of the designed dynamometric hold is to measure the force applied by the athlete to the hold during a world cup speed climbing competition. it does not have to modify shape and fixing of standard grips. To fulfill this requirement, circular mono-axial load cells (HBM KMR100) have been mounted behind the climbing wall as shown in Fig. 52.2, in order to be \u201cinvisible\u201d from the athletes point of view, and to be in agreement with the regulation (guaranteeing safety of athletes). The hand-grip is fixed to the wall through an M12 bolt. Between the wall and the nut, KMR load cell, packed with two washers, has been preloaded with a 87,3 Nm tightening torque, which corresponds to an axial preload of 37.7 kN, as specified by bolt DIN standards. Moreover a 0.1 mm thick washer (maximum allowed by rules) has been put between the wall and hand-grip in order to reduce friction with the wall surface, preventing the possibility of correctly measure forces applied by athletes. The reaction of the fixing system consequentially to the athlete\u2019s action on the other side of the wall is measured by the load cell along the direction of the bolt. The measured force, Fa, is thus the projection of the force applied by the athlete, F, along the axis of the bolt connecting the hold with the wall and so it accounts both for the vertical (Fv) and horizontal (Fh) force components according to: Fa \u00bc Fvsina\u00fe Fhcosa (52.1) being \u03b1 the overhang angle of the wall (Fig. 52.2c). four instrumented holds. It must be pointed out that the measurement devices were out from athletes view and athletes were not aware of their presence. The acquisition system is composed of a six channels HBM MGCplus conditioning unit and a laptop, which provide the data logging and the real time data elaboration and visualization of the test results (Fig. 52.4). The start buzzer was also recorded to synchronize different climbs in post-processing phase. Every instrumented hold has been verified by applying a known vertical force Fv on the grip and measuring the correspondent projection along the axis of the load cell Fa"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001411_s0022-0728(75)80292-0-Figure5-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001411_s0022-0728(75)80292-0-Figure5-1.png",
        "caption": "Fig. 5. Correlation of E~ for the redox couples RCo(DH)2(H20)+/RCo(DH)2(H2 O) in 1.0 M HC104, at 25.0\u00b0C, with pK,(RH), for the organic axial groups R: (1) C6HsCH2; (2) C6H5; (3) CH3;",
        "texts": [
            " The c-C6HllCH; deviates slightly from the above correlation. Again, a small steric effect may be operative. However, the benzyl group does not fit the correlation, and the deviation presented can not be attributed solely to steric effects (@ c-C6H 11CH 2 and i-C4H9). CYCLIC VOLTAMMETRY OF ORGANOCOBALOXIMES 203 (4) i-C4Hg; (5) n-CaHT; (6) C2H5; (7) i-C3HT; and (8) c-C6Iq,1. Except for the benzyl derivative, a linear free energy relationship can be found between the E~ and the pK,(RH) values of the hydrocarbon acid 14 (Fig. 5) corresponding to the organic group in axial position, indicating that the effect might be due to the charge donation from the organic group, decreasing the ionization potential of the particle. It is likely that the pKa(RH ) values also contain the contributions of steric effects. This type of linear correlation has also been found by Costa et al. 15 in the reductions of some organometallic cobalt chelates. A linear free energy plot with Taft's polar substituent constants o-* can be expected only if steric and resonance interactions between the substituents and the reaction site are constant for a series of substituents. Since branching at the fl-carbon of the substituent does not cause significant deviation from the correlation in Fig. 4, the behavior of the benzyl group is apparently due to direct resonance interaction with the reaction site. Similarly, in Fig. 5 an excellent correlation is observed by those groups which can interact via an inductive mode; strong deviation again implies a resonance interaction. Finally, when taken together, the correlations (Figs. 4 and 5) strongly indicate that (1) the transition state species is electron-deficient, and (2) the Co-alkyl bonding electrons are involved in the electron transfer. The E~ obtained and the rate constants for the dissociation of the oxidized species, explain the large range of limiting-rate constants for the chemical oxidation of these compounds by IrC12- (ca"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001535_s002565441001005x-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001535_s002565441001005x-Figure1-1.png",
        "caption": "Fig. 1.",
        "texts": [
            " In this case, the cylinder motion along the curve is a no-slip rolling, and the cylinder itself is assumed to be homogeneous in the present paper. The no-slip rolling is related to technical requirements of vibroprotection problems. 2. STATEMENT OF THE PROBLEM AND CONSTRUCTION OF THE OBJECTIVE FUNCTIONAL We assume that a homogeneous cylinder of mass m and radius r is in a homogeneous field of gravity and begins its motion from point O without initial velocity rolling without slip along a certain cylindrical cavity with directional curve OKL to point L (see Fig. 1). In this case, curve OKL lies in the vertical plane. The cylinder center is at point M . The problem is to find curve OKL such that moving along this curve the cylinder rolls from point O to point L in the least possible time. To determine the shape of curve OKL, we introduce the rectilinear Cartesian coordinate system OXZ with origin at point O and direct the axis OX horizontally to the right and the axis OZ vertically downwards. We assume that the equation of the desired curve OKL is described by the function z=z(x)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002273_1.3680609-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002273_1.3680609-Figure1-1.png",
        "caption": "Fig. 1. (a) Conventional swing of a new cricket ball results from the asymmetric air flow around the ball. The stitching is inclined at about 20 to the ball path and is maintained in that orientation by rotation of the ball about an axis perpendicular to the stitching. Black dots denote the boundary layer separation points. (b) Reverse swing of a new ball occurs at high ball speeds due to asymmetrical separation of the turbulent boundary layers on each side of the ball.",
        "texts": [
            " Regardless of the source of the asymmetry, if the air is deflected downwards by a ball in flight then the air exerts an equal and opposite force upwards on the ball. Similarly, if the air is deflected to the left by the ball, then the air exerts an equal and opposite force to the right on the ball. Deflection of the air flow is caused by early separation on one side of the ball and late separation on the other side. To illustrate how a side force can arise in practice, we will consider the case of a new cricket ball with a raised seam, as shown in Fig. 1. Separation is a boundary layer effect whereby air flowing in a thin layer adjacent to the ball surface is slowed by friction until it comes to rest at the separation point. Within the boundary layer, air flows from the front of the ball toward the rear. Air remains at rest right at the ball surface itself, increases in speed in a direction perpendicular to the surface, and decreases in speed in a direction along the surface. At the separation point, t\u00bc 0 and @v=@y \u00bc 0, where v is the air speed along the surface, and y is the coordinate perpendicular to the surface",
            " In that case, the separation point for a ball traveling horizontally through the air, when viewed side-on, is near the top and bottom of the ball or shifted slightly toward the front of the ball. If one side of a ball is rough or has a raised seam, then the air flow in the boundary layer on the rough or the seam side will become turbulent and separate from the ball further toward the rear of the ball. Turbulent air in the boundary layer mixes with higher speed air at the outer edge of the boundary layer, thereby increasing the average air speed near the ball surface and delaying separation. An example of this effect is shown in Fig. 1(a). The net transverse flow of air in Fig. 1(a) is upward in the figure (actually to the left side of the ball, Fig. 1 being a bird\u2019s-eye view) since air separates later on the right side of the ball than the left side. Since the ball acts to deflect air to the left, the air exerts an equal and opposite force on the ball to the right. At high ball speeds, air in the boundary layer can become turbulent even if the ball surface is smooth. In that case, air flows in turbulent boundary layers on both sides of the ball regardless of whether one side is rough or contains a raised seam. Delayed separation on both sides of a ball acts to reduce the drag coefficient, resulting in a so-called drag crisis.3 In the case of a high speed cricket ball with a raised seam, the air flow remains asymmetrical despite being turbulent on both sides of the ball. If turbulent air encounters a raised seam, then the boundary layer is thickened and weakened18 in which case there is only a slight delay or no delay at all in the separation point, as indicated in Fig. 1(b). The latter effect is responsible for the reverse swing of a new ball observed at high ball speeds in cricket. 290 Am. J. Phys., Vol. 80, No. 4, April 2012 Rod Cross 290 Downloaded 11 Nov 2012 to 132.210.244.226. Redistribution subject to AAPT license or copyright; see http://ajp.aapt.org/authors/copyright_permission Aerodynamic forces acting on a ball in flight increase with the speed and diameter of the ball but do not depend on the mass of the ball. The trajectory of a light ball therefore provides a more sensitive measure of the effect of the aerodynamic forces"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003761_9781118818428-Figure20.4-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003761_9781118818428-Figure20.4-1.png",
        "caption": "Figure 20.4 Finkelstein\u2019s (1960) notation for parallel-bore \u2018beta\u2019 configuration supplemented by notation required by the design charts. The layout maximizes power-producing potential per net cubic displacement, while minimizing the number of parameters required for thermodynamic study \u2013 and thus for thermodynamic design",
        "texts": [
            " This level of pressure looks to be within the capacity of a moulded, fibre-reinforced plastic crank-case. An engine delivering 100 We+ would extend the operating range of battery-powered mobility buggies used by the disabled, or power an entry to the Shell \u2018Eco-Marathon\u2019, or be a candidate for other modest duties. This prospect defines the hot air engine for the purposes of illustrating a method of design: The focus will be on internal design rather than mechanical construction. Placing of the crank mechanism uppermost in Figure 20.4 is of no significance except for consistency with ad hoc heating (e.g., by a natural convection flame) for preliminary trials. For an inspirational tour of machining and fabrication considerations, the 1993 account by Ross is unsurpassed. Figure 20.4 might be thought of as a \u2018virtual\u2019 engine, since all (absolute) numerical values are left floating. Table 20.1 fleshes out the geometry somewhat (dimensionless lengths \u2013 an angle is the ratio of lengths). The choice of values for kinematic volume ratio \u03bb and kinematic phase angle \u03b2 are not \u2018optima\u2019 by any quantitative criterion. They are, however, consistent with low compression ratio and corresponding low pressure swing. This should maximize prospects for easy starting and for operation at intermediate temperatures",
            " Literal implementation would involve monitoring the behaviour of some 1025 individual molecules, but the algorithms relevant to the eventual task can be set up and exercised on a tiny fraction of that number. The exercise is a way to estimating the minimum number of molecules required to give a picture of macroscopic gas behaviour \u2013 information which might otherwise have to await development of further generations of computer. The principle of the closed-cycle regenerative engine can be demonstrated in a simpler mechan- ical embodiment than is possible with other reciprocating cycles. The result is the \u2018hot-air\u2019 engine already illustrated at Figure 20.4. This can be realized in as few as three moving parts. A power range not well served1 by the IC engine is that between 25 and 100 W. The former figure would power a lap-top computer. The latter the portable oxygen concentrators used in medical emergency, and currently restricted to one hour\u2019s operation per 1 kg electric battery. It is possible that, with appropriate thermodynamic design, the use of modern materials, and by resorting to pressurization, this power range would come within the scope of the \u2018hot-air\u2019 type"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001145_02640410802641392-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001145_02640410802641392-Figure1-1.png",
        "caption": "Figure 1. Schematic of test set-up to measure the dynamic response of softballs.",
        "texts": [
            " Previous work that considered the load-displacement response of baseballs under quasi-static loading provided an unreasonably low modulus (Shenoy, 2000). The following describes a technique intended to represent the rate and magnitude of deformation that would occur in a ball during play. Balls were experimentally characterized by measuring forces from rigid wall impacts. Details of the method are presented elsewhere (Smith, 2008). The apparatus consisted of three piezoelectric load cells (PCB model 208C05) mounted between a rigid wall and a solid cylindrical impact surface, as shown in Figure 1. The design is under consideration as an ASTM (American Society for Testing and Materials) test method. It is robust, where the mean stiffnesses of 12 balls measured from two separate fixtures were within 3%. The 57-mm (2.2-inch) diameter cylindrical impact surface corresponded to the bat diameter. Test speeds ranged from 27 to 49 m s71 (60 to 110 mph). A representative force\u2013time curve is presented in Figure 2. Integration of the force\u2013 time curve provides an impulse that can be compared with ball speeds"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003283_00207721.2012.683834-Figure9-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003283_00207721.2012.683834-Figure9-1.png",
        "caption": "Figure 9. Flexible spacecraft with Stewart platform.",
        "texts": [
            " BFP i is given by BFP i \u00bc ki3\u00f0xi3 xi1\u00de \u00fe ci3\u00f0 _xi3 _xi1\u00de ki1\u00f0xi1 xi2\u00de ci1\u00f0 _xi1 _xi2\u00de: \u00f030\u00de The attitude and translational motion are written as mp \u20aczp \u00bc \u00f0PFB i 1i\u00de, \u00f031\u00de where z\u0308p is the displacement of payload centre of mass and mp is the payload mass. 1i denotes the unit vector of leg i and 1i\u00bc (Ai Bi)/kAi Bik. The payload is subject to moment M M \u00bc ri PFB i 1i, \u00f032\u00de where ri is the vector from the payload centre of mass to the ith payload joint. The application of 6-DOF vibration isolation and precision pointing/steering in astronautics is investigated. Figure 9 shows the flexible spacecraft with Stewart platform. State feedback control can be used for angular manoeuvring. However, there are attitude jitters which are difficult to remove due to sensor noise, disturbances and the coupling between the flexible structure and the controller. Stewart platform can suppress the jitters and vibrations. The dynamic modelling of flexible spacecrafts is mature (Likins and Fleischer 1971; Gennaro 2002). It is expressed by _) \u00bc j dBSxB, _j \u00bc Cj\u00fe K)\u00f0 \u00de \u00fe CdBSxB, J _xB \u00bc xB JxB \u00fe dTBSj \u00fe dTBS Cj\u00fe K) CdBSxB\u00f0 \u00de \u00feQrot, 8>>>>< >>>>: \u00f033\u00de where uB is the angular velocity, K and C are the stiffness and damping matrix of the flexible structure, respectively, J \u00bc Jsys dTBSdBS, where Jsys is the moment inertia of the spacecraft, dBS is the coupling matrix between the spacecraft and flexible structure, Qrot is the moment for attitude control"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001366_secon.2007.342901-Figure3-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001366_secon.2007.342901-Figure3-1.png",
        "caption": "Figure 3. Frames of reference used in the model",
        "texts": [
            " Model In this paper, we will use the vector convention from Fossen [2] where n eb v which denotes the vector quantity v of b frame with respect to e frame expressed in the n frame. 1-4244-1029-0/07/$25.00 \u00a92007 IEEE. 273 2.1. Quad-Rotor Dynamics As discussed above, the quad-rotor UAV is an inherently underactuated system. While the angular torques are fully actuated, the translational forces are only actuated in the zdirection. The forces and torques are expressed as 3 1 3 2 3 4 0 0 TF f T F t F u F u u u           = \u2208 = \u2208 (1) where ( )F f F t is the translational force vector and ( )F t F t is the torque vector. Figure 3 shows the three coordinate frames used to develop the kinematic and dynamic models. Three equations describe the rigid-body motion of the quadrotor UAV [3] I FI IF F IF R vx = (2) ( )1 3 F F F FF IF IF IF I fm mS v N mgR e Fv \u03c9      = \u2212 + \u22c5 + + (3) I I F F F IFR SR \u03c9      = (4) where ( ) 3I IF tx \u2208 is the time derivative of the position of the UAV frame, ( ) 3F IF v t \u2208 is the translational velocity of the UAV, ( ) 3F IF t\u03c9 \u2208 is the angular velocities of the UAV and is calculated directly without modeling the angular dynamics, and ( ) ( )3 I F R t SO\u2208 is the rotational matrix that transforms the vectors from the UAV frame to the inertia frame",
            " To convert between ( )I FR t and ( )I F t\u0398 , cos cos sin cos cos sin sin sin sin cos sin cos sin cos cos cos sin sin sin cos sin sin sin cos sin cos sin cos cos I F R \u03c6 \u03b8 \u03c6 \u03c8 \u03c6 \u03b8 \u03c8 \u03c6 \u03c8 \u03c6 \u03b8 \u03c8 \u03c6 \u03b8 \u03c6 \u03c8 \u03c6 \u03b8 \u03c8 \u03c6 \u03c8 \u03c6 \u03b8 \u03c6 \u03b8 \u03b8 \u03c8 \u03b8 \u03c8 \u2212 + +   = + \u2212 +   \u2212  (9) is used [4]. 2.3. Camera Kinematics As stated, the quad-rotor can thrust in the z-direction, but it cannot thrust in the x- or y-directions. Since the quadrotor helicopter is underactuated in two of its translational velocities, a two actuator camera is added to achieve six degrees of freedom (DOF) control in the camera frame. A tilt-roll camera is added to the front of the helicopter as seen in Figure 3. With the new camera frame, there are now three rotations and three translations, a total of six DOF, to actuate. To control any of the DOF, either the camera must move, the UAV must move, or both. 2.4. Tilt-Roll Camera on Front of UAV The rotation matrix between UAV frame and Camera frame seen in Figure 3 is sin cos sin sin cos sin cos 0 cos cos cos sin sin t r t r t F C r r t r t r t R \u03b8 \u03b8 \u03b8 \u03b8 \u03b8 \u03b8 \u03b8 \u03b8 \u03b8 \u03b8 \u03b8 \u03b8                 \u2212 = . \u2212 \u2212 (10) Since only two of the angles vary, the Jacobian can be redefined as ( ) 0 cos 1 0 0 sin t C front t J \u03b8 \u03b8                 = \u2212 (11) and finally ( ) [ ] T 2\u03c9 \u03b8 \u03b8 \u03b8 \u03b8= , = \u2208 F C C t rFC C front J (12) which facilitates the calculation of the angles of the camera. 3. Control Development 3.1. Translational error formulation The control system developments begins by defining the velocity error between the desired camera velocity, ( ) 3C ICdv t \u2208 , and the actual velocities in the camera frame, "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001366_secon.2007.342901-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001366_secon.2007.342901-Figure2-1.png",
        "caption": "Figure 2. DraganFlyer X-Pro",
        "texts": [
            " The flight control interface allows the user to command motions in the camera frame of reference \u2013 a natural perspective for surveillance and inspection tasks. A nonlinear velocity controller, derived using Lyapunov stability arguments, produces simultaneous complementary motion of the quad-rotor vehicle and the camera positioning unit. The controller achieves Globally Uniform Ultimate Boundedness (GUUB) on all velocity error terms. 1. Introduction The quad-rotor unmanned aerial vehicle (UAV) shown in Figure 2 can be used for civilian and military surveillance and inspection tasks. The six-degree of freedom rigid body craft is positioned by changing the relative speed of the four rotors. These speed differences can produce torques about the roll, pitch, and yaw axes in addition to the thrust produced as the sum of the four rotating blades. Since the system is inherently under-actuated, there are only four control inputs to directly control the six degrees of freedom, makes performing surveillance and inspection tasks challenging"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000301_gamm.200890001-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000301_gamm.200890001-Figure2-1.png",
        "caption": "Fig. 2 Vacuum gripper flanged to link n\u22121, load as link n, link frame n and selected forces (from [36]).",
        "texts": [
            " These limitations require elaborate control strategies to prevent the contact between gripper and load from breaking off especially in time-optimal motions with high acceleration forces active. A direct inclusion of the gripper and load model into the Newton-Euler formalism is possible. The gripper is fixed to link n \u2212 1 and modeled as part of this link; it thus affects the inertia tensor In\u22121, the center of mass CM n \u2212 1 and the vectors Pn\u22121 and pn\u22121. The load of mass mn is modeled as link n connected to link n \u2212 1 by the immobile (artificial) joint n (\u03b8\u0307n = 0, \u03b8\u0308n = 0). The position of link frame n relative to link frame n \u2212 1 is fixed. Fig. 2 gives a schematic view of the gripper and load geometry. c\u00a9 2008 WILEY-VCH Verlag GmbH & Co. KGaA, Weinheim Full contact between load and gripper is maintained as long as the following two constraints are not violated during operation C1(t, x, u) := \u3008n\u3009V\u0307n T \u00b7 D110 \u00b7 \u3008n\u3009V\u0307n \u2212 ( ft mn )2 \u2264 0 , (31) C2(t, x, u) := \u2212\u3008n\u3009V\u0307n \u00b7 ez|n \u2212 fp mn \u2264 0 , Fp = \u2212fp \u00b7 ez|n . (32) D110 := diag (1, 1, 0) \u2208 IR3\u00d73 is a diagonal matrix, fp > 0 denotes the perpendicular and ft > 0 the tangential force limit. (32) is an unilateral constraint"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000169_ijmic.2006.011940-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000169_ijmic.2006.011940-Figure1-1.png",
        "caption": "Figure 1 A ball on a plate",
        "texts": [
            " Jintao Su is an MS candidate at the School of Communication Engineering, Jilin University. His research interests include visual servo and intelligent control. A ball moving on a beam is a typical non-linear dynamic system, which is often applied to proof-test diverse control schemes. The ball and plate system is the extension of the traditional ball and beam system (Hauser et al., 1992; Xinzhe et al., 2003). A ball moving on a rigid plate is unconstrained only except contacting with the plate (Naoyoshi et al., 1997), as shown in Figure 1. The ball and plate system is a typical multi-variable and super articulated non-linear plant. The ball is able to move freely and has no ability to recognise the environment. So it cannot control its behaviour by itself (Naoyoshi et al., 1997). All of these factors make the movement of the ball on the plate difficult to control. Control problems of the system consist of the ball\u2019s position control, trajectory tracking and barrier passing, which is comparatively complex (Xinzhe et al., 2003). The ball\u2019s position control problem is to make the ball arrive at a point from any other point accurately and as soon as possible"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003642_0954410012467716-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003642_0954410012467716-Figure2-1.png",
        "caption": "Figure 2. Planar relationship between the aircraft and the defined path.",
        "texts": [
            " The thrust force FT is generated by the propeller FT \u00bc n 2D4CFT \u00f08\u00de where n represents the speed of the engine that is modeled as a first-order low-pass filter with the time constant 0.4 s; D is the diameter of the propeller; CFT is the dimensionless thrust coefficient.24 The framework of the path-following method is shown in Figure 1. In our article, the outer guidance logic is based on the notion of the virtual target. The philosophy of this method is similar to that of the missile guidance approach. Figure 2 illustrates the planar geometric relationship between the aircraft and the defined path with no wind. In Figure 2, the positive direction of the X axis designates the North and the positive direction of the Y axis represents the East. The flight path is specified by the waypoints Wk and Wk\u00fe1. There is equivalence between the line segment R and the line of sight (LOS) in the missile guidance. The virtual target begins to move along the desired path from the intersection point of the aircraft\u2019s velocity direction and the path. To track the virtual target, the aircraft adapts its heading so that its velocity direction aligns with the LOS and gets close to the desired path. Figure 1. Framework of the path-following method. at COLUMBIA UNIV on October 12, 2014pig.sagepub.comDownloaded from In Figure 2, the difference between the reference heading and the real heading is defined as \u00bc ref \u00f09\u00de where ref represents the reference heading of the flight path and represents the real heading of the aircraft. The dynamics of can be derived as : \u00bc : \u00bc Vw sin R \u00f010\u00de where Vw represents the speed of the virtual target. The asymptotic convergence performance of the proposed guidance logic can be analyzed by using the Lyapunov theory. Define the following Lyapunov function as V1 \u00bc 1 2 2 \u00f011\u00de Differentiating V1 and using equation (10) _V1 \u00bc _ \u00bc Vw sin R \u00f012\u00de The speed of the virtual target is explicitly selected as where V is the aircraft speed"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003427_mmar.2012.6347818-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003427_mmar.2012.6347818-Figure1-1.png",
        "caption": "Fig. 1. Comparison between ideal and real sliding mode control for z \u2208 R2 with the sliding surface s (z) = 0.",
        "texts": [
            " As shown in the following, this can be achieved by a variable structure control law. The switching component of the controller guarantees a motion of the system states towards a so-called sliding surface s (z) = 0 which coincides with those regions in the state-space that are consistent with the desired output yd (t). In phase 2, after reaching the sliding surface at the point of time t\u2217, a tracking control strategy is employed which is designed to guarantee s (z) = 0 for all t \u2265 t\u2217. This two-stage control procedure is visualized in Fig. 1a. As already mentioned above, it is not always guaranteed that s (z) = 0 holds in practical implementations of sliding mode control procedures for all t \u2265 t\u2217. This is caused by the fact that disturbances acting on the system as well as measurement errors and an imperfect observer-based state reconstruction can cause the system to leave the sliding surface even though the point of time t\u2217 at which it has been reached has been detected perfectly. Moreover, also the exact point of time t\u2217 at which the system reaches the sliding surface s (z) = 0 can most likely not be determined exactly in real control systems since sliding mode controllers are usually implemented on digital signal processors. Therefore, it is only possible to check reaching the sliding surface at discrete points of time. Both phenomena lead to chattering in practice. This means that the switching component of the control law changes its sign with high frequency, see Fig. 1b. To reduce actuator wear and noise caused by switched activation of the actuators, the amplitude and number of these switchings has to be reduced. One possibility to achieve this goal is the introduction of the regularization strategy presented in the following sections. Without significant loss of generality for practical control applications, systems (1), (2) are considered that can be reformulated in terms of a nonlinear input-affine set of ODEs x\u0307 (t) = a (x (t)) + b (x (t)) \u00b7 u (t) (3) with the state vector x \u2208 Rn, the scalar input u (t) \u2208 R, a (x (t)) \u2208 Rn, b (x (t)) \u2208 Rn, and the output y (t) = x1 (t)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001246_j.triboint.2009.05.001-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001246_j.triboint.2009.05.001-Figure2-1.png",
        "caption": "Fig. 2. Hydraulic circuit of the test rig showing the direction of main flow.",
        "texts": [
            " When rotating in a certain direction, a hydraulic motor can operate in driving mode when direction of torque is positive, or in breaking mode when direction of torque is negative. In this test two identical motors are mounted on the same shaft, one motor in driving mode and the other in breaking mode. When the two motors are pressurized, the only power needed to rotate the shaft is the power to overcome the friction losses and the oil flow losses in the system. These losses are called the hydromechanical losses. The hydraulic circuit is described in Fig. 2. The high pressure, PH, is the same for both motors and is held constant with a high pressure pump and a pressure control valve. The flow from the high pressure pump compensates for the leakage in the motors. The low pressure of the motor in driving mode, PLM, is held constant with a pressure control valve. The main flow is to the low pressure inlet of the motor in breaking mode and a pressure control valve is mainly used as a safety valve. The main flow is indicated with arrows in Fig. 2. The shaft torque of the hydraulic motor is dependent on the differential pressure between the inlet and outlet. In this case the high pressure of both the motors and the low pressure of the motor in driving mode are kept constant. The only pressure that is varied is the low pressure on the motor in breaking mode, PLP. When PLP is increased, the torque of the motor in breaking mode will decrease until it is equal to the torque of the motor in driving mode and the shaft starts to rotate. Because the displacement of the two motors is the same, the difference in pressure between PLP and PLM is a measure of the hydromechanical losses in the system",
            " To separate friction losses from oil flow losses, oil flow losses from earlier tests have been used. A loss coefficient C for the pistons is introduced, assuming that all friction losses originate from the piston assembly. Torque equilibrium for the motor in driving mode gives PH Ploss 2 \u00f01 C\u00de Ts PLM \u00fe Ploss 2 \u00f01\u00fe C\u00de Ts T \u00bc 0 ) C \u00bc Ts \u00f0PH PLM\u00de T Ts \u00f0PH \u00fe PLM\u00de Ploss \u00f0PH \u00fe PLM\u00de where Ploss is the pressure loss between inlet and outlet of the motor, Ts is the specific torque of the motor in Nm/bar, T is the torque measured during the test. Other parameters are as shown in Fig. 2. Using the oil flow losses from earlier tests, the loss coefficient C is 0.02 for piston A and 0.03 for piston B at 180 rpm and 300 bar. The roller in piston B has a low friction DLC coating expected to enhance scuffing resistance. However, the motor tests have shown that the piston B, with 50% higher losses, has a lower seizure resistance than piston A. The main reason for difference in friction losses is in view of the differences in form and surface finish, which influence the hydrodynamic lubrication"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002259_amm.229-231.710-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002259_amm.229-231.710-Figure1-1.png",
        "caption": "Fig 1 Geometry of the model used in the analysis Fig.2. Solid model of hard faced circular grid plate Fig.3. Meshed model used in analysis",
        "texts": [
            " Chemical composition of SS304 and hard faced colmonoy is shown in table 1. Table.1 Chemical composition of SS304 and hard facing alloy Elements % by weight SS304 Colmonoy C 0.1 0.6 Mn 1.6 0.1 Si 0.69 3.8 P 0.04 - S 0.022 - Cr 18.1 11.5 Ni 8.4 77.5 V 0.48 - Fe 69.93 4.4 B - 2.6 Cu 0.33 Finite element modeling and analysis are performed by using ANSYS 12.0 to predict residual stresses and distortion within the hard faced grid plate. Model Description. A Circular grid plate with inner and outer diameter of 780 mm and 1000 mm respectively and 30 mm thick as shown in Fig 1 and the 3D solid model is shown in Fig. 2. The hard faced material is modeled to a total thickness of 4mm by filling the groove of 3 mm and 1 mm above the surface of the plate. The hard facing process simulated is a single pass plasma transferred arc hard facing. The finite element model in this study employs the technique of element birth and death. All elements have to be created, including those that are born in the later stages of the analyses. Mesh creation. Using 8-noded brick element, the solid model is meshed"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003560_cdc.2012.6425834-Figure3-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003560_cdc.2012.6425834-Figure3-1.png",
        "caption": "Fig. 3. Schema of the camera\u2019s view.",
        "texts": [
            " Section VI presents the Quad-plane experimental platform. The behavior of the Quad-plane during real-time experiments is shown in Section VII. Finally, some conclusions and future works are presented in section VIII. The road following procedure can be detailed as follows. First, the vehicle performs an autonomous take-off, reaching a desired altitude zd over the road. Then, the heading of the vehicle \u03c8, is controlled to obtain a parallel positioning between the Quad-plane x-axis xh and the longitudinal direction of the highway xr (Fig. 3). The Quad-plane forward velocity x\u0307 is maintained constant while the distance between 978-1-4673-2064-1/12/$31.00 \u00a92012 IEEE 3110978-1-4673-2066-5/12/$31.00 2012 I xh and xr, expressed by ey, is controlled and maintained to a minimum value, achieving a flight path well aligned and centered w.r.t. the highway at constant forward velocity. Finally the Quad-plane performs a landing in a position near to the end of the highway. All of these actions should be accomplished completely autonomously without the previous knowledge of shape of the highway and the terrain",
            " For the purpose of control we will use the linear model (1) instead of the nonlinear version, since in general the vehicle operates in areas where |\u03b8| \u2264 \u03c0/2 and |\u03c6| \u2264 \u03c0/2, these constraints are satisfied even when the nonlinear model is used together with a feedback control [2], [7]. The corresponding coordinate system is represented in (Fig.2). x\u03071 = x2 \u03b8\u03071 = \u03b82 x\u03072 = \u2212\u03b8u1 \u03b8\u03072 = u2 y\u03071 = y2 \u03c6\u03071 = \u03c62 y\u03072 = \u03c6u1 \u03c6\u03072 = u3 z\u03071 = z2 \u03c8\u03071 = \u03c82 z\u03072 = 1\u2212 u1 \u03c8\u03072 = u4 (1) IV. VISION-BASED STATES ESTIMATION A sufficiently smooth road is represented in the camera\u2019s image as a group of lines, (Fig. 3). A straight line in the image can be seen as a segment of infinite length and whose center of gravity belongs to the straight line [8]. Using the probabilistic Hough transform method available in OpenCV [9], a straight line will be represented as \u03c1 = x cos \u03b8+y sin \u03b8. The center of gravity (xg, yg) of each line detected can be computed as xig = cos(\u03b8) and yig = sin(\u03b8). Where the index i is related to the line i. Let\u2019s define (xiI , y i I) as the initial point of the line, located below the image\u2019s lower margin, and let (xiF , y i F ) be the line\u2019s final point located above the image\u2019s upper margin",
            " The parallel lines extracted from the road\u2019s projection on the image are grouped together with the objective of obtaining a mean line, which will uniquely represent the road in the image with a pair of initial coordinates xI = xi I i , yI = yi I i and final coordinates xF = xi F i , yF = yi F i . Where i is the number of lines representing the road, (xI , yI) stands for the initial road coordinate and (xF , yF ) stands for the final road coordinate. The angle between the camera\u2019s xc axis and the point (xF , yF ) can be computed as \u03c8r = arctan (yF \u2212 yI , xF \u2212 xI), this value is used for obtaining the heading angle \u03c8d that will align the vehicle\u2019s longitudinal x-axis (xh) with the road\u2019s longitudinal axis (xr), (Fig. 3). Finally we can express \u03c8d as \u03c8d = \u03c8r + \u03c0 2 (4) the therm \u03c0 2 is added to adjust \u03c8d to zero when \u03c8r is vertically aligned with xh. Lets consider an image-based distance ecx along the camera\u2019s xc axis, which is defined between the road\u2019s center of gravity projection (xg , yg) and the vehicle\u2019s center of gravity projection (x0, y0), (Fig. 3). If xin > xfi, then ecx = ( xI\u2212xF 2 + xF ) \u2212 cw 2 , where cw represents the image\u2019s width in pixels. If xI < xF , xI must be replaced by xF and vice-versa. Using ecx, the lateral position of the aerial vehicle relative to the road is estimated as ey = z ecx \u03b1x (5) where z represents the Quad-plane\u2019s height and \u03b1x is the focal length of the camera, in the xc direction. By implementing an optical flow algorithm to the camera\u2019s image, translational velocities can be estimated. Consider that the arrangement camera-vehicle is moving in a 3-D space w",
            " In both cases the control objective is to regulate the y1 state to the origin, i.e. y1d = 0. The control schema when the line is detected, proposes a feedback control law given by \u03c61 = 1 u1 (\u2212kL3y1 \u2212 kL4y2 + kLI\u03be) (12) where kL3, kL4 and kLI are positive real numbers. Here we have introduced the additional state \u03be, where \u03be dynamics are given by \u03be\u0307 = y1d\u2212y1 = \u2212y1. Defining a distance named dc as the lateral position existing between the vehicle\u2019s center of gravity projection and the point where the camera loses the image of the highway (Fig. 3), a change of coordinates can be made such that yd1 = y1 + dc and y\u0307d1 = yd2 = y2 is its derivative. Thus, using the control (12), the closed-loop system of the y dynamics is given by y\u0307d1 = yd2 y\u0307d2 = \u2212kL3yd1 \u2212 kL4yd2 \u2212 kLI\u03be \u03be\u0307 = \u2212yd1 (13) with eyd = (yd1 , yd2 , \u03be)T . Then, (13) can be represented as e\u0307yd = ALeyd where AL =   0 1 0 \u2212kL3 \u2212kL4 kLI \u22121 0 0   (14) The control schema when the line is not detected proposes the same structure given by (12), with the difference of having a set of alternative gains"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000580_1081286507077107-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000580_1081286507077107-Figure1-1.png",
        "caption": "Figure 1. The pre-bifurcation state in a pre-stressed annular thin lm subjected to a uniform displacement eld on the outer boundary (r R2), and azimuthal shearing on the inner rim (r R1). The radial and orthoradial stress distributions shown in a are both tensile but, as indicated in b , the lm might undergo partial compression in the direction of one of the principal stresses (here 2).",
        "texts": [
            " An earlier attempt that questions TFT is the work of Rimrott and Cvercko [15], who studied gravityinduced wrinkling in vertical membranes, a problem already discussed by Mansfield a few years before [16] with the help of a particular version of TFT. The former authors showed that the introduction of a critical compressive stress perpendicular to wrinkles leads to results that better match the experimentally observed patterns. In this work we revisit the wrinkling instabilities of a stretched annular membrane subjected to uniform in-plane shear on the inner boundary (see Figure 1(a)). The point of view taken in what follows differs from the previous studies in that we regard wrinkling as a typical buckling instability for thin plates rather than membranes. In other words, the problem we investigate involves a bifurcation of plane-stress deformation to an out-of-plane bendingtype deformation, and this introduces several complications into the analysis. The most severe one is represented by the presence of variable coefficients in the governing equation which involves a complex-valued fourth-order partial differential operator",
            " The insight gained from this study proves useful in guiding us through the formal calculations in the sequel. Asymptotic approximations for the neutral stability curves and their envelope are obtained with the aid of a WKB-type analysis in Section 4 the versatility of this approach is further confirmed by comparisons with direct numerical simulations of the original eigenproblem. The paper concludes with a discussion of our results and considers some possible extensions to related problems. The general setting is that shown in Figure 1(a): an annular thin film of inner radius R1, outer radius R2, and thickness h (h R2 1) is initially stretched by imposing the uniform displacement field U0 0 along the outer edge (r R2) while the inner boundary (r R1) is rotated through some (small) angle by applying a torque M . The solid film will be modelled using thin plate theory [26]. A cylindrical system of coordinates r z is used to define various quantities of interest associated to the problem, and the notation employed below is standard (see [27], for example)",
            "comDownloaded from where we have introduced the differential operators 2 2 1 1 2 2 1 1 2 2 2 0 1 2 and two auxiliary constant expressions, A 1 1 2 and B 2 1 1 2 (6) Next, let 1 1 and 2 2 denote the principal stresses in the unwrinkled plate, so that (see [27], for example) 1 2 rr and 1 2 rr r 2 Clearly, 1 2 0 for all [ 1], but 1 2 depends on and this product may become negative in some parts of . When this happens, one of the two principal stresses, 1 or 2, is compressive and the film will experience compression in the direction of the negative principal stress (for the sake of definiteness, we have assumed that 1 2 in Figure 1). The minor principal stress becomes zero along the circle , with 4 4 1 2 2 1 2 2 4 1 2 1 4 (7) and it can be easily checked that , as long as 4 2 1 2 (8) The expression that appears on the right-hand side of the inequality in (8) represents the loading parameter threshold that marks the onset of compressive orthoradial stresses in the film. These are confined to the annular region between and , and an approximate distribution of the principal stresses is sketched in Figure 1(b). Furthermore, it can be seen that the compressive region will span the entire breadth of the annular plate if 2 1 2 4 1 2 1 2 (9) However, owing to the presence of finite bending rigidity, wrinkling will not be triggered as soon as the film experiences compression. As shown in Section 4, the critical must exceed , its precise expression involving the number of wrinkles, as well as the large parameter . Since and the bifurcation analysis undertaken in this work is linear, (9) will play no part in what follows"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001247_0954406jmes321-Figure11-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001247_0954406jmes321-Figure11-1.png",
        "caption": "Fig. 11 Applied coordinate system of tooth contact analysis",
        "texts": [
            " The numerical results for TCA show that the proposed double-crowning modification of the tooth geometry does indeed enable localization and stabilization of the bearing contact and absorbs the discontinuity of the adjacent motion curves. The continuous tangency of the involute cylindrical pinion surface 1 and the modified face-gear surface \u2032 2 is guaranteed by equations (30) and (31). However, to perform the TCA, the equations of the face-gear surface \u2032 2 and involute cylindrical pinion surface 1 should be represented in the same fixed coordinate system Sf . Figure 11 shows the TCA coordinate systems. An auxiliary coordinate system Sm is used for simulation of alignment errors such as axial displacement L, change to the shortest centre distance E , and change to the shaft angle \u03b3 . All errors of alignment refer to the pinion. Proc. IMechE Vol. 221 Part C: J. Mechanical Engineering Science JMES321 \u00a9 IMechE 2007 at UNIV OF CONNECTICUT on April 13, 2015pic.sagepub.comDownloaded from At any instant, the pinion surface 1 can mesh with the gear surface \u2032 2 in point contact"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001978_s10846-010-9408-9-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001978_s10846-010-9408-9-Figure2-1.png",
        "caption": "Fig. 2 Detection direction of omni-directional mobile robot",
        "texts": [
            " Inspired by the mechanism of Jerne\u2019s idiotypic network hypothesis, the artificial immune system in this paper is constructed as follows: The omni-directional mobile robot is taken as the immune agent, and it is equipped symmetrically but unevenly with eight virtual sensors around according to the different contributions of environment information to path planning. The detection distance of virtual sensor is dived into three grades: near, mid and far, and the detection direction is 1\u223c8 as shown in Fig. 2. The robot can move according to eight different moving instructions C = {a, b , c, d, e, f , g, h}, i.e. {forward, left forward, right forward, left move, right move, left back, right back and back}, and the stop instruction is adopted at the goal. The environment surrounding the robot is regarded as antigen, and the antigen epitope is the environment information including obstacles and goal. The robot action aiming at the environment is regarded as antibody, the antibody idiotope is environment information too, and the antibody paratope is the moving instruction",
            " . . ,8, AbGi \u2208 {0, 1}. The elements of the above two sets are defined as the same way as in antigen. Definition 3 In this paper, the antigen-antibody affinity \u03b6 is used to evaluate the matching degree between the antigen and antibody. In the biological immune system, the genes of antigen or antibody at different positions are of equal importance. However, in this paper, the coded information of obstacle is taken as genes of antigen or antibody, and the genes give different contributions. From Fig. 2, it can be seen that the contribution of the frontal information of the robot is bigger than its rear information, the foremost contribution is the biggest, the hindmost contribution is the least, and the left and right contribution are symmetrical and equal. To embody the different contributions of information at different positions in path planning, the antibody-antigen affinity \u03b6 is defined as follows: \u03b6 = \u239b \u239c\u239c\u239c \u239d 1 \u2212 8\u2211 i=1 AGi 2 \u239e \u239f\u239f\u239f \u23a0 /( 1 + 8\u2211 i=1 \u03b5i AOi ) (3) Where, AGi = (AgGi\u2295AbGi) denotes the matching information of goal between the antigen and antibody at direction i, the sign \u2295 represents XOR logical operator, AOi = (AgOi\u2295AbOi) denotes the matching information of obstacles between the antigen and antibody at direction i, \u03b5i is the contribution weight of obstacle information, and \u03b5 = {2, 0"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003953_j.compeleceng.2013.07.016-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003953_j.compeleceng.2013.07.016-Figure1-1.png",
        "caption": "Fig. 1. Construction tele-robotic system with force guidance.",
        "texts": [
            " The rest of the paper is structured as follows: in Section 2, we describe structure of tele-robotic system with force guidance, involving a bilateral section\u2013master and slave. Then we describe the 3-D reconstruction of task object and ground surface by the stereo vision camera in Section 3. Based on the artificial potential field theory, Section 4 shows the Generation of guiding forces, including the attractive force from task object and repulsive force from obstacles. In Section 5, comparison experiments are shown under the circumstances of force guidance and other situations. Finally, Section 6 concludes the paper. Fig. 1 shows the block diagram of the tele-robotic system with vision-based force guidance. The system involves a bilateral master and slave. The master is controlled by an operator and mainly consists of two force feedback joysticks and a screen with computer graphics (CG). For force feedback joysticks, Microsoft SiderWinder2 force feedback joysticks are adopted, which are capable of delivering around 100 different forces and 16 programmable buttons (8 action buttons plus 8-direction hat). The slave is composed of a construction robot and a stereo vision camera called \u2018\u2018Bumblebee\u2019\u2019"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003973_s12206-012-0811-y-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003973_s12206-012-0811-y-Figure2-1.png",
        "caption": "Fig. 2. The Pro/E wildfire model.",
        "texts": [
            " The piston-cylinder assembly consists of four components, which are piston, cylinder, connecting rod and crank shaft. The piston is made of aluminum alloy and the crank shaft and connecting rod are made of alloy steel. In order to maintain the compression ratio at a constant value, the tolerance of the clearance has to be maintained at a particular value. So the objective is to allocate appropriate tolerance in order to maintain sufficient clearance between the piston top surface and the cylinder top surface. First, a 3D model of the assembly is created using Pro/E wildfire 5.0 software as shown in Fig. 2. To determine the features which have an effect on the clearance measurement, a vector loop model of the assembly has been created as shown in Fig. 1. The assembly function that describes the quality value is: (0.7) (0) (0) (0) (50) (174) (69) (292.3) 2 4 6 3 5 7 1x x x x x x x x .\u2211 = + + + + + \u2212 (1) Once a three-dimensional model of the assembly is created, the next step is to develop a finite element model of the same to determine deformation of various components and their effect on the clearance"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001704_acc.2009.5160384-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001704_acc.2009.5160384-Figure1-1.png",
        "caption": "Fig. 1. Basic configuration",
        "texts": [
            " The local control is in the form of a tuple (\u03b8, d), which is interpreted by each agent in its local frame of reference. Here, \u03b8 refers to the angle by which each agent changes its orientation, and d is a scalar that refers to the distance by which each agent moves after effecting the orientation change. Note that, the broadcast mechanism (\u03b8, d) is the same for all the agents. Also, the local frame of reference for each agent is centered at the agent\u2019s location and its reference axis is oriented along its current orientation. As an illustration see Fig. 1, where agents are shown located initially at xi0 with initial orientation \u03b8i0 in the global reference frame. If the control command broadcast to all the agents is (\u03b8, d), then the agents implement it in their local coordinate frame by each of them changing their orientation by the same angle \u03b8 and advancing by the same distance d to reach the final destination xif . Even in this figure it can be seen that by doing this the agents have come closer to each other. Our objective is to determine a (\u03b8, d) such that the agents can achieve the closest proximity with each other"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001336_amc.2008.4516091-Figure5-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001336_amc.2008.4516091-Figure5-1.png",
        "caption": "Fig. 5. Generation of Force to Robot",
        "texts": [],
        "surrounding_texts": [
            "Upper structure of vehicle consists of three manipulator links and the vehicle body on the two coaxial wheels. Centre Of Gravity (COG) of upper structure is controlled to obtained the stable posture. Relational expression between joint velocity vector q = (\u03b80, \u03b81, \u03b82, \u03b83) and the COG velocity vector x is obtained by the following equation. x = f(q) (38) x\u0307 = J(q)q\u0307 (39) x\u0308 = J(q)q\u0308 + J\u0307(q)q\u0307 (40) where J is Jacobian matrix. Command of the COG is estimated such that the system is stable. COG command is converted to the joint command value using the Jacobian matrix. In the redundant manipulator, joint space acceleration reference, obtained by using pseudo inverse matrix, can be written as (41). In (41), first term is the work space acceleration and second term represents the null space motion. q\u0308ref = J+(x\u0308ref \u2212 J\u0307(q)q\u0307) + (I \u2212 J+J)\u03c6\u0308 (41) where, J+ is the weighted pseudo inverse matrix and define as (42). J+ = W\u22121(JT JW\u22121JT )\u22121 (42) In the above equation W is a diagonal weighting matrix. In the joint space disturbance observer based acceleration controller, W corresponds virtual inertia matrix Ivn and can be selected arbitrary[9]. By using the joint space and workspace observer in (41) can be rewritten as (43) without calculating the J\u0307(q)q\u0307 and (I \u2212 J+J) terms[10]. q\u0308ref = J+x\u0308ref + q\u0308 ref null (43) Equation (44) gives the acceleration reference in workspace. xref = Khp(x cmd \u2212 x) + Khv(x\u0307 cmd \u2212 x\u0307) (44) where, Khp and Khv are the proportional gain and derivative gain respectively. Figure (7) shows the detailed block diagram of manipulator control."
        ]
    },
    {
        "image_filename": "designv11_12_0001568_ie801291h-Figure9-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001568_ie801291h-Figure9-1.png",
        "caption": "Figure 9. Proposed degradation pathways of CI Reactive Blue 221 during electrooxidation in the presence of a graphite sheet as anode.",
        "texts": [
            " Extracts were dried over anhydrous sodium sulfate and concentrated by evaporation. From the data of GC-MS, 10 major fragments (a1 - a4, b1 - b5, and c) were identified as shown in Table 2. In the first step, cleavage of -NdN-, C-N, and O-Cu bonds in the parent molecule takes place and forms three compounds, viz., a1, b1, and c, which are identified at retention times (tR) of 33.0, 46.42, 7.48 min and having corresponding m/z values of 556, 370, and 154, respectively. Considering the identified fragments, it is expected that three pathways are involved, as shown in Figure 9. The fragments, which formed in the second step, are ultimately converted as CO2 and water. The interesting point here is that triazine (atrazine) was completely mineralized without forming the refractory cyanuric acid. Also, it is 2154 Ind. Eng. Chem. Res., Vol. 48, No. 4, 2009 noteworthy that formation of organic acids is consistent with the fact that the electrolyte pH was intended to shift toward the slightly acidic region during the course of the reaction process. The organically bound nitrogen was measured as total Kjeldahl nitrogen (TKN) before and after the electrooxidation process by the macro-Kjeldahl method (4500 - Norg) as per the standard methods"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000707_s11044-006-9032-4-Figure3-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000707_s11044-006-9032-4-Figure3-1.png",
        "caption": "Fig. 3 The 3-SPS/PU parallel manipulator with loads and its serial mechanism with loads",
        "texts": [
            " In this system, the workloads can be simplified as a wrench (F, T), which includes the inertia wrench of platform due to its mass, the damping wrench of platform due to its velocity, the equivalent inertia wrench of legs due to their masses [10], the gravity of platform and legs, and the external working wrench (such as machining wrench or operating wrench). Here, F is a central force and T is a central torque. (F, T) is applied onto m at o, and is variables versus time. (F, T) or its components (FX , FY , FZ , TX , TY , TZ ) are balanced by a group of active forces Fri exerted on ri (i = 1, 2, 3) and a constrained wrench (Fc, Tc) exerted on ro (Figure 3a). In order to solve the active forces and the constrained wrench, a virtual serial mechanism is constituted (Figure 3b). In the virtual serial mechanism, all the active legs of the parallel Springer manipulator are removed, m of the parallel manipulator is completely supported by the virtual serial mechanism, and each of joints in the virtual serial mechanism bears the wrench (F, T). Based on the principle of virtual work, the total virtual work by active forces Fri (i = 1, 2, 3) exerted on ri must equal the total virtual work by F and T exerted on m at o. Since F and T can be mapped into a group of loads Fei exerted on Ji of the virtual serial mechanism, the total virtual work of Fri exerted on ri of the parallel mechanism must equal the total virtual work of Fei "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002028_icnsc.2010.5461545-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002028_icnsc.2010.5461545-Figure2-1.png",
        "caption": "Fig. 2. A simplified STG",
        "texts": [
            " In comparison with traditional planetary gearbox, STG potentially offers the following benefits [3]: (1) high ratio of speed reduction at final stage; (2) reduced number of speed reduction stages; (3) lower energy losses; (4) increased reliability of the separate drive paths; (5) fewer gears and bearings; (6) lower noise. These benefits have driven the helicopter OEMs to develop products using the STG. For example, the Comanche helicopter was designed with a STG, and the new Sikorsky CH-53K will incorporate the STG design to transmit over 18,000 hps to the rotor blades. It is likely that STG will be incorporated into more designs in the future [4]. A simplified split torque gear drawing [3] is shown in Fig. 2 and a more representative gearbox design, such that seen from the Comanche STG is given in Fig. 3 [5]. T 62978-1-4244-6453-1/10/$26.00 \u00a92010 IEEE Because of the limited experience in building helicopter with STG, there is no condition based monitoring data on this type of gearbox. Studies have been conducted to model and analyze vibration dynamics of the STG [6], and analysis on gear loading has been conducted [5]. Yet, these studies do not give insight into fault detection of gears on this type of design. Gear diagnostics use time synchronous averages to separates in frequency gears that are physically close. As shown in Fig. 2, in a STG, to divide the torque evenly, several identical compound gears will mesh simultaneously with the bull gear. The effect of a large number of synchronous components (gears or bearing) in close proximity may significantly reduce the fault signal (decrease signal to noise ratio) and therefore reduce the effectiveness of current gear analysis algorithms. Only limited research on STG fault diagnosis using vibration sensors has been conducted to date. In a recent paper [7], an investigation on condition indicator performance on a notational STG type gearbox was reported"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003887_j.procir.2015.06.103-Figure3-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003887_j.procir.2015.06.103-Figure3-1.png",
        "caption": "Fig. 3. Screw system of nutation drive",
        "texts": [
            " The cross product of between the 3S and equation (4) can be given as below cos2 cos)(sin)( 21 2 2 2 1 2 22 2 1121122112 3 SSSS SaSaSSaaSShha (8) Then the magnitude, location, pitch and direction of the sum screw can be determined by equations (3), (7) and (8). Particularly, in the nutation drive, the cone vertex of the bevel gear pair is coincidence, and consequently the point A , B and C is coincidence, i. e. 01a and 012 aa . Compare with Fig. 1 and Fig. 2, the screw system of nutation drive are further illustrated in Fig. 3. In nutation gear drive system, bevel gear pair meshes only for pure rotation, and the pitch of the screw is 021 hh . Thus, the pitch of the screw and the distance between sum screw 3$ and coordinate system ),,( ZYXS is 03h and 03a . The meshing between the external and internal spiral bevel gears in the nutation drive can be considered as the external and internal spiral bevel gear meshing with crown gear (an imaginary gear), which has a pitch cone angle of 90 , and the pitch cone is at right angle to its axis"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003531_j.piutam.2011.04.015-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003531_j.piutam.2011.04.015-Figure1-1.png",
        "caption": "Fig. 1: The bipedal SLIP model is shown in panel A). The monopedal ASLIP [15] is shown in panel B). The bipedal ASLIP developed in this work is in panel C). The sign of each angle and torque follows the right hand rule",
        "texts": [
            " Each of these models has its drawbacks when considered in the context of human walking: the passive walking machines display a lot of torso sway relative to a human [14]; the partial gait model [12] assumes that the hip torque applied to the torso does not change the force vector applied to the hip; and the sagittal plane walking model [13] uses a heuristic controller, which has unknown stability properties and exhibits more torso sway than a human does while walking [14]. Recently, Poulakakis and Grizzle [15, 16] used input-output feedback linearization to embed desirable torso dynamics into the controller of a simplified planar robot (Fig. 1B). Their model, the asymmetric spring loaded inverted pendulum (ASLIP), consists of a planar biped with massless, telescoping legs that attach at the hip joint of the torso. The center-of-mass (COM) of the torso is not coincident with the hip joint, adding non-trivial torso dynamics to the system equations of the biped. Poulakakis and Grizzle used the hip torque and leg force (the linear force generated by the telescoping leg) of the model to control the state of the torso. This control was achieved by embedding the dynamics of the desired plant (Fig. 1A), the spring loaded inverted pendulum (SLIP model), into the control laws for the hip torque and leg force of the ASLIP using input-output feedback linearization [17, 18]. Interestingly there is a high degree of similarity between human COM kinematics and ground reaction force profiles to those of the SLIP model during both walking and running [19]. The COM kinematics and ground reaction force profiles can be made to fit simultaneously within \u00b11 standard deviation of human profiles [14] if the point contacts of the SLIP model are allowed to translate forward at a velocity that is similar to the center-of-pressure (COP) velocity in humans [20]",
            " The quality of fit between the SLIP model and human walking and the illustration that this gait can be embedded in more elaborate models [15, 16] inspired the current investigation to determine if human-like SLIP dynamics could be embedded into a sagittal plane human gait model. The investigation first begins by extending Poulakakis and Grizzle\u2019s ASLIP model and control laws to a bipedal ASLIP model in Sec. 2, simulates walking motions, and then proceeds to apply the same control framework to a multibody sagittal plane gait model in Sec. 3. The standard planar SLIP model (denoted with a subscript \u2018S \u2019 in equations) consists of a point mass (m) in a uniform gravity field, g, with two massless linear springs with a fixed resting length (r0) and no preload (Fig. 1A). Each leg behaves like a massless prismatic joint (actuated by forces of magnitude pS ,1 and pS ,2) connected to the ground (during stance) with revolute joints. Together both legs exert a net force of fS x and fS y to the SLIP point mass in the horizontal x and vertical directions y respectively. Although the dynamic equations of the SLIP model, Eqns. 1-2 , are very simple, it can be made to walk or run with human-like ground reaction force and COM kinematic profiles [19] using optimized initial conditions selected to yield limit-cycle walking or running",
            "6 radians apart during swing to maintain a stable limit cycle at this quick walking speed. The stiffness of the legs was chosen to be consistent with the frequency constant ( f = \u221a k/m) of Geyer et al.\u2019s SLIP simulation [19]. The SLIP model can be made to resemble the human form more closely by adding a hip joint and a torso above the massless legs. Poulakakis and Grizzle introduced an asymmetric monopedal (running) SLIP model (ASLIP, denoted with a subscript \u2018A\u2019) that included a torso, with the linear leg actuators terminating at a hip joint (Fig. 1B). The equations of motion of the ASLIP, Eqns. 8-10, are very similar to those of the SLIP but include torso (of mass m, inertia J, length L, oriented at angle \u03b8A relative to the inertial frame) dynamics, which are a critical component for an anthropomorphic gait model. x\u0308A = 1 m fAx (8) y\u0308A = 1 m ( fAy \u2212 mg) (9) \u03b8\u0308A = 1 J (L( fAx sin(\u03b8A) \u2212 fAy cos(\u03b8A)) + \u03c4A) (10) A statics analysis (Fig. 3) of the massless leg can be used to obtain the expressions for the forces and torques applied to the torso",
            " \u03bd\u03b8 = \u2212K\u03b8(\u03b8A \u2212 \u03b80) \u2212 D\u03b8\u03b8\u0307A (20) The hip torque that regulates the orientation of the torso to the error term, \u03bd\u03b8, can be found by substituting Eqn. 19 into Eqn. 10 and solving for \u03c4A. \u03c4A = \u03bd\u03b8J \u2212 L( fAx sin(\u03b8A) \u2212 fAy cos(\u03b8A)) (21) After substituting in Eqns. 14-15 for the single stance phase, Eqn. 21 becomes \u03c4A = lA \u03bd\u03b8J + LpA cos(\u03b1A) lA \u2212 L sin(\u03b1A) . (22) where all of the above subscript A terms refer to quantities associated with the monopedal \u2014 equivalent to single stance because the model has massless legs \u2014 ASLIP model (Fig. 1 B). After performing a coordinate transformation, Poulakakis and Grizzle arrived at a control law for the leg force, pA, that renders the dynamic equations of the ASLIP identical to the SLIP when \u03bd\u03b8 = 0. pA = lA \u2212 L sin(\u03b1A) rA pS (23) The nonlinear coordinate transformation required to derive Eqn. 23 is quite involved because both the SLIP and ASLIP models are underactuated (they have fewer actuators than degrees of freedom). Refer to Poulakakis\u2019s thesis [21] and Ch. 4 of Isidori [17] for details"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003252_j.triboint.2012.07.002-Figure4-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003252_j.triboint.2012.07.002-Figure4-1.png",
        "caption": "Fig. 4. Friction measurement test apparatus.",
        "texts": [
            " The aims of this study were therefore to extend the previous work by examining: friction forces across a range of normal forces that would cover those used when catching/carrying (relatively high) and passing/manipulation (lower); the effect of wear of the surface textures (for a materials containing a relatively high and low amounts of natural rubber); the effect of water and mud in the contact; the effect of wearing mitts of different designs (using the palm rather than finger pad and comparing with bare palm). A bespoke finger friction rig (see Fig. 4) was used for the friction tests. This has been used in a number of previous studies [2,10,16]. The rig consists of two Vishay Tedea\u2013Huntleigh Model 614 50 kg f rated load cells, one to measure the normal force applied and the other to measure the friction force. The load cells are accurate for a force applied along either their upper or lower face, but not for offcentre loading. Therefore, all friction tests were conducted linearly along the centre of the top plate, for the length of the load cell"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001664_s1672-6529(08)60110-9-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001664_s1672-6529(08)60110-9-Figure1-1.png",
        "caption": "Fig. 1 Cantilever model of seta-setule.",
        "texts": [
            " Hence, the elasticity and dimensions of the unique spatulae level seem not able to bend and adhere in a good manner on no-flat surfaces, result that can be reasonably achieved with upper hierarchical levels. To complete the characteristics and dimensions of the model, the evaluations of the stiffness and the adhesive force have to be made. An equivalent cantilever beam model with a round section is considered to compute the stiffness of the spatula (gecko) and the setule (spider), as shown in Fig. 1. The applied force F is considered perpendicular to the substrate and can be decomposed into F\u00b7cos( ) and F\u00b7sin( ) responsible of the bending ( b) and compressive ( c) deformation 3 cos sin, 3b c Fl Fl EI AE , (3) where l is the length, I the moment of inertia ( R4/4), R the radius, A the cross-sectional area ( R2). The total stress becomes 2 3 2sin cossin cos 3n c b Fl Fl AE EI (4) The stiffness of a single element becomes 2 2 2 2 2 4 cotsin (1 ) 3 R Ek ll R (5) Considering a Young modulus of 2 GPa (materials such as keratin have Young modulus on the order of a few GPa) the stiffness of the setules-spatulae can be calculated"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003263_s11012-013-9830-8-Figure3-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003263_s11012-013-9830-8-Figure3-1.png",
        "caption": "Fig. 3 Structure of planetary gear unit for bucket wheel excavator drive",
        "texts": [
            " In order to achieve these goals, it is necessary that the decomposition process includes the following: (a) decomposition of TS structure into components which have to be replaced in complete in the maintenance process, (b) creation of decomposition model of the total desired reliability that will contain the elementary reliabilities of components, reliabilities of component relations and reliabilities of secondary functions not in directly related with the TS components, and (c) creation of analytic model which can provide calculation (processing) of TS desired total reliability. The main goal of decomposition process is to extract elementary reliabilities of TS components from the total desired reliability. The case study presented here for reliability based design is a planetary gear drive unit (planetary reducer) applied for bucket wheel drive at the mine bucket wheel excavator with power 375 kW, trans- mission ratio 182, input speed of rotation 1480 rpm and output speed of 8.15 rpm (Fig. 3a). The gear unit consists of three sections. The first one (I) is an input section that contains one planetary gear set and bevel gear pair (Fig. 3b). Sections II and III are two planetary gear sets i.e. planetary stages in power transmission for torque and speed transformation. One planetary gear set (Fig. 3c) consists of the central pinion, three gear satellites and one inside toothed gear ring. This ring is fixed and by rotation of the central pinion, planetary gears rotate together with the satellite carrier. Input shaft is connected to the central pinion and output shaft with the satellite carrier. In section I the design structure is different and presents the combination of the bevel gear pair and planetary gear set with the input shafts. The main power input (375 kW) provides bevel gear pair with a corresponding (side) shaft",
            " For this purpose, it is necessary to extract elementary reliabilities of components from other various influences on the overall reliability of TS. In this sense, the model of total reliability decomposition has to include reliabilities against appearance (a) malfunctions of design components, (b) malfunctions of connections and relations between design components and, finally, (c) appearance of malfunctions of secondary technical processes such as lubrication, cooling, control etc. In the case of planetary gear unit presented in Fig. 3, the three sections in design structure are identified. Each of them contains a corresponding gear set, bearings, gaskets, satellite carriers and various connecting fits. The model, in the form of reliability tree, obtained by total reliability decomposition contains total reliabilities of design components Ra , connection fits (spline joints, bolted joints, interference-clarence joints etc.) Rb and secondary processes in the unit (lubrication, cooling, . . .) Rc (Fig. 4). Total reliability of design components Ra is structured in three branches of reliability tree, each for every structure sections I, II and III",
            " The damage that occurs first causes the replacement of a complete component. This is the reason why the reliability for design of a component is equal to elementary reliability in relation to failure with minimal elementary reliability. In the case of a planetary gear set, it is elementary reliability against teeth flanks failure of planetary central pinion. In the model in Fig. 4 these are RPGS-II and RPGS-III for sections II and III. In the case of section I, the bevel gear pair and the planetary gear set (Fig. 3b) create one gear set complete, i.e. one component. Failure of teeth flanks of the bevel gear pair or planetary gear set will cause replacement of the complete set. In the model in Fig. 4, reliability of input gear set RIGS is equal to the lower value of elementary reliability of teeth flanks of the bevel or planetary set. In general, gear teeth can be broken due to bending and similar stresses. In the case of involutes\u2019 gears this can be accidental or the result of errors in the production. This is the reason why this possibility has to be included in the part of total reliability Rc in Fig",
            " 4, but not in the reliability for design of a gear set. 3.4 Elementary reliability of bearings and gaskets In planetary gear transmission unit bearings are the executors of auxiliary functions, but they are also very important components from reliability aspect. The main function of power transmission is carried out by gear sets and this process is supported by auxiliary functions of bearings. Every planetary gear set contains two groups of bearings: one group is applied to support satellite carrier (marked by BC in Fig. 3a) and the second one to support satellite gears, marked by BG. Bearings of satellite carriers BC operate similarly to all bearings of common shafts in the gear units. Additionally, as the forces of planetary gear sets are in a relative balance, these bearings are not significantly loaded. Unlike the previous group, the bearings of satellite gears BG are highly loaded and settled in a small space inside the satellite gears. High load and small space (of relatively small bearing dimensions) indicate that reliability of these bearings is not enough",
            " 4 shows, the branch of reliability for design model for the section I contains elementary reliabilities RBV 1 \u00b7 \u00b7 \u00b7RBV 4, then elementary reliabilities RBG1-I \u00b7 \u00b7 \u00b7RBG4-I, and one RBG-I which represents elementary reliability of all three bearings in satellite gears. Elementary reliabilities of bearings in satellite gears in all three sections have great influence on the total reliability of planetary gear transmission unit. The components with short exploitation life and those that have to be frequently replaced are gaskets which seal the contact between shafts (input and output) and housing. Damage of a gasket can produce damage of gears and bearings. This probability has to be included in an analytic model for decomposition. In Fig. 3, the two input shafts contain two seals in section I of the planetary gear unit, marked by SE. In section III there is one seal (SE) included in the reliability system (Fig. 4). These elementary reliabilities RSE are linked to the number of shaft revolutions in the course of service life only. The objective of reliability based design is to provide the design parameters (DP) of components which will satisfy desired elementary reliability obtained by total reliability decomposition. In this way, a set of additional objectives can be also satisfied",
            " planetary gear sets. According to previous discussion, the central pinion in the planetary gear set is exposed to higher teeth stress with higher stress cycles number, compared to other gears in the set. This is the reason why elementary reliability, associated with the central pinion flank failure, presents elementary reliability of the planetary gear set assembly. As a case for this elementary reliability calculation, the central pinion of planetary gear set in section III of gear drive unit in Fig. 3a, is selected. In Fig. 2 and by Eq. (1) the calculation process based on operation load spectrum (load probability) and gear failure probability, is presented. Operation load spectrum of bucket wheel excavator, where gear drive unit is applied, is presented in Fig. 6. This machine operates in a regime, where resistance torque at the bucket wheel is not uniform and varies its value. The level of this torque depends of the material resistance, operating intensity, handling way etc. The load spectrum presented in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003311_j.jelechem.2011.02.020-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003311_j.jelechem.2011.02.020-Figure1-1.png",
        "caption": "Fig. 1. Dimensionless electrode system modelled by simulation.",
        "texts": [
            " Material is considered to diffuse within in the simulation space, where migratory [7] and convection effects are negligible, by way of Fick\u2019s second law to solve the electrolysis current as a function of time: @c @t \u00bc DCO 2C; C \u00bc \u00bdA ; \u00bdB \u00f02\u00de for typical voltammetric and chronoamperometric measurements, where C is the concentration of species A or B, DC is the diffusion coefficient of the species of C, and t is time. The finite difference system used in this work is similar to that found in Refs. [14,61] which have been used previously to great effect. Parameters are noted in Table 1 and dimensionless parameter transforms are noted in Table 2. Due to system symmetry, only half the system need be modelled with the appropriate boundary conditions, as shown in Fig. 1, where 2d is the shortest dimensionless separation distance from the generator electrode edge to the collector electrode edge. Also, a dimensionless diffusion coefficient ratio DR is defined as the ratio of the diffusion coefficients between species B and A, and thus Eq. (3) applies, and the time ordinate transform is relative to DA. DR \u00bc DB DA \u00f03\u00de Potential-step chronoamperometry was simulated at the electrode surfaces by solving the dimensionless three-dimensional form of Eq. (2), as shown in Eq"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000041_j.mechmachtheory.2006.05.006-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000041_j.mechmachtheory.2006.05.006-Figure2-1.png",
        "caption": "Fig. 2. Definition of coordinate systems.",
        "texts": [
            " From the dynamics point of view, the LAP together with the spindle attached to the moving platform can be considered as a system of interconnected rigid bodies, and the generalized principle of D\u2019Alembert states that for such systems the total virtual work done by the effective forces along any kinematically admissible virtual displacement and rotation is zero [17], or dW a \u00fe dW e \u00bc 0; \u00f01\u00de where dWa is the total virtual work done by the gravity forces, inertia forces, and inertia moments, and dWe is that done by the externally applied forces and moments. As shown in Fig. 2, the origin of the coordinate system U\u2013V\u2013W attached to the moving platform is denoted as point P, which is located at the center of mass of the moving platform. The spindle is rigidly attached to the moving platform and its axis is coincident with the W axis of the moving frame. The position vector of the center of mass of the spindle is given by Ps \u00bc P\u00fe qw; \u00f02\u00de where q is the distance between the two mass centers and w is a unit vector which denotes the positive direction of the W axis. Consequently, the virtual displacement of the center of mass of the spindle and its virtual rotation can be, respectively, written as dPs \u00bc dP\u00fe dW qw \u00f03\u00de and dWs \u00bc dW; \u00f04\u00de where dP is the virtual displacement of the mass center of the moving platform, and dW is a 3 \u00b7 1 vector representing the virtual rotation of the moving platform [20]",
            " In addition, the orientation and the angular velocity and acceleration of the moving platform at any time instance can be, respectively, written as \u00bdRp\u00f0/\u00de \u00bc \u00bdR\u00f0w;/\u00de \u00bdR0p ; \u00f017\u00de xp \u00bc x0p \u00fe _/w; \u00f018\u00de ap \u00bc a0p \u00fe x0p _/w\u00fe \u20ac/w; \u00f019\u00de where [R(w,/)] is a 3 \u00b7 3 spatial rotation matrix [20], in which / is the rotation angle of the moving platform about the w axis, and \u00bdR0p \u00bc \u00bdu j v j w is a 3 \u00b7 3 orthogonal matrix which represents the given orientation of the moving platform. x0p and a0p are, respectively, the specified angular velocity and acceleration of the moving platform, and _/ and \u20ac/ are, respectively, the angular speed and angular acceleration of the moving platform about the W axis. Noting that the linear acceleration of the mass center of the tool bit can be written as (see Fig. 2) as \u00bc ap \u00fe as qw\u00fe xs \u00f0xs qw\u00de; \u00f020\u00de where ap is the specified linear acceleration of the mass center of the moving platform, and xs \u00bc xp \u00fe _/sw; \u00f021\u00de as \u00bc ap \u00fe xp _/sw; \u00f022\u00de are the angular velocity and the angular acceleration of the tool bit, in which _/s is the constant rotation speed of the spindle. Besides, according to the rotation axis theorem, the inertia matrix of the tool bit with respect to the fixed coordinate system at any instance is given by \u00bdIs \u00bc \u00bdRp\u00f0/\u00de \u00bdsIs \u00bdRp\u00f0/\u00de T; \u00f023\u00de where [sIs] is the constant inertia matrix of the tool bit evaluated with respect to the moving coordinate system"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003159_978-1-4419-9985-6-Figure1.25-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003159_978-1-4419-9985-6-Figure1.25-1.png",
        "caption": "Fig. 1.25 PCB of the final demonstrator",
        "texts": [
            " To show the performance of the F/T-sensor a demonstrator was constructed, that \u2013 Could to be connected to a PCB or a notebook \u2013 Had USB as the interface \u2013 Was based on the PCB according to Fig. 1.24 including a Micro-Controller, an A/D-converter and a memory to store the calibration matrix and to calculate the forces and torques \u2013 Included a manageable housing of aluminum that can be flanged to a robot and that includes an overload protection, see Fig. 1.16 \u2013 Allowed calibration and measuring via a revised graphical user interface (GUI) (see Fig. 1.25) With this hardware and GUI it is possible to switch between measuring the single resistors as shown in Fig. 1.26, to present the calculated forces and torques. Figure 1.25 shows the user interface for measurement and force and torque calculations. The following functionality was implemented: Connection settings: Settings for the USB device of the measurement hardware and the PC. \u201cMeasurement settings\u201d: The number of measurement samples, the amount of values to calculate the arithmetic mean and the delay between the measurements can be decided here.Moreover, the name of and path to the file to store the data in are defined. \u201cResistors\u201d: Here one can choose which of the 24 resistors that will be measured including the resistors for temperature compensation"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003274_ssp.210.26-Figure6-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003274_ssp.210.26-Figure6-1.png",
        "caption": "Figure 6 shows the characteristics of the forces as a function of displacement for gas spring of hydropneumatic strut (citroen BX) for static load. It is preferred that with the increase of static load progression of dynamic characteristics is increased too.",
        "texts": [
            " In hydropneumatic suspension the role of spring elements takes gas spring with constant mass of gas (nitrogen). The role of damper takes a valve controlling the flow of hydraulic fluid in the actuator. Gas and oil are separated by special membrane. Valve consists of a central hole, which enables direct the flow of fluid to and from the sphere causing the damping effect. Additional channels are half-closed by disk spring, arranged around a central hole, forming the desired damping characteristics of the compression and expansion . Fig.6 Hydro pneumatic gas spring characteristic -own research. In the case of the damping characteristics of hydropneumatic column (citroen BX) it was designated in common with the gas spring on the indicator test stand-figure 7 [3,4,5]. Fig.7 Hydro pneumatic damping characteristic-own research. Suspension based on mechanical solutions with passive elements are some compromises in terms of the criteria of safety and comfort, and price. They are cheap and the main disadvantage is the inability to control the characteristics of mechanical spring (some resolution is used an additional gas spring elements or hydraulic actuator and shock absorber with controlled damper characteristics but they are more expensive solutions than hydropneumatic suspension)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002513_3.4984-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002513_3.4984-Figure2-1.png",
        "caption": "Fig. 2 Schematic of two tethered bodies.",
        "texts": [
            " The circular path of F* in N implies the assumptions that 1) the only forces significantly affecting the motion of the vehicle are those exerted by E and 2) the orbit radius R is such a great distance in comparison to the largest dimension of the vehicle that attitude changes of the vehicle and its parts have negligible effect on the motion of F*. Accordingly, the orbital angular rate 12 is constant and equal to 12 = (GM/R*)1'2 (1) where G is the universal gravitational constant. D ow nl oa de d by U B G ie ss en o n Fe br ua ry 1 1, 2 01 5 | h ttp :// ar c. ai aa .o rg | D O I: 1 0. 25 14 /3 .4 98 4 DECEMBER 1968 TWO TETHERED UNSYMMETRICAL EARTH-POINTING BODIES 2283 In Fig. 2, the description of the vehicle is shown in greater detail. Let BI, B^ J33 be the principal axes of B for B* and Ci, C2, C% be the principal axes of C for C* with corresponding moments of inertia /i, 72, h for both B and C. Take QB, the B tether connection point, to be on the axis BI a distance x from B* and take Qc, the C tether connection point, to be on the negative extension of axis Ci a distance x from C*. The tether with a natural length L is considered to be composed of a material with properties analogous to a linear spring and viscous damper in parallel",
            "3 to bring the axes fixed in B into their final orientation. The angles 71, 72, 73 used to describe the orientation of axes C\\, \u20ac2, C$ with respect to axes AI, A2, AS are defined in an identical manner. Taking the distance from V* to B* to be small in comparison to the orbit radius R, one may approximate the resultant of the gravitational forces acting on body B by (F*)grav = -(GM/R*)m{ [I + (3r/fi)-ai]ai - r/R} (3) where m is the mass of B, r, the position vector from B* to r = (4) and the unit vectors ai, bi, Ci, d are as shown in Fig. 2. Likewise, the resultant of the gravitational forces acting on C is Fig. 3 Attitude angles for end body B. approximated by (5) At this point it should be noted that, consistent with the assumption made previously about the circular motion of Y*, the resultant of (Fs)grav and (Fo)sr*v is independent of r. The moments of the gravitational forces about the mass center for each body are to the first approximation and TB = 3(GM/& To = \u20223(GM/Rs)(al X (6) (7) where IB and Ic are the inertia dyadics for bodies B and C, respectively, with respect to their mass centers",
            " The equations of motion for this nine-degree-of-freedom system may now be derived and expressed in the nondimensional first-order form y' = F(y) (8) where y is a 18-dimensional column matrix whose elements are 2/i = ft 2/4 = 7i 2/7 = Bi 2/io = 2/13 = 2/16 = 2/2 = P2 2/5 = 72 2/8 = 62 2/14 = 2/6 = 73 2/9 = d/L 2/i5 = 2/18 = (9) The variables ui, u^ HZ are the measure numbers of the angular velocity of B in N expressed in the unit vector basis bi, b2, bs fixed in B, while u\u00b1, u5} u& are the measure numbers of the angular velocity of C in N using the unit vector basis Ci, c2, c3 fixed in C (see Fig. 2 for unit vectors). Finally 1/7, w8, ^9 are the ai, a2, a3 measure numbers of the velocity of C* with respect to B* in reference frame N. In mathematical notation i = 1,2,3 i = 4,5,6 i = 7,8,9 (10) The prime in Eq. (8) denotes differentiation with respect to the independent variable r which has been introduced to accomplish the nondimensional form and is defined by the relationship = tit (11) As the derivation of the differential equations of motion is a rather lengthy and involved matter and is not of principal interest here, its details are omitted and the equations are tabulated in the Appendix A"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002611_jmes_jour_1969_011_023_02-Figure7-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002611_jmes_jour_1969_011_023_02-Figure7-1.png",
        "caption": "Fig. 7. Roller bearing models c Hydrodynamic forces and surface velocities.",
        "texts": [
            " 6 shows cage slip for a given shaft speed and - 20 40 60 80 100 120 140 160 180 200 220 A P P L I E D L O A D - l b f Shaft speed 2530 rev/min; radial clearance 0.002 in. Fig. 4. Effect of lubricant on cage slip J 0 U R N AL M E C H A N I C A L E N G I N E E R I NG S C I E N C E Vol 11 No 2 1969 at UNIV NEBRASKA LIBRARIES on June 5, 2016jms.sagepub.comDownloaded from lubricant viscosity and U value, subsequently reducing the cage and roller slip. THEORY The roller bearing model analysed by Dowson and Higginson (2) is shown in Fig. 7 and illustrates the force components and surface velocities for a typical roller. For steady-state motion the equilibrium equations for Fig. 7 may be written as X P , , + F i = P x , + F , . . * (1) Y P,,-P,, = m(r+Rl)wC2 . . . (2) Moments FI+FZ=O . . - * (3) Neglecting the inertia loading term equation (2) becomes Y P,2-P,l = 0 Expressions for the hydrodynamic force components per unit length of cylinder from conventional hydrodynamic theory are given by lubricant plotted against load for various radial clearances. Decreasing the radial clearance from 0.002 in to 0.001 in has very little effect on cage slip. It is only when the clearance is reduced to 0",
            " It is now necessary to compute for values of load, speed and viscosity the corresponding coefficients A , B and K. The starting point for such calculations is Reynolds equation, which for a compressible, viscosity-pressure dependent fluid neglecting side leakage can be written in the form: dx (14) where h = x2/2R+ho. Dowson and Whitaker (6) from The relevant fluid properties are calculated after . . aP P I P 0 = I+- 1 +bp at UNIV NEBRASKA LIBRARIES on June 5, 2016jms.sagepub.comDownloaded from For the model shown in Fig. 7 normal and tangential surface force components can be defined as follows, reference (7). Px, = -J\"'pRsin 81 Ode . . (17) PV1 = P ~ , J ' e a p R c o ~ Ode . . (18) 81 . . * (19) . . * (20) P,, = 0 f = ho/h The numerical step-by-step integration of Reynolds equation (14), was based on the method suggested by Dowson and Whitaker (5). The only difficulty in the technique is the location, for a given inlet point, of the outlet boundary where p = dp/dx = 0 at x = xl. An iteration procedure is used where for a given inlet position xz, minimum film thickness, oil viscosity, pressure-dependent fluid property coefficients, speed, and for an initial guessed value at outlet of xl, equation (14) is integrated"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000264_3-540-36224-x_4-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000264_3-540-36224-x_4-Figure2-1.png",
        "caption": "Fig. 2. Vertical view of an omni-directional robotic platform with 6 degrees of freedom and 3 nonholonomic constraints [12,19]. This device is capable of highly accurate positioning, high payloads, and high speed motion. In its fully actuated configuration, the robot is endowed with 6 motors at the three wheels and at the three joints (\u03b21, \u03b22, \u03b23). However, underactuated configurations can arise because of failures or intentional design.",
        "texts": [],
        "surrounding_texts": [
            "The research in Robotics is continuously exploring the design of novel, more reliable and agile systems that can provide more efficient tools in current applications such as factory automation systems, material handling, and autonomous robotic applications, and can make possible their progressive use in areas such as medical and social assistance applications. Mobile Robotics, primarily motivated by the development of tasks in unreachable environments, is giving way to new generations of autonomous robots in its search for new and \u201cbetter adapted\u201d systems of locomotion. For example, traditional wheeled platforms have evolved into articulated devices endowed with various types of wheels and suspension systems that maximize their traction and the robot\u2019s ability to move over rough terrain or even climb obstacles. The types of wheels that are being employed include passive and powered castors, ball-wheels or omni-directional wheels that allow a high accuracy in positioning and yet retain the versatility, flexibility and other properties of wheels. A rich and active literature includes (i) various vehicle designs [38,41,44,46], (ii) the automated guided vehicle \u201cOmniMate\u201d [2], (iii) the roller-walker [15] and other dexterous systems [17] that change their internal shape and constraints in response to the required motion sequence, and (iv) the omni-directional platform in [19]. A. Bicchi, H.I. Christensen, D. Prattichizzo (Eds.): Control Problems in Robotics, STAR 4, pp. 59--74, 2003 Springer-Verlag Berlin Heidelberg 2003 Other types of remotely controlled autonomous vehicles that are increasingly being employed in space, air and underwater applications include submersibles, blimps, helicopters, and other crafts. More often than not they rely on innovative ideas to affect their motion instead of on classic design ideas. For example, in underwater vehicle applications, innovative propulsion systems such as shape changes, internal masses, and momentum wheels are being investigated. Fault tolerance, agility, and maneuverability in low velocity regimes, as in the previous example systems, are some of the desired capabilities. A growing field in Mobile Robotics is that of biomimetics. The idea of this approach is to obtain some of the robustness and adaptability that biological systems have refined through evolution. In particular, biomimetic locomotion studies the periodic movement patterns or gaits that biological systems undergo during locomotion and then takes them as reference for the design of the mechanical counterpart. In other cases, the design of robots without physical counterpart is inspired by similar principles. Robotic locomotion systems include the classic bipeds and multi-legged robots as well as swimming snake-like robots and flying robots. These systems find potential applications in harsh or hazardous environments, such as under deep or shallow water, on rough terrain (with stairs), along vertical walls or pipes and other environments difficult to access for wheeled robots. Specific examples in the literature include hyper-redundant robots [13,16], the snakeboard [32,40], the G-snakes and roller racer models in [26,27], fish robots [23,25], eel robots [21,36], and passive and hopping robots [18,35,42]. All this set of emerging robotic applications have special characteristics that pose new challenges in motion planning. Among them, we highlight: Underactuation. This could be owned to a design choice: nowadays low weight and fewer actuators must perform the task of former more expensive systems. For example, consider a manufacturing environment where robotic devices perform material handling and manipulation tasks: automatic planning algorithms might be able to cope with failures without interrupting the manufacturing process. Another reason why these systems are underactuated is because of an unavoidable limited control authority: in some locomotion systems it is not possible to actuate all the directions of motion. For example, consider a robot operating in a hazardous or remote environment (e.g., aerospace or underwater), an important concern is its ability to operate faced with a component failure, since retrieval or repair is not always possible. Complex dynamics. In these control systems, the drift plays a key role. Dynamic effects must necessarily be taken into account, since kinematic models are no longer available in a wide range of current applications. Examples include lift and drag effects in underwater vehicles, the generation of momentum by means of the coupling of internal shape changes with the environment in the eel robot and the snakeboard, the dynamic stability properties of walking machines and nonholonomic wheeled platforms, etc. Current limitations of motion algorithms. Most of the work on motion planning has relied on assumptions that are no longer valid in the present applications. For example, one of these is that (wheeled) robots are kinematic systems and, therefore, controlled by velocity inputs. This type of models allows one to design a control to reach a desired point and then immediately stop by setting the inputs to zero. This is obviously not the case when dealing with complex dynamic models. Another common assumption is the one of fully actuation that allows to decouple the motion planning problem into path planning (computational geometry) and then tracking. For underactuated systems, this may be not possible because we may be obtaining motions in the path planning stage that the system can not perform in the tracking step because of its dynamic limitations. Furthermore, motion planning and optimization problems for these systems are nonlinear, non-convex problems with exponential complexity in the dimension of the model. These issues have become increasingly important due to the high dimensionality of many current mechanical systems, including flexible structures, compliant manipulators and multibody systems undergoing reconfiguration in space. Benefits that would result from better motion planning algorithms for underactuated systems. From a practical perspective, there are at least two advantages to designing controllers for underactuated robotic manipulators and vehicles. First, a fully actuated system requires more control inputs than an underactuated system, which means there will have to be more devices to generate the necessary forces. The additional controlling devices add to the cost and weight of the system. Finding a way to control an underactuated version of the system would improve the overall performance or reduce the cost. The second practical reason for studying underactuated vehicles is that underactuation provides a backup control technique for a fully actuated system. If a fully actuated system is damaged and a controller for an underactuated system is available, then we may be able to recover gracefully from the failure. The underactuated controller may be able to salvage a system that would otherwise be uncontrollable."
        ]
    },
    {
        "image_filename": "designv11_12_0003713_978-1-4614-2419-2-Figure16.8-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003713_978-1-4614-2419-2-Figure16.8-1.png",
        "caption": "Fig. 16.8 Mode shapes from preliminary testing",
        "texts": [
            " The experiments are completed in 12 batches (roving the accelerometer location and/or direction each round). Hammer excitation is conducted in the same location for all 12 rounds, and the hammer impact location is shown in Fig. 16.7. Based upon analysis and findings from the preliminary hammer test, an initial full hammer test is completed with at hammer impact location #1 (Fig. 16.7). A full set of measurements for all 88 degree of freedom is collected three separate times. Using Reflex software from B&K, natural frequencies (see Table 16.1) and mode shapes less than 55 Hz (see Fig. 16.8) are extracted from the experimental data. Rational fraction polynomial parameter estimation techniques are used to create the stability diagram used for mode selection (see Fig. 16.9). Based upon analysis of the stability diagram, Modes 1, 2, 7, 10, and 11 are identifiable. The standard deviation for all three undamaged tests is near 0.025 for all extracted natural frequencies. FE model calibration will be completed taking this experimental uncertainty into account. One must also note that the experimental uncertainty provides a limiting state for the desirable agreement for the test-analysis correlation (i"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002187_gt2010-22086-Figure11-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002187_gt2010-22086-Figure11-1.png",
        "caption": "Figure 11-Early hybrid foil-magnetic bearing concept that nests the foil bearing inside electromagnetic coils to create a \u201csmart\u201d bearing [22].",
        "texts": [
            " Notwithstanding, long-term applications of large foil bearings for heavy rotors could benefit from hydrostatic or electromagnetic load sharing, especially during low speed operation. An early demonstration of such a hybrid approach included the side-by side demonstration of a large (100mm diameter) foilmagnetic bearing [21]. Such concepts have been patented [22, 23] and recent demonstrations indicate that this technology marriage is more capable than either bearing technology alone. Hybridization leads to a \u201csmart\u201d bearing in which the rotor static weight loads can be relieved and active damping can be added via electromagnetics [24]. Figure 11 depicts the hybrid \u2018smart\u201d bearing approach. While externally pressurized hydrodynamic foil and more conventional rigid surface gas bearings are an old concept [25], recent demonstrations show that heavy rotors using large foil bearings (100 mm diameter) can be augmented with integral, pressurized air supply [26]. In this work, as expected, the augmentation air enabled essentially friction free start-up and elimination of foil surface wear. For this particular hybrid bearing which was not optimized for stability, at higher speeds, the hydrostatic bearing component needed to be curtailed to prevent pneumatic hammering and whirl instabilities"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001970_20100915-3-de-3008.00048-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001970_20100915-3-de-3008.00048-Figure2-1.png",
        "caption": "Fig. 2. Coordinate systems.",
        "texts": [],
        "surrounding_texts": [
            "Dynamic Positioning Systems are only concerned with the low-frequency horizontal motions of the vessel, that is, surge, sway and yaw. The motions of the vessels are expressed in two separate coordinate systems (see Fig. inertial system fixed to the Earth, OXYZ; and the other, O\u2019x x2 x6, is a vessel-fixed non-inertial reference frame. The origin for this system is the intersection of the m section with the ship\u2019s longitudinal plane of symmetry. The axes for this system coincide with the principal axes of inertia of the vessel with respect to the origin. The motions along of the axes O\u2019x1, O\u2019x2 e O\u2019x6 are called surge, sway e yaw, respectively. The low frequency motion of a dynamically positioned vessel can be described by: where represents the vector of mid velocities relative to the vessel F F F is the vector of environmental forces (surge and sway) and yaw moment, is the vector of thruster forces and moment and and D are the inertia and 2 by to reduce and , paper in Finally, the 2): one is the 1 id-ship (1) -ship -fixed frame, linear damping matrix, respectively. Such by: 00 0 0 0 ! 0 0 ! 0 0 where M is the vessel mass, M the moment of inertia around the vertical axis, distance between vessel center of mass the O\u2019 position and heading relative to the earth expressed in vector form by: \" # # #$ % & ' The coordinate transformation between the vessel the earth-fixed velocity vectors is accomplished matrix ( ) : \" ( ' ( ' *+,',-.'0 where ' . In order to design the controller, the accelerations must be isolated from (1): ! /0 /0 Rewriting equation (4) in terms of the accelerations and velocities in the OXYZ fixed coordinate system yields the complete model of the system with three degrees of freedom: \"1 2 \", \" 4 \" where 2 \", \" 5!& ' % ' and 4 \" ( ' /0. 3. CONTROLLER"
        ]
    },
    {
        "image_filename": "designv11_12_0002192_6.2009-2047-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002192_6.2009-2047-Figure2-1.png",
        "caption": "Figure 2. Quadrotor reaction torque schematic.",
        "texts": [
            "Unlimited Conference 6 - 9 April 2009, Seattle, Washington AIAA 2009-2047 Copyright \u00a9 2009 by the American Institute of Aeronautics and Astronautics, Inc. All rights reserved. American Institute of Aeronautics and Astronautics 2 II. Quadrotors A quadrotor is a vertical takeoff and landing vehicle (VTOL) that is propelled by 4 rotors. It consists of two sets of counter-rotating fixed-pitch propellers. Both lift and reaction torque of each propeller is used to manipulate altitude, attitude, and heading of the vehicle. Each set of blades is mounted on one axis of the cross shape vehicle, as shown in Figure 2. The altitude of the vehicle is controlled by varying the speed of the four rotors collectively. Assuming that rotor 1 is the front of the vehicle, the pitch can be controlled by adjusting the relative speed between rotors 1 and 3. Similarly, the roll of the vehicle can be controlled by varying the relative speed between rotors 2 and 4. Finally, the yaw of the vehicle is controlled by varying the speed of two rotors on one axis relative to the speed of the rotors on the other, while keeping the collective lift of the vehicle constant"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002680_1.4001838-Figure4-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002680_1.4001838-Figure4-1.png",
        "caption": "Fig. 4 Open loop characteristics of the EPS system",
        "texts": [
            " Figure 3 shows a block diagram of an EPS system from input sw to output c. In this analysis, the small frictional torque Tf is neglected. The region within the dotted line shows a conventional controller consisting of a phase compensator, an assist map, and a current controller. It is assumed that the current is controlled to compensate for errors between the motor torque and the reference assistance torque, and satisfies the equation Ta = Kti 6 The open loop characteristics from sw to c are shown in Fig. 4. When the assist gain is increased to reduce the steering torque, the crossover frequency increases while the phase characteristics remain static. Therefore, the phase margin \u201cC\u201d decreases while the tendency to oscillate increases. 3 Control Strategy to Reduce EPS Oscillation 3.1 Advantages of Pinion Angular Velocity Control. It is not practical to expand the range of the phase compensator to increase the phase margin because it functions like a differentiator and increases noise sensitivity. To examine the relationship between the steering mechanism parameters and the oscillatory behavior, the closed loop transfer function without the phase compensator is considered OCTOBER 2010, Vol"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000101_j.clinbiomech.2006.11.010-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000101_j.clinbiomech.2006.11.010-Figure1-1.png",
        "caption": "Fig. 1. Osseo-ligamentous finite element model of the trunk: (a) postero-anterior view and (b) typical functional unit in the lateral view.",
        "texts": [
            " In this study, the generated FEM represented the geometry of a normal subject (height:147 cm, age:11 years). The latter, although normal, has small variations in vertebral symmetry at the different vertebral levels. As a slight geometrical asymmetry may lead to progressive spinal deformities (Villemure et al., 2004), a perfect spine was computed by averaging coordinate points of the subject and by making the model symmetrical in the sagittal plane. The final FEM contains 1725 nodes and 3593 elements corresponding to a wire-frame representation of the thoracic and lumbar spine, rib cage, and pelvis (Fig. 1a). Vertebrae, intervertebral discs, pelvis, ribs, sternum, costo-chondral cartilages and costo-vertebral ligaments were represented using 3D elastic beam elements. Articular facets were represented using shell and contact elements (Villemure et al., 2002). The mechanical properties of those anatomical structures were collected from experiments and published data (Descrimes et al., 1995). The model was built using Ansys 8.0 finite element package (ANSYS Inc., Canonsburg, USA). Each vertebral body was modeled by 10 3D beam elements interconnected within a rigid crossbar system (16 beam elements) (Fig. 1b). Eight of these 10 beams are distributed along the vertebral edge to enable a representative distribution of variable internal forces within the vertebral body and therefore represent vertebral deformation (wedging) as presented in the literature (Villemure et al., 2002). The eight following muscle groups of the trunk were introduced in the FEM: Multifidus, Iliocostalis, Longissimus, Rectus Abdominis, Obliquus Externus, Obliquus Internus, Quadratus Lumborum, and Psoas. They were represented by 160 fascicles linking the anatomical attachment points in straight lines (Beause\u0301jour, 1999)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003263_s11012-013-9830-8-Figure5-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003263_s11012-013-9830-8-Figure5-1.png",
        "caption": "Fig. 5 Teeth wear distribution in planetary gear set",
        "texts": [
            " All of them have to be included in the model for analytic hierarchy processing of total reliability which has to be developed based on presented model in Fig. 4 and which is not the subject of this article. 3.3 Elementary reliability of gear sets Planetary gear units are very compact structures with a very high specific load capacity. The main disadvantage of these structures is non-uniform failure of planetary gear sets, small space for bearings and heating of gear unit caused by small dimensions, which reduces heat radiation. The subject of further discussion is non-uniform failure of gears in a planetary gear set. Figure 5 shows the distribution of teeth flanks failures at gears in planetary gear set. This is the result of laboratory testing of planetary gear units at back-toback testing rig with a permanent torque during testing process. In this structure the central pinion makes almost three meshes (1) in one revolution. Also, the teeth flank of this pinion is exposed to a very high surface stress, which is the result of the small pinion diameter. The result of these conditions is progressive damage of the teeth active flank",
            " This progressive failure process has taken out a hardened (carbonised) layer, and in further operation process the failure continues with acceleration. This is untypical process according to standardised models of gear teeth failure. Satellite gears with one teeth flank are meshing with the central pinion and with another one with the inside toothed ring. The first one is also exposed to the same high surface stress as the central pinion, but teeth mesh frequency (stress cycles number) is much smaller. The picture in Fig. 5 shows that at this flank, pitting started in the middle region of the flank. The opposite teeth flank of satellite gear in the mesh (2) with inside toothed ring is exposed to lesser surface stress, because concave and convex shapes of flanks are in contact. Stress cycles number is less compared to opposite flank. The result is the beginning of micro-pitting of this flank, as shown in Fig. 5. Furthermore, the teeth flanks of inside toothed ring are exposed to less stress cycle\u2019s number and the wear process of the teeth flanks was not so serious at first. It is possible to notice small surface damages that are produced by plastic penetrations of particles from other gears taken by the oil. The situation is that a complete planetary set has to be removed after central pinion failure occurs. For the gear pairs, the rule is that when teeth flanks of one gear get significant, it is not possible to replace just this gear, on the contrary, it is necessary to replace a complete gear pair. In the case of a planetary gear set (Fig. 5), after the teeth flanks failure of central pinion, it is necessary to replace the complete gear set. In general, various damages can occur in one TS component. The damage that occurs first causes the replacement of a complete component. This is the reason why the reliability for design of a component is equal to elementary reliability in relation to failure with minimal elementary reliability. In the case of a planetary gear set, it is elementary reliability against teeth flanks failure of planetary central pinion"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001680_biorob.2010.5626817-Figure4-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001680_biorob.2010.5626817-Figure4-1.png",
        "caption": "Fig. 4. The reachable space of locomotions in the library. (a) Nonholonomic locomotions sampled by curvature from radius 0 (turn on spot) to radius 1 (m), 2 (m), 3 (m) (turn on the left, right) and infinitive radius (walk on a straight line). (b) Holonomic locomotions sampled by tangent direction \u03b8 of trajectory from initial to final configuration. This tangent direction \u03b8 represents sideward steps (\u03b8 are 0 and \u03c0) and diagonal steps (\u03b8 are from \u03c0/3 to 5\u03c0/6).",
        "texts": [
            " We observed that by changing the curvature of the followed path, the three characteristic velocities change either. The curvature for straight walking is infinitive and corresponds to positive tangential velocity, whereas the angular and the lateral ones are null. The curvature for turning on the spot converges to zero and corresponds to positive or negative angular velocity only. Therefore, we sampled nonholonomic locomotions along curvatures from zero to infinitive on the left and the right to obtain the variance of tangential and angular velocities (see figure 4 (a)). Some examples of these recorded locomotions are shown in figure 5 (a). The sideward walking is an holonomic type of locomotion. In this case, the tangential direction of trajectories are considered. This means that the variance of the tangential direction causes the variance of other velocities. For example, a diagonal walking and a sideward walking have different velocities. As a result, we sampled holonomic locomotions from zero to \u03c0 to obtain the variance of lateral and tangential velocities (see figure 4 (b)). Some examples of these recorded locomotions are shown in figure 5 (b). For each of above types of locomotion, we recorded with 3 levels of tangential velocity: low, normal and fast. Straight running and rest (stop) postures were also captured for answering the needs of various scenarios. The subject performed 68 captures (68 locomotion primitives) total to compose the Motion library. 38 markers were tracked. Motion data were recorded as independent markers trajectories (points cloud) using the Cortex software from MotionAnalysis"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001024_09544062jmes574-Figure6-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001024_09544062jmes574-Figure6-1.png",
        "caption": "Fig. 6 General arrangement (GA) of a woodpecker hammer",
        "texts": [
            " With rather more information it would be possible to predict the mechanical behaviour of the animal under a variety of conditions and to see how much of its behaviour is dictated by mechanics. One of the reasons for studying the woodpecker was to derive a design for a lightweight hammer. It was reasoned that the woodpecker is a bird, therefore has to fly and therefore is constructed as light as possible. The mechanism, which has emerged as a result of the model reported here \u2013 momentum transfer from body to head of the woodpecker \u2013 has been used in the design of a novel hammer (Fig. 6 \u2013 designed by Dr G. Whiteley). A rotating crank is connected by means of a rod to the casing, so that the motor plus its mounting oscillates about a central pin. The motion is transferred to the hammerhead by a parallel springs. The constants can be calculated from the woodpecker data. This hammer has a number of advantages over conventional design. It was originally conceived for use in space exploration, where it has no net inertia Proc. IMechE Vol. 221 Part C: J. Mechanical Engineering Science JMES574 \u00a9 IMechE 2007 until it comes into contact with an object, and even then the force delivered can be tuned"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003973_s12206-012-0811-y-Figure7-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003973_s12206-012-0811-y-Figure7-1.png",
        "caption": "Fig. 7. The loads and constraints.",
        "texts": [
            " The three contact pairs are Piston \u2013 Piston pin, Piston pin \u2013 connecting rod and connecting rod \u2013 crank shaft. Material properties required for this analysis are modulus of elasticity, Poisson\u2019s ratio and density. The material properties of various elements of the analysis are listed in Table 1. The next step is to define loads and constraints required for the analysis. To account for an inertia effect like gravity, appropriate values for g (9.81 m/s 2 ) are given. All the degrees of freedom of elements in that portion of the crank shaft which is seated on main bearings have been constrained. Fig. 7 shows the loads and constraints applied on the finite element model. Fig. 8 shows the deformed shape of the assembly. The deformation legend diagram is given (Fig. 9). The maximum de- formation occurs at the piston top surface and it is equal to - 0.04812 mm (along downward direction). Fig. 10 shows the von Mises stress legend diagram. The tolerance allocation of the piston - cylinder assembly is carried out as follows. The assembly parts are listed in Table 2. The associated tolerances T1, T2, T3, T4, T5, T6 and T7 must be determined so that the clearance x\u03a3, between the piston top surface and cylinder top surface must fall within the functionality limits, 0"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002815_6.2010-7639-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002815_6.2010-7639-Figure1-1.png",
        "caption": "Figure 1. Missile Axis Definitions for VLSAM",
        "texts": [
            " Finally, the results of the study are discussed and the following studies are introduced. II. Modeling of the Vertical Launch Surface to Air Missile In this study, a Vertical Launch Surface to Air Missile is conceptually designed to be in the range of the ones presented in the previous literature.\u2020 The VLSAM, analyzed in this paper, is axi\u2013symmetric and has a blunt nose. It is a tail controlled missile and uses both the aerodynamic tail fins and jet vanes. The shape of the missile and the axis definitions are seen in Fig. 1. Also, the main physical parameters of the missile are presented at Table 1. The missile has a slenderness (length/diameter) ratio of 16.5. This value is chosen such that the moment of inertia values are suitable for the desired high maneuverability specifications. Dynamic modeling of a vertical launch surface to air missile is carried out by implementing the well known Newton-Euler equations with rigid body assumption. Two main coordinate systems as the body coordinate system (B) and the earth fixed inertial coordinate system (E) are defined and the equations of motion are derived with respect to them"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003524_j.rcim.2012.02.004-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003524_j.rcim.2012.02.004-Figure2-1.png",
        "caption": "Fig. 2",
        "texts": [
            " To facilitate the calculation and analysis, the following assumptions have been made: (i) ignore the effect from the spread of work piece, which may change the location of contact point on the groove surface; (ii) the work piece is deformed uniformly in the deformation zone; (iii) overlook the effect from thermal expansion, which may affect the dimensions of outgoing work piece; (iv) rolls are rigid bodies, and the rolling load is distributed uniformly along the contact surface. According to the descriptive geometry method proposed by Ragab and Samy [20], in the process of symmetrical rolling, namely the symmetry axis of work piece is superimposed on that of roll gap; the contact area can be presented by At \u00bc ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi A2 x\u00feA2 y q \u00f010\u00de where, Ax and Ay are the project areas of contact in the horizontal (x\u2013y) plane and the vertical (y\u2013z) plane respectively. As shown in Fig. 2, for a given point is (b cosj, a sinj) located on the surface of oval bar, the projected coordinates in the horizontal plane can be calculated from y\u00bc \u00f0asinj z\u00decot\u00f0b=2\u00de z\u00bc bcosj ( \u00f011\u00de where, b is the half minor axis, a is the half major axis, and j is the eccentric angle. Value of z can be calculated from Eq. (2) and value of b can be obtained from b\u00bc cos 1\u00bd\u00f0R0\u00fe0:5hk asinj\u00de=\u00f0R0\u00fe0:5hk asinj\u00de \u00f012\u00de where, R0 is the minimum radius of rolls, hk is the height of groove. The projected area of horizontal plane Ax can be calculated using numerical integration method",
            " 3 shows the outline of round groove drawn superimposed on the cross section of an oval work piece. It is easy to get the intersections of the oval curve and the round groove profile curve, and then the project areas of contact in the vertical plan Ay can be calculated through integral operation for the blank area in the center of Fig. 3. After that the value of contact area can be obtained through solving Eq. (10). Contact area calculation for other roll passes can use the same approach. For a given point C located in the deformation zone, Fig. 2 shows the start point of deformation C1, Fig. 3 gives the end point of deformation C2, thus, the length of contact arc related to point C can be presented by x\u00bc asinj y\u00bc b\u00f0R0\u00fehk=2 z\u00de ( \u00f013\u00de For any point located on the surface of deformation zone, the lengths of contact arc and sliding distance have a leading role in its rate of wear. For example, assuming the load is distributed uniformly along the deformation zone, in the flat rolling process, due to the contact arc length as well as the sliding distance of every deformation point equals those of other deformation points, rolls can achieve a uniform wear, and the useful life of roll can be extended"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001369_j.jsv.2008.09.045-Figure4-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001369_j.jsv.2008.09.045-Figure4-1.png",
        "caption": "Fig. 4. Pulley deformation model.",
        "texts": [
            " It is to be noted that the contact forces exerted on the pulleys would be equal and opposite of the contact forces exerted on the links i.e. kNp \u00bc kN, kTp \u00bc kT. Since it has been observed [2,4,8\u201312,19,20,22\u201324] that the variation of local groove width caused by the elastic deformation of the pulleys significantly influences the thrust ratio and slip behavior of a belt/chain CVT, simple trigonometric functions (as outlined in Refs. [8\u201312,19,20,23,24]) are used to describe the varying pulley groove angle and the local elastic axial deformations of the pulley sheaves. Fig. 4 depicts the model for pulley deformation. The following equations describe pulley deformation effects in Fig. 4: b \u00bc b0 \u00fe D 2 sin z yc \u00fe p 2 . (3) Using this approximation for the sheave-angle deformation, the pitch radius of a chain link in the deformed pulley sheave can be easily computed. The amplitude of the variation D in the pulley groove angle is always much smaller than unity, however, it is not constant during speed-ratio changing phases due to variations in the pulley axial (clamping) forces. Sferra et al. [24] proposed the following approximate correlations for the amplitude variation D and the center of pulley wedge expansion yc in terms of the transmitted torque t, and the chain pitch radius on driver and driven pulleys, r and r0, respectively, D \u00bc 0:00045 r r0 0:55 , yc \u00bc p 3 r r0 \u00fe 23p 180 ",
            " In this figure, Fr and Ft represent the components of the resultant friction force vector Ff between a chain link and the pulley, which act in the plane of the pulley sheave, and N is the normal force between the link and the pulley. It is necessary to quantify the bolt spring force, Fb, in order to derive the contact forces. The bolt force depends on the bolt length lb and stiffness Kb as well as on the local distance z between the pulley\u2019s surfaces. Since the pulley sheaves also bend, additional axial width variation (refer to Fig. 4) affects the bolt force. So, the bolt force, Fb, can be written as Fb \u00bc Kb\u00f0lb z u\u00de 8\u00f0z\u00fe u\u00deplb; 0 8\u00f0z\u00fe u\u00de4lb: ( (5) It is to be noted that the chain link slips in the plane of the pulley sheave. The slip angle, g, defines the plane where the friction force between the chain link and the pulley acts (i.e. g defines the slip direction). It is the angle which the resultant friction force vector, Ff, makes with a unit tangential direction vector to the pulley. So, in order to get the friction force vector, it is crucial to keep track of the relative velocity vector between the chain link and the pulley"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001893_icems.2009.5382875-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001893_icems.2009.5382875-Figure1-1.png",
        "caption": "Fig. 1. The analyzed motor configuration.",
        "texts": [
            " Position has been detected through saturation effects [1], [2], using back-emf [3], voltage equations, state estimators [4], eddy currents spatial forced distribution [5], magnetic anisotropy [6] and magnetic or mechanical saliency [7]. An exhaustive review of the existing techniques is presented in [8]. Most of the presented techniques work only with moving rotor, and it is known that position detection is more difficult to obtain at standstill and low speed. In this paper, a new simple technique for the rotor position detection that works at low speed without any motor modification is introduced. During a previous work [9], a small industrial PM motor (100 W, 3 phases, 10 poles) with non-salient rotor is analyzed (Fig. 1). Surprisingly, the phase self inductance Ls (Fig. 2) and the phase resistance R (Fig. 3) measured at high frequency with a precision impedance analyzer (Agilent 4294A) are position dependent. For both values, there is a minimum when the measured phase is in front of a magnetic pole (north or south) and a maximum when it is in front of a transition between two magnetic poles. It is found that the variations are due to the iron B-H hysteresis characteristic, and more accurately to the B-H local loops followed during AC signal injection (Figs"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003450_09540911311309077-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003450_09540911311309077-Figure1-1.png",
        "caption": "Figure 1 The optical system for curvature measurement",
        "texts": [
            " Alumina plates (96 per cent purity) of 0.62mm thick with the same metallization as that of the silicon wafer on one side were cut into the same size as silicon dummy samples. Test assemblies were constructed by bonding the metalized surface of silicon to the alumina with nano-silver paste. The assemblies were then sintered using the profile suggested by the manufacturer (Lei et al., 2007, 2010). Three different bondline thicknesses of sintered nano-scale Ag paste, as listed in Table I, were prepared for the experiments. Figure 1 shows the optical system used to measure the curvature of the bonded assemblies. The assemblies were placed on a plate that was secured on top of a linear translation stage. Positioning of the stagewas achieved via a steppingmotor controlled by a microcomputer. A HeNe (4mW) laser beam scanned across the polished surface of die (silicon) along the length direction of the samples as it was moved by the translation stage. The position of the reflected laser beam was detected by a linear-position sensitive detector whose output was acquired by the microcomputer",
            " Since the variation in bondline thickness (maximum 50mm) was very small compared to the total thickness (more than 1mm) of the assembly, it was assumed that therewas nodifference in temperaturewithin the assembly for all of the samples. At least three samples for each condition were taken out from the thermal cycling chamber and their curvatures weremeasured at room temperature after 50, 100 and 300 cycles (Figure 4). Finite element analysiswasemployed to simulate thebehaviour of the joint subjected to temperature cycling from 2408C to 1258C by ANSYS. Due to the symmetry, the finite element mesh of a quarter model is shown in Figure 1, which is made up of Solid45 element with a total of 1,076 elements and 1,080 nodes. Also, the lower corner at the substrate (alumina) side was restrained in three-directions to prevent rigid rotation. Thematerial properties of thedie, adhesive and substrateused for the simulation are presented in Tables II and III. SEM images were taken for as-sintered samples and samples after 300 cycles to investigate the impact of temperature cycling on the reliability of sintered silver attachment. In this study, a criterion for the detection limit for observing cracks Note: C2 and C3 of Generalized Garofalo model are highly correlated parameters Evolution of curvature under thermal cycling in sandwich assembly Yunhui Mei, Gang Chen, Xin Li, Guo-Quan Lu and Xu Chen Volume 25 \u00b7 Number 2 \u00b7 2013 \u00b7 107\u2013116 D ow nl oa de d by U ni ve rs ity o f A la ba m a at T us ca lo os a A t 0 5: 15 1 8 Ju ne 2 01 6 (P T ) was defined as when the length of a crack reached bondline thickness"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003602_icmech.2013.6518547-Figure9-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003602_icmech.2013.6518547-Figure9-1.png",
        "caption": "Fig. 9. Composition of parts Fig. 10. Flange cross-section",
        "texts": [],
        "surrounding_texts": [
            "A. Outline of the Artificial Rubber Muscle We initially utilized a straight-fiber-type artificial muscle as a segment and an actuator [6]. The artificial muscle is made of low ammonia natural-rubber-latex (LA-NR.LATEX). It is light, inexpensive, flexible, and highly biocompatible. The materials are formed in a tube, and a micro carbon sheet is inserted in the long-axial direction. The range of this sheet can be extended when the rubber elongates [7]. The actuator is controlled by air pressure. When supplying air to the actuator, the artificial muscle contracts in the axial direction and expands in the radial direction because the rubber expands, but the carbon sheet limits the expansion. B. Bellows-Type Artificial Rubber Muscle We developed a bellows-type artificial rubber muscle. It is more flexible than the straight-fiber type artificial rubber muscle and, therefore we postulate that a robot made of this material can pass through complex pipelines. Both segments and actuator of the robot are represented by bellows-type artificial muscles. Fig. 2 shows the outline of the bellowstype artificial rubber muscle and Fig. 3 shows two sections. Fig. 4 shows the outline of the muscle in the pressureless and pressurized state. Fig. 5 shows the relation between air pressure and contraction/expansion rate. IV. PERISTALTIC CRAWLING ROBOT A. Outline of Peristaltic Crawling Robot Fig. 6 shows the outline of a peristaltic crawling robot. The robot\u2019s specifications are shown in TABLE I. The robot consists of four units and three connectors that constitute a rubber tube over the inner bellows. Each unit is constructed from one artificial rubber muscle over of the each inner bellows. B. Mechanism of Peristaltic Crawling Robot Fig. 7 shows the mechanism of the robot. Each air tube travels through the inner bellows to a bellows-type artificial rubber muscle. When air is supplied to each chamber between the artificial rubber muscle and the bellows, each flange slides on each air tube along the guide in the flange. Only a single point on the air tube is fixed in the flange. Thus, each unit is independently controllable. Cross-section Longitudinal section Air tube (PU) Flange (ABS) Inner bellows (PE) C. Locomotion Control System Fig. 11 shows the robot\u2019s control system. An H8-3052F microcomputer controls the robot. The digital signal output is converted into an analog signal by using a D/A converter (MCT, MCP4920). Then, this signal is input to a proportional solenoid valve (SMC, TV0030-2M). This valve controls the air compressor (JUN AIR, 6-12) pressure, which should be kept proportional to the input voltage. Accordingly, each unit can be supplied with air from the valve. D. Movement Patterns This robot can have the two movement patterns. Fig. 12 shows two movement patterns of the robot. Pattern 7SM is divided into a seven Step Motion. Pattern 2-1-1 (\u201ccontracted units\u201d - \u201ctransmitted waves in the contracted units\u201d - \u201cnumber of waves\u201d) imitates the earthworm. V. DRIVING TESTS OF PERISTALTIC CRAWLING ROBOT A. Straight Pipe Driving Test First, we performed an experiment in a straight pipe using two movement patterns: 2-1-1 and 7SM. In addition, to examine the effect of slimy surfaces on the robot\u2019s motion, we coated a pipe with a lubricant. The viscosity of the lubricant was 10000\u201315000 mPa\u00b7s and the inner diameter of the pipe was 50 mm. Fig. 13 shows the speed of the robot for each movement pattern. We then computed the average pattern speeds. For the 2-1-1 pattern, the speed was 3.87 mm/s in the lubricated pipe and 3.25 mm/s in the pipe without lubricant. For the 7SM pattern, the speed was 4.47 mm/s in the lubricated pipe and 3.56 mm/s in the pipe without lubricant. These results indicate that for two-pattern motion, the robot speed is greater in the lubricated pipe. We attribute this to the reduction in friction between the robot and pipe wall. The friction occurs to allow sagging under the robot\u2019s own weight when the robot extends. In addition, the friction inhibits the robot\u2019s motion when it is extended. This friction is reduced with lubrication, and the speed is therefore greater in the lubricated pipe for each pattern. Fig. 8. Robot parts Fig. 7. Mechanism of the air tube slide robot (4 units) Fig. 19. Vertical pipe Fig. 20. Robot speed in vertical pipe -10 40 90 140 190 0 10 20 30 40 50 60 70 \u6f64\u6ed1\u7121\u3057 \u901f\u5ea62.30 [mm/s] \u6f64\u6ed1\u6709\u308a \u901f\u5ea60.94 [mm/s] Time (s) \u2014 With lubricant \u2015 Without lubricant D is ta nc e (m m ) B. Bent pipe Driving Test Second, we performed an experiment in a bent pipe with or without lubricant using two movement patterns. The inner diameter of the bent pipe was 50 mm, and the outer curvature radius was 90 mm. The bent pipe is shown in Fig. 14. Fig. 15 shows the process of the experiment, and the arrow indicates the robot tip position. Fig. 16 shows the speed of the robot for each movement pattern. . From these results it was observed that the robot could pass through a bent pipe for each movement pattern. With a lubricant, the speed was 2.71 mm/s in 2-1-1 and 3.90 mm/s in 7SM. Without lubricant, the speed was 2.58 mm/s in 2-1-1 and 3.00 mm/s in 7SM. These results indicate that the speed of the robot in a bent pipe is slower than that in a straight pipe. We attribute this to the increased friction in the robot\u2019s extended state, when the robot straightens itself. Therefore, the normal force in the extended state in a bent pipe is greater than in a straight pipe, and the speed decreases as the friction increases. C. Double-Bent Pipe Driving Test Third, we performed an experiment wherein the robot moved through a pipeline constructed from a 200 mm straight pipe connected to two bent pipes. Note that a conventional endoscope cannot pass smoothly through a pipeline with bent pipe connections. The front and rear of the robot will pass simultaneously through each bent pipe because the robot is longer than 200 mm. Therefore, we can confirm whether the robot can pass through successively bent pipes. In this experiment, the robot uses movement pattern 2-1-1. Fig. 17 shows the experimental environment and Fig. 18 shows the robot speed. These results indicate that the robot can pass through two bent pipes connected by a 200 mm straight pipe. If the number of successively bent pipes increases within a short distance, it will be difficult to move a conventional endoscope through the pipe. Thus, this robot can be used in environments such as that shown in Fig. 17, wherein conventional endoscopes cannot operate. D. Vertical Driving Test Next, we performed an experiment to measure the robot speed when it moved through a vertical pipe, with lubricant and without lubricant, with movement pattern 2-1-1. Fig. 19 shows the experimental process and Fig. 20 shows the robot speed in a vertical pipe, with lubricant and without lubricant. From this result, we confirmed that the robot can move through a vertical pipe, with lubricant and without lubricant. We then computed average speeds from Fig. 20. The speed was 0.94 mm/s in the pipe coated with a lubricant and 2.30 mm/s in the pipe without a lubricant. The speed of the robot decreases significantly in a lubricated pipe. We attribute this to the slippage due to the lubricant. Thus, the robot did not achieve a sufficient gripping force. E. Complex Pipeline Driving Test Next, to prevent the robot from slipping in a lubricated pipe, we attached four pieces of friction material to the surface of each robot unit as shown in Fig. 21. We postulated that this should increase the driving power of the robot by increasing the friction force in the expanded state. Next, we performed an experiment wherein the robot moved through a complex lubricated pipeline using movement pattern 2-1-1. Fig. 22 shows the experimental apparatus for the complex pipeline. Fig. 23 shows images taken every 80 s. The arrow indicates the position of the robot tip. To prevent the robot from slipping in the pipeline four pieces of friction material were attached to the surface of each robot unit. We confirmed that the robot could pass through the entire length of the complex pipeline. Thus, we postulate that this robot is suitable to carry out sewer pipe inspections. VI. DRIVING TEST OF PIPELINE IN WATER Then, we performed an experiment wherein the robot moved through a pipeline in water. In this experiment, we computed the average speeds for a straight pipe and confirmed that the robot could pass through a bent pipe in the vertical direction in water. A. Straight Pipe in Water driving Test We sunk the pipeline in a 0.1 m deep tank and placed the robot in a straight pipe in the water. We measured the speed of the robot three times for each of the patterns 2-1-1 and 7SM. Fig. 24 shows the process of the robot driving in a straight pipe in water while Fig. 25 shows the distance as a function of time for each movement pattern. For the 2-1-1 pattern, the speed was 3.12 mm/s and for the 7SM pattern, it was 2.14 mm/s. These results indicate that the pattern 2-1-1 movement is faster than that of pattern 7sm. In addition, the speed of the robot with pattern 7SM decreases by up to 60 % compared to the speed without the water. We believe this is because for pattern 7SM all units are briefly extended, which is not the case for pattern 2-1-1. The robot is subjected to a force in the opposite direction by the water pressure in the pipeline when the robot moves forward pushing the water. Therefore, when all units extend, the robot is pushed back by the water pressure. The travel speed of pattern 2-1-1 is approximately the same as the speed without the water. This is because for pattern 2-1-1 the robot grasps the inner pipe with at least one of the four units. Therefore, the robot with pattern 2-1-1 is not pushed back by the opposing force. B. Bent Pipe in Water Driving Test Next, to confirm that the robot can pass through a bent pipe in water, we performed an experiment wherein the robot moved through a pipeline composed of a horizontal pipe connected to a vertical bent pipe. Fig. 26 shows the process of the robot driving in the vertical bent pipe in water. The result indicated that the robot can indeed pass through a bent pipe in water. In addition, we confirmed that water was ejected from the top of the inner unit when the robot emerged from the water. We believe this is because high water pressure developed behind the robot due to the water from around the robot being sent backwards by the peristaltic crawling locomotion. This is because peristaltic crawling locomotion functions as a pump. The head unit extended and ejected water while the water pressure behind the robot increased when one or two rear units of the robot expanded. Although the water pressure behind the robot increased, the water pushed backwards passed through the inner robot and ejected from the head unit. Therefore, the water level and water pressure did not rise. VII. CONCLUSION AND FUTURE WORK A. Conclusion A peristaltic crawling robot was developed for inspecting sewer pipes. We performed several experiments with 50 mm inner diameter pipes. We used straight, bent, horizontal, vertical, and in water pipes. In addition, we examined the effect of slimy pipe surfaces on the robot\u2019s motion. The results are listed below. 1. We confirmed the robot\u2019s capability in straight and bent pipes with and without a slimy pipe surface using the 2- 1-1 and 7SM movement patterns. 2. The robot can pass through double-bent pipes. The distance of each bent pipe was shorter than the length of the robot. 3. The robot can pass through vertical pipes with or without slimy surfaces using the 2-1-1 patterns. In addition, the robot movement became more stable by attaching four pieces of friction material to each robot unit. 4. The robot can pass through the entire length of a complex pipeline. 5. The robot can transition between a straight and vertical bent pipe in water. From these results, we believe that this robot is suitable for inspecting sewer pipes. B. Future Work We plan to perform robot driving tests in actual sewer pipes to improve the current experimental data. In addition, we plan to construct an analytical model of the robot and compute its theoretical speed. REFERENCES [1] H. Uchida and K. Ishi, \u201cBasic Research on Crack Detection for Sewer Pipe Inspection Robot Using Image Processing\u201d, Proceeding of the 2009 JSME Conference on Robotics and Mechanics, No. 09-4, 2009. [2] P. Li, S. Ma, B. Li, and Y. Wang, \u201cDevelopment of an Adaptive Mobile Robot for In-pipe Inspection Task,\u201d Proc. IEEE International Conference on Mechatronics and Automation, pp. 3622\u20133627, 2007. [3] T. Okada and T. Sanemori, \u201cMOGER: A Vehicle Study and Realization for In-pipe Inspection Tasks,\u201d IEEE Journal Of Robotics And Automation, Vol. RA-3, NO. 6. December 1987. [4] A. H. Heidari, M. Mehrandezh, R. Paranjape, and H. Najjaran, \u201cDynamic Analysis and Human Analogous Control of a Pipe Crawling Robot,\u201d Proc. IEEE/RSJ International Conference on Intelligent Robots and Systems, pp. 733\u2013740, 2009. [5] H. Sugi, \u201cEvolusion of muscle motion,\u201d The University of Tokyo Press, p. 72, 1977 (in Japanese). [6] T. Nakamura and H. Shinohara, \u201cPosition and Force Control Based on Mathematical Models of Pneumatic Artificial Muscles Reinforced by Straight Glass Fibers,\u201d Proceedings of IEEE International Conference on Robotics and Automation (ICRA 2007), pp. 4361\u2013 4366, 2007. [7] T. Nakamura, N. Saga, and K. Yaegashi, \u201cDevelopment of Pneumatic Artificial Muscle Based on Biomechanical Characteristics,\u201d Proceedings of IEEE International Conference on Industrial Technology (ICIT 2003), pp. 729\u2013734, 2003. [8] Y. Hidaka, M. Yokojima, and T. Nakamura, \u201cPeristaltic Crawling Robot with Artificial Rubber Muscles Attached to Large Intestine Endoscope,\u201d Proceedings of 1st International Conference on Applied Bionics and Biomechanics (ICABB 2010), October 2010. 272 Powered by TCPDF (www.tcpdf.org)"
        ]
    },
    {
        "image_filename": "designv11_12_0000178_isic.2007.4450908-Figure3-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000178_isic.2007.4450908-Figure3-1.png",
        "caption": "Fig. 3. (a) schematic structure of the lower body of humanoid robot. (b) the planes defined by the DOFs of the humanoid lower body.",
        "texts": [
            " During the single support time, the support area is the support foot area; during the double support phase, the support area is the convex area defined by the two feet. By combining the ZMP constraints (Equ.6) and the motion of IPM (Equ.4), the following parameter constraints are acquired    Lxsaxs \u2264 Lfs/2 Lxdaxd \u2264 Lfs + Ls (7) whereLfs is the length of foot print, Ls is step length. III. INVERSE KINEMATICS CONSTRAINTS After 3D planning in the Cartesian space, the angular position of every joint can be obtained through 3D kinematic constraints. The six DOFs of one leg are labeled in Fig.3(a). The six DOFs and three links constitute four planes (Fig.3(b)). The 3D kinematic constraints are described as follows: 1) The shank link (Lsh) and thigh link (Lth) constitute a plane S1. 2) The plane S2 is another plane that passes through DOF1 and parallel to the foot print plane. The intersection between S1 and S2 is always a line Ll. 3) DOF5 and the hip link constitute a plane S3. 4) DOF6 and the hip link constitute a plane S4. In this walking planning, it is assumed that S2 is parallel to S3 and the ground. With the above 3D kinematic constraints, the angles of every DOFs can be obtained in the following ways: DOF1 the plane angle between plane S1and plane S2 DOF2 the angle between shank link and plane S1 DOF4 the angle between thigh link and plane S3 DOF3 the sum of DOF2 and DOF4 DOF5 the plane angle between plane S1and plane S3 DOF6 the plane angle between plane S1and plane S4 Humanoid walking is a periodic motion which alternates between the double-support phase and the single-support phase[10]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003352_imece2013-62193-Figure5-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003352_imece2013-62193-Figure5-1.png",
        "caption": "Fig. 5 The finite element meshed model.",
        "texts": [
            "org/about-asme/terms-of-use ) [ )]( ) ) ) ( ) (7) A 3-D thermal transient model based on the experimental geometry characteristics is developed by ANSYS 11.0. The SOLID-70 element with an 8- node linear brick shape was used in this model, and a non-uniform mesh was adopted for a higher density mesh in the irradiated area with a high thermal gradient. According to the mesh sensitivity study and a stable molten pool, the time increment (\u0394t) of 0.0015s was selected through a number of small time steps. The finest mesh size was 0.15 mm for the deposited area, and 1 mm was selected for the coarsest mesh size in the substrate far from the clad (Fig. 5). Thermo-kinetic model and hardness prediction In order to estimate the hardness distribution and solid-state phase transformation, the temperature evolution calculated from the thermal model was combined with the thermo-kinetic relations. The phase transformation in a typical steel-based deposit by LC yields to changing the carbon solubility at different temperatures and thermal loading cycles. During the LC process, the temperature of the molten pool reaches 100% above the AC3 (i.e., end temperature of austenite transformation)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001960_s12239-009-0053-x-Figure4-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001960_s12239-009-0053-x-Figure4-1.png",
        "caption": "Figure 4. Force components acting on a helical gear tooth.",
        "texts": [
            " A multibody dynamic analysis model for Figure 2 was constructed using MSC/ADAMS and is represented in Figure 3. This model is based on the following three assumptions: (1) shafts and gear teeth are flexible, and bearings are considered to be bushings with 6 degrees of freedom; (2) gear meshing stiffness due to bending varies along a moving contact point between two helical teeth; and (3) fluctuating torque or acceleration is transmitted through the clutch input to the transmission input shaft. 2.1. Bending Stiffness of a Gear Tooth Figure 4 shows a schematic of three components of forces acting against a helical gear tooth. The tangential component is also a transmitted load, which results in a transmitted torque. The important tangential force has the following relationships: (1) (2) where r is the gear radius; \u03b8 is the tooth rotational angle (bending slope) due to bending at a moving contact point during one rotation; the subscripts g and p are a gear and a pinion, respectively; Ft is the transmitted tangential load (N); n is speed (rpm); T is torque (Nm); and V is pitch line velocity (m/s)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003713_978-1-4614-2419-2-Figure29.5-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003713_978-1-4614-2419-2-Figure29.5-1.png",
        "caption": "Fig. 29.5 Eight damage scenarios (damage area marked in blue) (color figure online)",
        "texts": [
            " All damage scenarios represent cracks, which were modeled by a local reduction of the Young\u2019s modulus of the material. Although this is not the most realistic way to model a crack, it is considered a sufficient way for a preliminary evaluation of the performance of the scenario-based identification technique. As the performance of the algorithm is tested with simulated \u2018test data\u2019 the way the damage is modeled does not have an impact on the results (which would certainly not be the case when real test data is used). Figure 29.5 provides an overview of the considered damage scenarios; the blue regions are the areas where the stiffness is reduced for that particular damage scenario. Figure 29.6 shows the impact of two arbitrarily chosen damage scenarios on the resonant frequencies of the structure. This figure illustrates that the impact of every damage scenario results in a very particular \u2018fingerprint\u2019. Recognizing one of these patterns in the frequency differences of the structure in its current and undamaged state indicates the presence of that particular damage scenario"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001465_robot.2010.5509554-Figure3-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001465_robot.2010.5509554-Figure3-1.png",
        "caption": "Figure 3. Definition of the optimization variables",
        "texts": [
            " The second stage is to proceed with a linear motion along the secondary trajectory to approach the final planar coordinates. Once the wheelchair reaches its final position, the third stage will be to orient the wheelchair to its final desired orientation. Figure 2 shows the three stages implemented for the secondary trajectory. be followed by the wheelchair The three stages to be applied for the secondary trajectory will only involve the position \u201cX\u201d and orientation \u201c \u03d5 \u201c variables of the wheelchair. As shown in figure 3, knowing the initial and final transformations of the wheelchair base, the trajectory angle \u03b1 can be defined as: [ ]f i f ia tan 2 (y y ) , (x x )\u03b1 = \u2212 \u2212 (10) That defines the amount of motion needed for the three stages to be followed in the following order: 1) Rotation by the amount of 1 i\u03b2 = \u03b1 \u2212 \u03d5 2) Translation by the amount of 2 2 f i f itr (x x ) (y y )= \u2212 + \u2212 3) Rotation by the amount of 2 f\u03b2 = \u03d5 \u2212 \u03b1 The above three wheelchair motion values can be utilized in the weight matrix as criteria to enforce the wheelchair motion"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003501_2041302510394742-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003501_2041302510394742-Figure1-1.png",
        "caption": "Fig. 1 SkyTote",
        "texts": [
            " The T-Wing is a VTOL unmanned air vehicle that is capable of both wing-borne horizontal flight and propeller-borne vertical mode flight including hover and descent. Prominently, Air Force Research Lab (AFRL) in USA has designed a novel unmanned aircraft called SkyTote [20\u201322]. The objective of this effort is to combine the VTOL and hover capabilities of helicopters with the high-speed cruise capability of a fixed-wing aircraft. The mission phases of this aircraft are hover, takeoff and landing, and transition to forward flight from hover (see Fig. 1). SkyTote is a novel approach to precision delivery of cargo within a limited region [19]. It was originally conceived as an airborne conveyor belt that would use a VTOL capability to minimize ground handling. The concept demonstrator is a \u2018tail-sitter\u2019 configuration and utilizes co-axial, counter-rotating rotors. A relatively large cruciform tail provides directional control in the airplane mode as well as serving as landing gear in the helicopter mode. Transition between hovering flight and wing-borne flight will be the main focus of the initial flight test series"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002510_epepemc.2010.5606557-Figure5-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002510_epepemc.2010.5606557-Figure5-1.png",
        "caption": "Fig. 5. Cross section of the used RSM",
        "texts": [
            " \u2206isHF = ( iss \u2212 i\u0302 s s ) \u2212 \u2206isFM (39) = es prd \u2212 \u2206isFM (40) S1-12 And according to equation (37) it can be forecasted using the estimated angle of a PLL structur. \u2206isHF = Y\u2206S (\u03b8pll)u s L\u2206t (41) Assuming that the rotor position would be estimated correctly by the PLL both vectors (39) and (41) would have the same orientation. If the estimated angle is not correct the orientations will differ which can be found in the result of a vector product and fed back as an error epll to close the PLL. epll = ( iss \u2212 i\u0302 s s \u2212 \u2206isFM )T J S(\u03b8pll) u s L (42) The entire linear tracking scheme is shown in Fig. 4. Fig. 5 shows a cross section of the used RSM. This transverse laminated motor has two pole pairs and distributed stator windings. The rotor is characterized by a large saliency caused by the flux barriers: the d-axis is in the direction of maximum permanence and the q-axis is aligned crossing the flux barriers. For very low currents the magnitudes of the flux in d- and in q-axis are nearly the same (see the measured results in Fig. 6). In order to make the saliency measurable in the stator currents (to make the saliency \"visible\"), it is necessary to always apply a small q-axis current of iq0 = 0"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000802_1.2908921-Figure9-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000802_1.2908921-Figure9-1.png",
        "caption": "Fig. 9 Vertical displacement of the",
        "texts": [
            " The whirl of the rotor is undesirable because, during the whirling motion, the rotor revolves with high frequencies. This motion produces high loads that may destroy the rotor and the bearing 17 . Two types of whirling motions are presented in the literature, namely, cylindrical and conical modes 4 . The difference between the cylindrical mode and the conical mode is demonstrated in Fig. 8. In Ref. 4 , Fumagalli explained that for the rotor under examination in this study, the cylindrical motion causes five times higher retainer bearing forces than the conical motion does. Figure 9 presents a vertical displacement of the rotor as a function of time at the locations of the retainer bearings in example g . In the figure, it can be seen that the cylindrical mode of the whirling has occurred due to the misalignment of the retainer bearings. From the simulated scenarios, it can be noticed that the horizontal misalignment causes the cylindrical whirling motion of the rotor. In the case of the horizontal misalignment, the first contact of the rotor with the misaligned retainer bearing will produce a force component in the rotor transverse direction",
            " This, in turn, leads to backward whirling motion at both ends of the rotor. Finally, friction forces compel the Table 5 Misalignments of the retainer bearings in the simulations Case Retainer bearing 1 Retainer bearing 2 X-direction Y-direction X-direction Y-direction a 0 m \u2212180 m 0 m 180 m b 0 m \u2212190 m 0 m 190 m c 0 m \u2212210 m 0 m 210 m d 0 m \u2212220 m 0 m 220 m e \u2212180 m 0 m 180 m 0 m f \u2212190 m 0 m 190 m 0 m g \u2212210 m 0 m 210 m 0 m h \u2212220 m 0 m 220 m 0 m rotor into a cylindrical whirling motion see Fig. 9 . In the case of Transactions of the ASME ata/journals/jotre9/28757/ on 02/06/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use t t t B s t t ring J Downloaded Fr he vertically misaligned retainer bearing see Fig. 10 b , the ransverse contact force component is negligible. As a result, in he simulations, the behavior of the rotor remains oscillatory. ased on the above presented results, it can be concluded that the ufficiently high friction force effecting on the rotor can lead to he whirling motion of the rotor"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003067_s0022-0728(68)80275-x-FigureI-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003067_s0022-0728(68)80275-x-FigureI-1.png",
        "caption": "Fig. I. Rotated ring-disk electrode. (S), brass shaft, top to chuck of rota tor motor; (DW), wiper arm for disk electrical contact (typically mult i -s t rand copper wire) ; (I), insulation layer of electrical tape; (SL), brass sleeve connected via wire W to ring electrode; (RW), wiper arm for ring electrical contact (typically mult i -s trand copper wire) ; OAr), copper wire connecting sleeve to ring; (C), channel to back-surface of ring; (E), plastic body of R R D E ; (D), disk cut-out, packed with carbon paste; (R), ring cut-out, packed with carbon paste.",
        "texts": [
            " The RRDE had electroactive surfaces of carbon paste. I t was constructed in the following manner. An elec- trode \"blank\" was molded using a self-setting resin (Quick-Mount, Fulton Metallurgical Products Corp.) around an axial I/4-in. brass rod. After the resin had set, the blank was machined on a lathe into a cylinder. The central disk portion was then drilled out (radius of the disk=o.22 cm). Electrical contact to the carbon-paste disk was via the central brass rod and a wiper contact, DW, as seen in Fig. I. The ring portion was cut, j . Electroanal. Chem., 16 (1968) 41-46 with a cutting tool, into the electrode to a depth of ca. 1/16 in. A channel, C, was then drilled from the top of the cylinder down through the back-side of the ring. A copper wire contact, W, from the ring, passed through the channel and up to the top of the cylinder. It made contact to a brass sleeve, SL, insulated from the main shaft, S, via a simple electrical tape wrapping marked I in Fig. I. Electrical contact was then made via tile ring wiper arm, RW, which contacted the brass sleeve. The radius from the center of the disk to the inner boundary of the ring was 0,29 cm. The radius from the center of the disk to the outer circumference of the ring was 0.44 cm. Obviously, the gap between the disk and ring is not particularly narrow but this type of electrode is relatively easy to fabricate. The electroactive surfaces were packed with carbon paste in the usual fashion ~5. Care must be taken that excess paste on the plastic surface does not cause a conducting path between disk and ring. The RRDE of Fig. I is immersed to a depth no greater than about \u00bd the height of the plastic cylinder. The rate constants for the hydrolysis of the quinoneimines of p-aminophenol (PAP) and 3-methyl-p-aminophenol (3-MePAP) are easily measured by reverse current chronopotentiometry 16. These electrode reactions correspond to eqns. (I) and (2) and were used to calibrate the RRDE. The PAP and 3-MePAP were purified by sublimation and recrystallization. Triply-distilled water was used for all solutions. The disk electrode was held at a constant potential on the limiting current plateau for the oxidation of the aminophenols"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002488_085106-Figure9-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002488_085106-Figure9-1.png",
        "caption": "Figure 9. Design chart of the threaded gauge.",
        "texts": [
            " ljL max, ljR max, ljL min and ljR min respectively indicate the maximum and minimum distances between the ball track center and those collected points at left and right sides of the j th gothic-arc type ball track; TMin and TMax respectively indicate the minimum and maximum distances among all ball tracks; Tp is the ball track cross-section error, which is normally compared with the permissible error to determine screw grade. The proposed thread profile measuring system is verified by the specific gothic-arc threaded gauge (produced by Chinese Hanjiang Machine Tool Co., Ltd), as shown in figure 9. The sampling frequency is set to be 200 Hz, rotation speed of feed motor is 12 rev min\u22121 and the total number of sampling points is 85 006. The rotary angle of the light curtain has been calibrated and the error should be compensated in order to minimize the fitting error. According to the measurement data of the thread profile, the radius and the ball center position can be calculated through the proposed algorithm. Comparing with the design value of the threaded gauge, the measurement results of the three experiments are confirmed in the range of tolerance"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003887_j.procir.2015.06.103-Figure4-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003887_j.procir.2015.06.103-Figure4-1.png",
        "caption": "Fig. 4. The meshing coordinate system between the crown and external spiral bevel gears",
        "texts": [
            " The meshing between the external and internal spiral bevel gears in the nutation drive can be considered as the external and internal spiral bevel gear meshing with crown gear (an imaginary gear), which has a pitch cone angle of 90 , and the pitch cone is at right angle to its axis. The pitch cone angle of the internal bevel gear is larger than 90 , and the pitch cone angle of the external bevel gear is less than 90 . The motion between the crown gear and the external or internal bevel gear can be considered as the pure rolling of two pitch cones. To establish the meshing function of spiral bevel gears, five coordinate systems are established in Fig. 4 and Fig. 5 to describe the meshing between the crown and external bevel gear and the meshing between the crown and internal bevel gear. This paper takes the meshing between crown gear and external spiral bevel gear as an example. The screw 1$ and 2$ represent instantaneous screw of spiral bevel gear and crown gear that relative to the frame, respectively, and screw 3$ presents relative spiral of 1$ to 2$ . The fixed coordinate system ),,( 0000 kjiS represents the original location of the external and internal spiral bevel gears attached coordinate system ),,( kkkkS kji )2,1(k in Fig. 4 and Fig. 5. The coordinate system ),,( 0000 kjiS fixed to the crown gear, represents the original location of the crown gear rotatable system. The coordinate system ),,( 1111 kjiS is attached to the external spiral bevel gear and rotates about the axis 1k of the centre of the external spiral bevel gear by angular speed 1\u03c9 in the Fig. 5, with 1 represents the rotational angle of the external spiral bevel gear. The coordinate system ),,( 2222 kjiS is attached to the internal spiral bevel gear and rotates about the axis 2k of the centre of the external spiral bevel gear by angular speed 2\u03c9 in the Fig",
            " The angular velocity of the crown gear is c and that of the spiral bevel gear is k with a unit value, which then has the relationship k k c k c cki sin (12) Where )2,1(kk is the pitch cone angle of the spiral bevel gear. According to the theory of instantaneous screw motion of rigid body, the relative velocity of the crown gear to the spiral bevel gear is ckOAckckOAckck hh \u03c9r\u03c9SSrv 3333 (13) Where ),,( cccOA zyxr represents the vector of the origin of coordinates to the meshing point. In the coordinate system ),,( ccccS kji of Fig. 4, there exists ccc kji\u03c9 coscossinsinsin 221 ccc kk\u03c9 11 sin Where 190 represents the angle between the screw 1$ and 2$ . And the relative angular velocity of crown gear with respect to external spiral bevel gear can be obtained as cccc ji\u03c9\u03c9\u03c9 212111 coscossincos ccc ccc OAc zyx 0coscossincos 21211 kji r\u03c9 (14) Thus, combining equations (13) and (14), the relative velocity of crown gear with respect to external spiral bevel gear is )cossin(cos sincos coscos 221 21 21 1 cc c c c xy z z v (15) Similarly, the relative velocity of crown gear with internal spiral bevel gear can be obtained as cccc ji\u03c9\u03c9\u03c9 222222 coscossincos )cossin(cos sincos coscos 222 22 22 2 cc c c c xy z z v (16) Then, according to the common normal of crown gear and equations (15) and (16), the meshing function can be obtained. In Fig. 4, the matrix transformation from coordinate system 2S to coordinate system 1S can be obtained, and the details of the each transformation matrix are given as follows. The transformation matrix from cS to 0S is 1000 0100 00cossin 00sincos 22 22 0 cM (17) The transformation matrix from 0S to 0S is 1000 0sincos0 0cossin0 0001 11 110 0M (18) The transformation matrix from 0S to 1S is 1000 0100 00cossin 00sincos 11 11 1 0M (19) The transformation matrix from cS to 1S is 1000 0 0 0 333231 232221 131211 00 0 1 0 1 aaa aaa aaa cc MMMM (20) Where 2112111 sinsinsincoscosa 2112112 cossinsinsincosa 1113 cossina 2112121 sinsincoscossina 2112122 cossincossinsina 1123 coscosa 2131 sincosa 2132 coscosa 133 sina (21) The parameter 1 represents the pitch cone angle of external spiral bevel gear"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000546_pes.2007.386135-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000546_pes.2007.386135-Figure1-1.png",
        "caption": "Fig. 1. Full-pitch winding distribution.",
        "texts": [
            " The advantages of short-pitch winding arrangement are found on the basis of computed results and discussion. In order to compare between the full-pitch and short-pitch winding distributions in the stator, at first, an example of the full-pitch winding distribution is shown below. In the bearingless motors with the stator core, in which the n-pole motor windings are arranged at the full-pitch, the suspension force is generated by the additional (n+2)- or (n-2)-pole suspension windings in the stator. Fig. 1 shows an example of the full-pitch arrangement in the 3-phase bearingless motor. Here the number of stator slots is 12 and the numbers of poles of motor and suspension windings are four and two, respectively. The 4u+ and 4u- are 4-pole, U-phase motor windings, the 4v+ and 4v- are 4-pole, V-phase motor T 1-4244-1298-6/07/$25.00 \u00a92007 IEEE. windings, and the 4w+ and 4w- are 4-pole, W-phase motor windings, respectively. The 2u+, 2u-, 2v+, 2v-, 2w+ and 2w- are 2-pole, 3-phase suspension windings, respectively",
            " The voltage is induced by the PM field excitation in each suspension winding; however, it is successfully canceled. Because both of the amplitudes are equal, and the sign is opposite in the induced voltages of suspension windings where wound on two stator teeth in both sides across the rotor as described in Fig. 3. The rotational torque and suspension force are computed by FEM using a machine model with short-pitch winding distribution as shown in Fig. 2. They are also computed for a machine model with full-pitch winding as shown in Fig. 1, and the computed results are compared and discussed between full-pitch and short-pitch winding distributions. The rotor structures are cylindrical as described in Fig. 3. The numbers of PM poles are four and eight in the full-pitch and short-pitch winding models, respectively. These PMs are pre-magnetized to be sinusoidal. The specifications of the full-pitch winding and short-pitch winding models are same except for the number of PM poles, and it is summarized in Table I. In the computation by FEM, the motor torque current and the suspension winding current are commanded so that the rotational torque is generated in the counter-clockwise direction and the suspension force is generated in the xpositive direction"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002929_j.elecom.2010.08.010-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002929_j.elecom.2010.08.010-Figure2-1.png",
        "caption": "Fig. 2. Voltammetric responses of solid electrodes (area: 0.8 mm2) Used substrates: (A) Smooth gold and (B) Gold covered with Pd0. (average thickness: 16 nm). Electrolyte: AN+TBABF4. Reference: SCE. Scan rate: 50 mV s\u22121. (C): Responses after addition of 2- iodopropane (concentration: 8 mmol L\u22121) at 1) glassy carbon, 2) smooth silver, and 3) Au\u2013Pd electrodes. The Pd coverage thickness increases according to the arrow: 16, 23, and 55 nm.",
        "texts": [
            " The Au\u2013Pd macro-electrode used for coulometric measurements was built from a thin sheet of commercial gold (Aldrich, area: 4 cm2) galvanostatically palladized by means of PdCl2 (total amount of electricity from 1 to 2.5 C cm\u22122). Coulometric measurements and electrolyses of alkyl halides reported in this work were carried out using three-electrode cells with a total catholyte volume of about 5 to 10 mL. A fritted glass separated the two compartments. The experiments were completed on small substrate amounts (typically from 0.1 to 0.3 mmol). Efficient argon bubbling was employed in all cases. Fig. 1, SEM image A, shows the structural aspect of a thin palladium deposit onto gold. Fig. 2 exhibits the available potential range when Au (curve A) and Au\u2013Pd (curve B) are tested as electrode materials in the absence of organic substrate in AN+TBABF4. The cathodic limit corresponds at bothmetals to thehydrogendischarge due to the presence of residual water. Using in situ activated neutral alumina can extend the limit towards more negative potentials [1]. The anodic limit with Au\u2013Pd (approximately+1.3 V vs. SCE) presumably corresponds to the oxidation of residual water. A noticeable improvement towards other palladized substrates (such as Cu\u2013Pd, Ni\u2013Pd, and Ag\u2013Pd alloys) is the presence of an available range between 0 V and +1 V",
            "Moreover, the presence of a small gold plate covered by palladium placed in an electrochemical cell and without electric connection seems to react with the RI in solution and produces (possibly via cathodic discharge of a soluble organo-metallic transient) gold deposits at electrified substrates like carbon. The simultaneous presence of Au (as substrate) and palladium (as traces for triggering the catalytic system) could create a synergy. Thus, growing amounts of deposited Pd0 were found to progressively accelerate the catalytic process (see Fig. 2, curves C3) and support reaction 1. Presumably, the reduction potential of the organo-gold intermediate (reaction 3) could explain the exceptional energy shifts towards pure palladium and silver. It is important to note that neither smooth gold nor smooth palladiumpermits reductions of RIs at a potential N\u22121 V vs. SCE. In particular, no step was observed in conditions of Figs. 2 and 3, with 1-iodobutane at gold. RX + Au\u2013Pd \u2192 R\u2013PdII ;X\u2212 h i\u2260 + Au0 \u00f01\u00de R\u2013PdII ;X\u2212 h i\u2260 + Au\u2013Pd\u2192slow R\u2013Au complex form\u00bd \u2260 + Pd0 \u00f02\u00de R\u2013Au complex form\u00bd \u2260 + e\u2212\u2192 Rd \u2193 coupling or grafting + Au0 \u00f03\u00de Doping of smooth gold by thin layers of palladium permits proposing a new family of efficient electrocatalytic materials"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002187_gt2010-22086-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002187_gt2010-22086-Figure2-1.png",
        "caption": "Figure 2- Journal and thrust foil bearings used to control radial and axial shaft motion, respectively.",
        "texts": [
            " Figure 1 shows a sketch of a typical bump style foil journal bearing. This material is declared a work of the U.S. Government and is not subject to copyright protection in the United States. Approved for public release. Distribution is unlimited. Downloaded From: http://proceedings.asmedigitalcollection.asme.org/pdfaccess.ashx?url=/data/conferences/gt2010/70392/ on 02/21/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use 2 There are two distinct types of foil gas bearings, journal and thrust bearings as depicted in Figure 2. Journal bearings support radial loads and thus control rotor orbit. Thrust bearings control axial motion. Though their geometry differs, both types of foil bearings operate under the same basic principles, namely that the moving surface relies on viscous action to drag fluid into the bearing generating hydrodynamic pressure that pushes the inner foil surface away from the shaft. In turbomachinery systems, thrust loads can be minimized through careful sizing and design of aerocomponents and judicious selection of operating points"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003824_jas.2014.7004670-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003824_jas.2014.7004670-Figure1-1.png",
        "caption": "Fig. 1. Aircraft flexible wing model (BMB system).",
        "texts": [
            " System with continuous control actuation will be studied in terms of performance analysis. This approach is applied to control the heave dynamics of an aircraft\u2032s flexible wing model. This application comes under the field of aeroelastic studies, which deals with the interaction of structural, inertial and aerodynamic forces. Current model is inspired from a \u201cbeam-mass-beam (BMB)\u201d system[17\u221218], where two Euler-Bernoulli beams connected to rigid mass represent a flexible wing with fuselage. A schematic of BMB system is shown in Fig. 1. In real world applications, the physical parameters used in the PDE models are often unknown. Online adaptive controllers can be used to estimate these parametric uncertainties and mitigate the loss in performance. Although the theory of adaptive controllers for finite dimensional systems is well developed[19\u221220], for PDE\u2032s adaptive controllers exist only for a few classes of problems. Most of the available adaptive controls schemes are for parabolic type PDE\u2032s. See [21\u221223] for some of the adaptive control techniques proposed for linear parabolic systems",
            " The offline and online operations are as follows: 1) Offline operations a) Train adaptive critic network controller for the system. 2) Online operations a) Measure the state variables at time tk across the spatial dimension. b) Compute the X\u0302k using the basis functions c) Using X\u0302k in the neural network, compute the control u\u0302k. d) Get the desired control profile u (tk, y) from U\u0302k, using the basis functions. A flexible wing aircraft model is represented by using two Euler-Bernoulli beams connected to a rigid mass as shown in Fig. 1. The BMB system primarily represents the heave dynamics of an aircraft model, which is initially assumed to be in a level flight with its wings straight and the lift force balancing the weight. Any perturbation in the wing\u2032s shape (defined by transverse displacement w (t, y)) causes a change in the local angle-of-attack distribution over the wing and this in turn leads to perturbation in lift distribution. The objective is to achieve the level flight conditions (w (t, y) = 0) using the proposed control action"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000894_iros.2008.4651029-Figure10-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000894_iros.2008.4651029-Figure10-1.png",
        "caption": "Fig. 10. The two kinds of elasticity models implemented: on the left the type 1, and on the right the type 2 flexure hinge.",
        "texts": [
            " After having approximated the links to a two concentrated masses model and estimated the inertial moments for the platform along the x and y axis according to the figure 9, the mass matrix elements can be easily defined. Concerning with the Fse vector, the equation (10) must be verified s se siF F F= + (10) with Fsi the column vector of the inertial forces applied to the mobile platform. In order to completely determine the external forces exerted on the platform the last considered contribute is introduced by the material elasticity. If the elastic phenomena are idealized as concentrated into the flexure hinges at the joints, two kinds of elasticity models has been developed as figure 10 shows. For each one of the hinges the compliance has been estimated as expressed in (11), where the table 2 indicates the adopted values, and if considering also the relation (13), with \u2206dis representing the introduced displacement, the Fse vector becomes as reported in (14). Note that (13) expresses in a matrix form the s HJ S \u2202 = \u2202 (4) q HJ Q \u2202 = \u2202 (5) relation between the compliance considered at every hinge and at the platform, while (12) represents the limit condition to neglect the beam compliances"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000388_14644193jmbd97-Figure3-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000388_14644193jmbd97-Figure3-1.png",
        "caption": "Fig. 3 Loci after the displacement due to the preload",
        "texts": [
            " 2 that the loci of the centres of the inner and outer ring raceway groove curvature radii are, respectively, expressed by <i \u00bc 1 2 dm \u00fe rgi db 2 cos\u00f0a0\u00de \u00f07\u00de <o \u00bc <i Bdb cos\u00f0a0\u00de \u00f08\u00de or <o \u00bc dm 2 \u00fe rgi \u00fe db 2 cos\u00f0a0\u00de \u00f09\u00de If the outer ring of the bearing is considered fixed in space as load is applied to the bearing, then the inner ring will be displaced. It can be shown that the distance between the centres of curvature of the JMBD97 # IMechE 2008 Proc. IMechE Vol. 222 Part K: J. Multi-body Dynamics at UNIV PRINCE EDWARD ISLAND on August 5, 2015pik.sagepub.comDownloaded from inner and outer ring raceway grooves at ith ball before and after the preload is equal to the contact deflection and can be found from Fig. 3 and equation (4), where Fig. 4 is a simplified two-dimensional representation of the view normal to a plane at the angle ui from the x-axis. d0 \u00bc Bdb cos\u00f0a0\u00de cos\u00f0aP\u00de 1 \u00f010\u00de The preloaded contact angle was obtained from the following formula aP \u00bc tan 1 Bdb sin\u00f0a0\u00de \u00fe z0 Bdb cos\u00f0a0\u00de \u00f011\u00de The actual deflections caused by the vibrations in the x, y and z directions will be dealt with from this point onwards. Figure 5 describes the situation after this three-dimensional movement of the inner ring centre has taken place"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002686_amm.86.348-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002686_amm.86.348-Figure1-1.png",
        "caption": "Fig. 1 Layout of the Design Unit back-to-back gear test rig.",
        "texts": [],
        "surrounding_texts": [
            "Gear test results showed that, for the as ground gears, surface damage started from initial micropitting and progressive micropitting, followed by the occurrence of macropitting at 32 million cycles and severe scuffing before the end of the test. As for the superfinished gears, only mild scratching and initial micropitting was seen after running for 50 million cycles. Photos showing the sequence of contact fatigue failures on both as-ground and superfinished gears are presented in Fig. 2. Patches of micropitting were seen appeared on gear flanks at an early stage of running and these developed continuously throughout the test. Although micropitting occurred mostly in the gear dedendum, there was a also a considerable amount of micropitting found at the tooth addendum. Gear profile measurement results showed that the progressive occurrence of micropitting caused gear profile deviation at both the dedendum and addendum. The plot of maximum profile deviation against running cycles is shown in Fig. 3, which suggests that as-ground gears had 16 to 20 \u00b5m of maximum profile deviation before the appearance of large scale macropits at 32 million cycles. On the other hand, superfinished gears did not have a measurable amount of profile deviation before 32 million cycles and only had 2 \u00b5m of maximum profile deviation after running for 50 million cycles. Small scale macropits with a size of between 0.1 to 1 mm in diameter were seen to begin appearing along the pitch line position of the as-ground gear flanks after micropitting had resulted in a considerable amount of profile deviation at both sides of the pitch line. Larger macropits appeared later at close to pitch line position, but a little distance away as shown in Figure 2. The maximum depth of a large macropit was found to be between 200 to 300 \u00b5m in this study."
        ]
    },
    {
        "image_filename": "designv11_12_0003713_978-1-4614-2419-2-Figure7.1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003713_978-1-4614-2419-2-Figure7.1-1.png",
        "caption": "Fig. 7.1 CAD model (a) and photo (c) of the structure used to validate the magnetic model, in figure (b) a detail of the chucks used to lock the magnets is shown",
        "texts": [
            ": clearances between moving and fixed parts, optimal ratio between magnets volumes) on so small prototypes. Moreover some simulation results about different distance between the coils and different preload magnet sizes have to be verified. This led to the decision to make a large scale parametric prototype that makes it possible to fulfill the requirements and verify the scale miniaturization effects. To validate the force-displacement magnetic model used in the simulations, an appropriate test structure has been designed (Fig. 7.1). To perform the test two magnets are placed vertically in repulsive configuration, the bottom one is fixed while the other one is free to slide in vertical direction. The test rig makes it possible to take into account small misalignments between magnets in the vertical force effects. Varying a known load acting on the moving magnet and measuring the corresponding air gap between the magnets, it is possible to reconstruct the force displacement magnetic characteristic. Two different empirical formulas, that express the modulus of magnetic force Fm as a function of the air gap x, have been used to fit the experimental data"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000901_tpwrd.2007.905551-Figure12-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000901_tpwrd.2007.905551-Figure12-1.png",
        "caption": "Fig. 12. Photograph of the instrument.",
        "texts": [],
        "surrounding_texts": [
            "From 1999 to 2004, 76 towers located in Lanzhou and Zhengzhou power utility were detected by the instrument developed by us. These towers are 110 kV or 220 kV towers, and overcast several types."
        ]
    },
    {
        "image_filename": "designv11_12_0002380_oceans-yeosu.2012.6263594-Figure9-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002380_oceans-yeosu.2012.6263594-Figure9-1.png",
        "caption": "Figure 9. Four vectored thrusters of P-SURO II AUV.",
        "texts": [
            " From this point of view, the above assumption is reasonable in practical applications. Simulation result is shown in Fig. 8, from which we can the effects of dropout with less than 20s are somewhat endurable. However, in practice, this kind of effect will be accumulated with time increases. To recovering from the drift, certain position aiding sensors such as range sonar, GPS, USBL, etc. may be needed. IV. TVC IN HORIZONTAL PLANE As mentioned before, P-SURO II AUV has four vectored thrusters mounted on its horizontal plane, see Fig. 9. For the horizontal motion, the thrusters can be controlled as following ),( ),45tan( ),( 12 \u03c8 \u03c8\u03c8 \u03c8 \u0394+= \u2212\u2212\u22c5\u2212= \u0394\u2212= hTFRTBL TFLTFR hTFLTBR fFF FF fFF (27) where 1\u03c8 is vehicle heading, 2\u03c8 is the direction of vehicle movement, and 1\u03c8\u03c8\u03c8 \u2212=\u0394 reference is heading error and )(\u22c5hf is a certain control function. In the case of PD control, it can be taken as \u03c8\u03c8\u03c8 \u0394\u22c5+\u0394\u22c5=\u0394 DPh CCf )( . From (27), we can see that once TFLF is determined, the other thrusters also can be determined according to (27). Here TFLF can be determined according to different control purpose in practice"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002742_icarcv.2010.5707422-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002742_icarcv.2010.5707422-Figure2-1.png",
        "caption": "Fig. 2. The potential field method and its problem description",
        "texts": [
            " The total force applied to each position p is the sum of the attractive force and the repulsive force as Ftot(p) = Fatt(p) + Frep(p), which determines the robot direction and speed for reaching a goal with an obstacle avoidance. As mentioned in the previous section, we consider that a robot moves with constant speed. Thus, the total force Ftot(p) can be re-formulated as Fc tot(p) = Ftot(p) |Ftot(p)| = Fatt(p)+Frep(p) |Fatt(p)+Frep(p)| , where Fc tot is the total force in the condition of constant robot speed. Then, given the speed sr(m/s) and sampling time Ts(s), a robot moves sr \u00b7 Ts with the direction of Fc tot(p) or (Fatt(p) + Frep(p))/(|Fatt(p) + Frep(p)|) every Ts. Fig. 2(a) illustrates the forces Ftot(pr) and Fc tot(pr) onto a robot by addition of the attractive force Fatt(pr) and the repulsive force Frep(pr). Throughout this paper, we call Fc tot a unit total force. Similarly, we call Fc att and Fc rep a unit attractive force and a unit repulsive force. When |Fatt(pr)| < |Frep(pr)|, a robot moves away from a goal and an obstacle until |Fatt(pr)| > |Frep(pr)|. Then the robot moves back toward an obstacle and a goal, and finally oscillates between two positions. The local minima problem condition is summarized as |Fatt(pr)| > |Frep(pr)|, (5) \u03c1(pr, po) > sr \u00b7 Ts, (6) \u2207\u03c1(pr, po) = \u2207\u03c1(pr, pg)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001313_09544100jaero155-Figure3-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001313_09544100jaero155-Figure3-1.png",
        "caption": "Fig. 3 Gravity forces and propulsive force in the vertical plane",
        "texts": [
            " The dynamic model as supplied by the manufacturer has been improved in this study and the DC motors are simulated with respect to the corresponding equations. The TRMS possesses two permanent magnet DC motors; one for the main and the other for the tail propelling. The motors are identical with different mechanical loads. The mathematical model of the main motor, as shown in Fig. 2, is presented in equations (1) to (5). The mathematical model of the remaining parts of the system in vertical plane is described in equations (6) to (8) (see Fig. 3) Uv = Eav + Raviav + Lav diav dt (1) Eav = kav\u03d5v\u03c9v (2) Tev = TLv + Jmr d\u03c9v dt + Bmr\u03c9v (3) Tev = kav\u03d5viav (4) TLv = ktv|\u03c9v|\u03c9v (5) In equation (6) the first term denotes the torque of the propulsive force due to the main rotor, the second Proc. IMechE Vol. 221 Part G: J. Aerospace Engineering JAERO155 \u00a9 IMechE 2007 at UNIV CALGARY LIBRARY on May 25, 2015pig.sagepub.comDownloaded from term refers to the torque of the friction force that covers all viscous, coulomb, and static frictions, and the torque of gravity force is shown in the third term"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001284_s10015-006-0401-0-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001284_s10015-006-0401-0-Figure1-1.png",
        "caption": "Fig. 1. The mechanism of the manipulator",
        "texts": [
            " This proposed search region is used as the initial range of the iteration method, and the region is shifted along the axis of the coordinates of angular displacement to minimize the energy consumption of the motor. The dynamic characteristics of manipulator control based on the above-mentioned trajectory are analyzed theoretically and investigated experimentally. When the mechanism of the manipulator is open-loop, the energy consumed increases with the weight of the motor, which is on the moving link. We therefore take a manipulator whose mechanism is closed-loop. The dynamic equations of the manipulator with two degrees of freedom, as shown in Fig. 1, which is able to move in a vertical plane are as follows: \u03c4 \u03b8 \u03b8 \u03b8 \u03b8 \u03b8= ( ) + ( ) + ( )H C G\u02d9\u0307 \u02d9, (1) where \u03c4 = [\u03c41, \u03c42] T is the torque which acts on link 1 and link 2, \u03b8 = [\u03b81, \u03b82] T is the angular displacement of the link, C(\u03b8, \u03b8 . ) is the torque about the centrifugal force and the Coriolis\u2019 force, and G(\u03b8) is the torque about gravity. The applied voltage of the servomotor is e b b b b f= + + + ( )1 2 3 3 \u02d9 \u02d9\u0307 \u02d9\u03b8 \u03b8 \u03c4 \u03c4 \u03b8sign (2) where b1 = kv + (Ra/kt)Dm, b2 = (Ra/kt)Im, b3 = Ra/kt, ia is the electric current of the armature, Ra is the resistance of the armature, Im is the moment of inertia of the armature, and Dm is the coefficient of viscous damping"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001358_icelmach.2008.4799845-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001358_icelmach.2008.4799845-Figure2-1.png",
        "caption": "Fig. 2: Equivalent vector circuit of SynRM including total iron losses [2, 16] (top), Vector diagram of SynRM in steady state, including the total iron losses [2, 3, 16] (bottom), in the synchronous reference frame. The rotor geometry is the first proposed SynRM by J. K. Kostko in 1923 [1].",
        "texts": [
            " INTRODUCTION The Synchronous Reluctance Machine (SynRM), for its torque production, utilizes the reluctance concept and rotating sinusoidal Magneto Motive Force (MMF), which is produced by the traditional Induction Machine (IM) stator. The reluctance torque concept has a very old history and it can be traced back to before 1900 [1, 2, 3], see Fig. 1. The first theoretical and technological attempt to realize the SynRM was made by J. K. Kostko in 1923 [1], see his proposed SynRM\u2019s rotor geometry in Fig. 2. SynRM under closed-loop control can easily be controlled and operated, due to the completely new possibilities through the development in power electronic based drives. Thus the drawback of this machine under direct operation and supply, especially stability and start up torque can be overcome. One of these simple control possibilities is based on the machine current angle control. Therefore different operating conditions, which are achievable under current control of the machine, are briefly developed and presented using a simple and quite new operating diagram of SynRM",
            " The stator current is responsible for both the magnetization (main field), and the torque production which is trying to reduce the field distortion, this can be done by controlling the current angle of the machine [2]. 978-1-4244-1736-0/08/$25.00 \u00a92008 IEEE 1 The general equations describing a conventional wounded field synchronous machine are known as Park\u2019s equations [2]. Therefore, SynRM can be modeled with those equations as well, but in case of SynRM, the field and damper winding equations have to be eliminated from Park\u2019s equations. Therefore, the resultant SynRM vector equations in the dq-axis (synchronous reference frame) can be written as follows (amplitude invariance), see Fig. 2: ,ssiRev += (1) .\u03c9\u03bb\u03bb j dt d e += (2) The machine vector single line diagram based on these equations is shown in Fig. 2, cR represents the machine total iron losses at the operating point. In (1) \u2013 (2), v is the machine\u2019s terminal voltage vector, \u03bb is the stator flux, sR is the winding resistance, cs iii += is the stator current vector, \u03c9 is the reference frame electrical angular speed and e is the stator electromotive voltage (internal voltage of stator winding). The stator flux \u03bb according to the magnetization current i can be defined as follows, slotting effect is neglected [2]: , ),(0~ 0~),( i iiL iiL iL qdq qdd \u22c5 \u2245\u22c5=\u03bb (3) ",
            " The stator winding is sinusoidally distributed and as a result it can be assumed that the flux harmonics in the air gap contribute only to an additional term in the stator leakage inductance [2, 9]. After magnetizing inductances, the most important parameters of the SynRM are: the machine saliency ratio ),( qd ii\u03be that is defined according to (5), and machine load angle \u03b4 , current angle\u03b8 , torque angle \u03b2 and internal power factor angle i\u03d5 . These angles are interconnected together by (6), see Fig. 2. ),( ),( ),( qdq qdd qd iiL iiL ii =\u03be (5) ,\u03b4\u03b2\u03b8 += i\u03d5\u03b8\u03b4\u03c0 +=+ 2 (6) The SynRM main magnetic and electric parameters, using the machine model based on (4) \u2013 (6), are interconnected together through (7). ( ) \u03b8 \u03be\u03bb \u03bb \u03b4 \u03d5\u03b8 tan 1 tan tan 1 ==== + \u2212 dd qq d q i iL iL (7) The stator electromotive voltage e in steady state, using (4), can be calculated according to (8). ( ) .),(),( ,),(),( 0 dqddqqdq qdqqddss iiiLjiiiL iijiije dt d \u22c5\u22c5+\u22c5\u22c5\u2212\u2245 \u2245\u22c5+\u22c5= \u2245 \u03c9\u03c9 \u03bb\u03bb\u03c9 \u03bb (8) The machine internal power factor (IPF) is defined in Fig. 2. If (6) is introduced in that definition IPF can be evaluated according to (9) by some standard trigonometric manipulation [16]. ( ) ( ) \u2212+=== \u03b8\u03b4\u03c0\u03b2\u03d5 2 cossincos iIPF (9) The electro-magnetic interconnection between \u03b8 and \u03b4 according to (7) can be introduced into (9), this gives (10). ( )( ) .sincottancos 1 \u03b8\u03b8\u03be\u03b8 <\u22c5+\u2212= \u2212IPF (10) An important conclusion from (10) is: IPF is always less than (sin\u03b8 ) at any operating point, see Fig 4. This equation shows that IPF strongly depends on the operating point \u03b8 as well as the machine \u03be ",
            " Introducing interconnection relation between \u03b8 and \u03b4 according to (7) into the torque equations (12) \u2013 (13) gives new equations for torque according to (14) \u2013 (15)[16]. ( ) ( )( )\u03b4\u03be tantan2sin 22 3 12 \u22c5\u22c5\u22c5\u2212= \u2212ILL p T qdag (14) \u22c5 \u22c5 \u2212= \u2212 \u03be \u03b8 \u03c9 tan tan2sin 11 22 3 1 2 E LL p T dq ag (15) If the machine air gap performance according to (8), (10), and (12) or (15) is evaluated for each operating point (I, \u03b8 , E, ) the terminal performance in the steady state can be determined by using (1) and the vector diagram in Fig. 2. The internal power factor based on (9) \u2013 (10) for different \u03be and as function of \u03b8 is shown in Fig. 4 (left). IPF has a maximum according to (16) (MTPkVA). 1 1 tan max + \u2212=\u21d4= \u03be \u03be\u03be\u03b8 MTPkVAor IPF (16) In Fig. 4 (right) based on (11), machine \u03be as a function of \u03b8 and IPF as parameter is shown as well, which shows a SynRM with IPF > 0.9 has to have a \u03be of almost 20. This saliency ratio is not practically feasible [11]. Similar to IPF, the SynRM torque as function of \u03b8 and at constant current, (12), and constant flux, (15), conditions is demonstrated in Fig",
            " The SynRM OpD clearly shows that the machine can provide certain demanded torque 0T , see Fig. 6, at certain speed , by choosing different operation strategies. Points \u201cA\u201d, \u201cG\u201d and \u201cC\u201d represent MTPA, MTPkVA and MTPV respectively. Running the machine in all these conditions produces same torque 0T . Points \u201cD\u201d & \u201cB\u201d are of particular interest with respect to point \u201cA\u201d. All these points have the same stator current, which is equal to the point \u201cA\u201d current, DBA IIII ===0 . The SynRM stator flux in these points can be evaluated using (8), (16) \u2013 (18), SynRM vector diagram in Fig. 2 & OpD in Fig. 6, according to (19) \u2013 (21). ( ) 222 )()( 22 2 022 2 0 2 0 2 0 2 0 222 0 I LL I L I L ILIL E qdqd qAqdAd AA \u22c5+= + = =+= = = \u03bb \u03c9 \u03bb (19) Similarly for points \u201cD\u201d & \u201cB\u201d there are, [16]: ( ) 2 02 0 22 21 2 1 1 2 \u22c5 + =\u22c5= = \u03bb \u03be \u03be\u03c9 \u03bb ILL E qd DD (20) . 21 4 1 1 2 2 2 0 2 2 022 2222 \u22c5 + = =\u22c5 + \u22c5 \u22c5= = \u03bb \u03be \u03be \u03c9 \u03bb I LL LLE qd qdBB (21) On the other hand, torque at these points are according to (22) \u2013 (24) using (12), (15), (16) \u2013 (18), [16], (k = constant). ( ) 0 2 0 TILLkT qdA =\u22c5\u2212\u22c5= (22) ( ) 0 2 0 1 2 1 1 1 2 1 1 TILLkT qdD \u22c5 + = + \u22c5\u22c5\u2212\u22c5= \u03be \u03be \u03be \u03be (23) ( ) 0 2 0 1 2 1 1 1 2 1 1 TILLkT qdB \u22c5 + = + \u22c5\u22c5\u2212\u22c5= \u03be \u03be \u03be \u03be (24) The machine IPF using (10), (16) \u2013 (18) can be calculated as well",
            " This fact can be used to estimate online the machine \u03be and inductances at the operating point, for a similar method see [18] where by using the machine voltage vector equation (1) in dq-frame, instead of terminal values, the inductances are estimated. Solving (7) for \u03be as function of i\u03d5 gives \u03be estimator equation using the machine power factor according to (27). ( ) ( ) ( )( )IPF ii iqd 1costantan tantan, \u2212+\u22c5\u2212= =+\u22c5\u2212= \u03b8\u03b8 \u03d5\u03b8\u03b8\u03be (27) Solving (7) for \u03b4 as function of i\u03d5 gives \u03b4 estimator using the machine power factor according to (28) see Fig.2. ( ) ( ) ( )( ) \u03b8 \u03b8 \u03c9 \u03b8 \u03d5\u03b8 \u03c9 \u03b8 \u03b4 \u03c9\u03b8 \u03b4\u03bb\u03bb sin coscos sin cos sin sin sin sin , 1 IPF I E I E I E ii iiL i q q qdq \u2212+\u2212\u22c5= =+\u2212\u22c5= =\u22c5=\u22c5== (28) The machine dL can be estimated using (5), (27) \u2013 (28). The performance of this machine saturation level estimator is simulated in Simulink-Matlab environment, the results are shown in Fig. 8 that shows in steady state it can estimate the actual machine inductances in less than 1s. This estimator uses the machine terminal values instead of the dq-frame voltage and current, which is the case in [18]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003618_jmes_jour_1970_012_066_02-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003618_jmes_jour_1970_012_066_02-Figure1-1.png",
        "caption": "Fig. 1. Journal bearing system",
        "texts": [
            " 1970 and accepted for publication on 14th M a y 1970. 33 Member of the Institution. J O U R N A L M E C H A N I C A L E N G I N E E R I N G S C I E N C E 4 Attitude angle. w Journal speed, rad/s. Recent experimental work (2) (3) has supported the use of the short-bearing approximation to describe the dynamic behaviour of modern turbine bearings for small amplitudes of journal or rotor motion (I). The short-bearing approximation will be used here. No approximations will be made as to the smallness of the journal motion. Consider a journal (Fig. 1) under the action of an external static load P and an external centrifugal load Pc due to unbalance, Oil-film forces P, and P, are called into play as shown in Fig. 1. Let the journal (or proportion of rotor) mass be m. Then } (1) P,+P cos ++pC cos (wt-+> = mc(i'-&) P, - P sin Q + P, sin (wt - $) = mc(q5 + 2 4 ) where P,, P, are functions of E,,, E , <, (I, and $. c0 is given (1) by 4 l ' ~ ~ / ~ l ~ R ~ = ~ o [ 2 ( 1-coz)+ 1 6 ~ o ~ ] ~ ' ~ / ( 1 - ~ o ~ ) ~ Vol I2 No 6 1970 at UNIV CALIFORNIA SAN DIEGO on March 22, 2016jms.sagepub.comDownloaded from . In non-dimensional form these equations become Y1+Y cos ++, cos (i-+) = E \u201d - I ) \u2018 ~ E Non-dimensional forces ",
            " 2a shows a family of whirl-loci for Y = 1.0 and c0 = 0-7, being typical of a mediumsized turbogenerator rotor. The family is obtained by adopting in turn different values of ujc. The increasing importance of non-linearity effects as ujc is increased can be clearly seen. In all cases the transient oscillations have been allowed to die away. This takes a non-dimensional time of about 12.0. Further evidence of the non-linearity can be obtained from Fig. 3a which shows the displacement in vertical and horizontal directions (as defined in Fig. 1) against non-dimensional time wt. The particular value of u/c used here is 0.8. Also taking place are large oscillations of the oil film about the attitude line (Fig. 1) as shown in Fig. 3b. If a .Y value of 10.0 is chosen then the corresponding family of whlrl-loci is shown in Fig. 26, indicating much reduced orbital amplitude as a proportion of radial clearance. For small high-speed steam turbines, values of B and c0 of about 0.2 and 0.4 respectively are typical and increase in ujc has greater effect (Fig. 2c). It may, however, be noted that non-linear effects are reduced, due primarily to the reduction in c0. Fig. 4 shows the departure from linearity of the vertical waveform as vibration amplitude increases, estimated from the families of curves for which P = 0"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003398_156855112x629531-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003398_156855112x629531-Figure2-1.png",
        "caption": "Figure 2. Alignment of micro-braiding for compression molding. This figure is published in color in the online version.",
        "texts": [
            " Micro-braided yarn was fabricated as fiber volume fraction of 50% with a tubular-braiding machine. In the present study, since bamboo rayon fiber was very fine and was tender, four fibers were bundled with a quiller machine and were used in the micro-braiding process. Fabrication and structure of microbraided yarn are illustrated in Fig. 1. Micro-braided yarn obtained was wound on the metallic frame 50 \u00d7 2 times. Then, the micro-braided yarns wound on the frame were placed on the pre-heated mold die and were molded with a hot-press system. Figure 2 shows a schematic view of the molding process. After a given molding cycle, the heating platens were cooled by the water flow in the pipes equipped through the platens. During cooling, molding pressure was applied until the temperature of the platens became 40\u00b0C and the molded pieces were obtained. The molding piece was a unidirectional composite with 200 mm length and 20 mm width. Here, molding pressure, molding temperature and molding time are defined as the pressure during compression molding, the temperature of the heating platens and the duration of molding temperature, respectively"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000399_s10999-008-9077-z-Figure5-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000399_s10999-008-9077-z-Figure5-1.png",
        "caption": "Fig. 5 (a) Prescribed tangential contact stresses on the boundaries (sub-problem I), and (b) prescribed normal contact stresses together with the actual contact surface CC (sub-problem II)",
        "texts": [
            " In this case, the constitutive relations take the following form: F\u00f0trij; tj\u00de \u00bc 0; d te p ij \u00bc tdk oF o trij \u00f014\u00de trij r dt \u00bc tC e ijrs\u00f0d ters d te p rs\u00de \u00f015\u00de with Ce ijrs being the material elastic constitutive tensor. 3.2 Special sub-problems of VI formulations Currently, there is no direct method available for the solution of the general frictional contact problem of Eq. 9 with all the terms included. In our approach, we divide the general elasto-dynamic variational inequality into two consistent sub-problems. Details are provided below. If the tangential stresses rT are assumed to be known everywhere on the boundaries of the two bodies at time t + Dt, then variational inequality (9) reduces to (Fig. 5a): hq\u20acu; _v _ui \u00fe A\u00f0u; _v _u\u00de R1\u00f0 _v _u\u00de _v 2 K \u00f016\u00de where R1\u00f0 _v\u00de \u00bc R\u00f0 _v\u00de \u00fe Z t\u00feDtCC rT _vd t\u00feDtS \u00f017\u00de The effect of friction is now included in the expression of R1\u00f0 _v\u00de as known tractions over the boundary CC. Note that for the unilateral contact model, the rate of work done by the normal contact stresses hP\u00f0u\u00de; _v _ui is always negative or equal to zero. Therefore, variational inequality Eq. 16 is reduced to: hq\u20acu; _v _ui \u00fe A\u00f0u; _v _u\u00de R1\u00f0 _v _u\u00de \u00f018\u00de In sub-problem II, where the normal stresses rN together with the actual contact surface CC are known (Fig. 5b), the variational inequality (9) reduces to: hq\u20acu; _v _ui \u00fe A\u00f0u; _v _u\u00de \u00fe j\u00f0 _v\u00de j\u00f0 _u\u00de R2\u00f0 _v _u\u00de; _v 2 K \u00f019\u00de where j\u00f0 _v\u00de \u00bc Z t\u00feDtCC t\u00feDtr\u0302N _vTj jd t\u00feDtS \u00f020a\u00de R2\u00f0 _v\u00de \u00bc R\u00f0 _v\u00de \u00fe Z t\u00feDtCC t\u00feDtrN _vNd t\u00feDtS \u00f020b\u00de The above VI formulation (19) has a non-differentiable frictional term j( ) (Eq. 20a). In order to overcome this difficulty, two approaches, regularisationand nondifferentiable optimisation, were adopted (Czekanski and Meguid 2001). 3.3 Modelling of friction The above VI formulation (19) has a non-differentiable frictional term j( ) (Eq"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002649_1.4706582-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002649_1.4706582-Figure1-1.png",
        "caption": "FIG. 1. Principle of cladding using scanned beam.",
        "texts": [
            " The difference is that laser light is guided using scanner instead of the static optics. Scanning optic adds flexibility to cladding process compared to using static optics.4 Flexibility of cladding process is based on possibility to adjust numerically the dimensions of laser interaction area, scanning amplitude. Scanning amplitude of laser beam on the top of work piece is mainly dependent on scanning mirrors tilting angle. When this angle can be adjusted numerically, it means also that laser interaction area can be adjusted numerically.4,10,11 Figure 1 shows the principle of cladding with scanning optics. Some scanners include the feature of power adjustment, which means that laser power can be adjusted according to the target location of laser beam or, in other words, according to tilting angle of scanning mirror.12 This kind of feature increases process flexibility even more when numerically can be adjusted how much laser power is input in certain part of melt pool. This study concerns on this matter. Laser cladding quality is typically graded by dilution ratio, such that the lower the dilution is, the better is the quality"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003713_978-1-4614-2419-2-Figure55.10-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003713_978-1-4614-2419-2-Figure55.10-1.png",
        "caption": "Fig. 55.10 Dynamic grid points",
        "texts": [],
        "surrounding_texts": [
            "To improve the analytical analysis an update of the stiffness distribution of the dynamic model with ground resonance test results is required. In the last two decades many analytical methods have been developed for updating, which can be classified as local and global methods. This methods do not taking into account the structural damping and any non-linearity of the structure. Sensitivity analysis is required to localize the errors in the dynamic modelling. With the results of the sensitivity analysis the model can be checked for robustness about the dynamic analysis and the areas shown of the airframe structure where changes in the model have fundamental effects. Using the above described method of the dynamic model all updating methods run into problems because the dynamic model is not available as finite element model due to the dynamic reduction. To improve the updating process, the quantity and quality of test measurements have to be improved. Methods where non-linearity and structural damping can be considered are required for further investigation."
        ]
    },
    {
        "image_filename": "designv11_12_0003713_978-1-4614-2419-2-Figure49.2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003713_978-1-4614-2419-2-Figure49.2-1.png",
        "caption": "Fig. 49.2 Measurement principle to determine the object coordinates using two known image points",
        "texts": [
            " With a photogrammetric evaluation strategy digital images from an object under load are used to obtain 3D coordinates of previously applied measurement points. White targets with black surrounding are used as measurement point markers. The grayscale deviation between the white dot in the middle and the black surrounding allows the PONTOS software to accurately determine 2D point coordinates in each camera image. Once the software recognizes a measurement points in a pair of captured images, its position in three dimensional space is determined using spatial triangulation (see. Fig. 49.2). H. Berger (*) \u2022 O. Erne \u2022 M. Klein GOM mbH, Mittelweg 7, Braunschweig 38106, Germany e-mail: h.berger@gom.com R. Allemang et al. (eds.), Topics in Modal Analysis II, Volume 6, Conference Proceedings of the Society for Experimental Mechanics Series 31, DOI 10.1007/978-1-4614-2419-2_49, # The Society for Experimental Mechanics, Inc. 2012 481 The 3D coordinates of all object points are calculated from the intersection point arisen from the center point rays from the two observation image points. The main measurement results are 3D coordinates of an unlimited number of measurement point markers"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003761_9781118818428-Figure21.3-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003761_9781118818428-Figure21.3-1.png",
        "caption": "Figure 21.3 Upper: notation for collision prediction sequence. Lower: basic case of wall collision \u2013 wall stationary",
        "texts": [
            " The inertia of the molecule \u2018cloud\u2019 entering the compression (upper) space from the annular gap causes a rarefaction above the upper horizontal face of the displacer. Heavy concentrations in the \u2018cloud\u2019 correspond to the opposite \u2013 elevated density and pressure. Molecules are spheres in perfectly elastic collision. Integration time step \u0394t can be chosen sufficiently short that there is only one collision (moleculeto-molecule, or molecule-to-enclosure) per \u0394t. The method does not cope with simultaneous impact between 3 bodies. Ultimate Lagrange formulation? 239 The upper diagram of Figure 21.3 represents two molecules, j and k, on course to potential collision. At time t = t0 particle j is at xj, yj and travelling with constant velocity components uj, vj, while k is located at xk, yk and travelling at uk, vk. At time t0 the origins are separated by (scalar) distance s0: s0 = \u221a {(xk \u2212 xj) 2 + (yk \u2212 yj) 2} After a small time increment dt separation s is given by: s = \u221a {(xk + ukdt \u2212 xj \u2212 ujdt) 2 + (yk + vkdt \u2212 yj \u2212 vjdt) 2} d, a collision occurs. Equating s to d in terms of coefficients of dt: dt2{(uk \u2212 uj) 2 + (vk \u2212 vj) 2} + 2dt{(xk \u2212 xj) (uk \u2212 uj) + (yk \u2212 yj)(vk \u2212 vj)} + s0 2 \u2212 d2 = 0 240 Stirling Cycle Engines This is a quadratic equation dt of the form adt2 + bdt + c = 0 having the usual solutions (involving the square root): a = (uk \u2212 uj) 2 + (vk \u2212 vj) 2 b = 2{(xk \u2212 xj) (uk \u2212 uj) + (yk \u2212 yj) (vk \u2212 vj)} c = s0 2 \u2212 d2 For pairs of molecules for which the argument of the root is positive, there are two real solutions, one or both of which may be positive. If both are positive, then it is the smaller which is relevant, since the larger relates to the (virtual) instant of separation where distance s passes through the value dm for a second time. Under assumptions to this point, the algebra of wall collisions (lower diagram of Figure 21.3) is elementary. Molecules, identified by index j are scanned in sequence (j= 1, nmol). Each is first examined as to whether current location xj, yj and velocity components uj, vj lead to contact with a wall. If so, point of contact and amount dt remaining out of time interval \u0394t are calculated, and post-contact location xj \u2032, yj \u2032 determined. By hypothesis this particle is not also in collision with another individual molecule during dt, so can be re-positioned. The process repeats until a molecule-to-molecule collision with k is detected, whereupon the collision algebra (below) is applied to give uj \u2032, vj \u2032 and thence xj \u2032, yj \u2032",
            " Using superscript R to denote components rotated to the new system: uR j = |Vj| cos(\u03b8j \u2212 \u03b8); vR j = |Vj| sin(\u03b8j \u2212 \u03b8) Re-introducing prime \u2032 to denote values after impact: uR j \u2032 = 0; uR k \u2032 = uR j; vR j \u2032 = vR j; vR k \u2032 = 0 Post-collision moduli are unchanged by rotation back through the original contact angle \u03b8: |VT j \u2032| = \u221a {02 + vR j \u20322} = vR j \u2032 |VT k \u2032| = \u221a {uR k \u20322 + 02} = uR k \u2032 Respective angles required for re-orientation to the earlier transformed plane are: \u03b8T j \u2032 = atan(vR j \u2032\u2215uR j \u2032) \u03b8T k \u2032 = atan(vR k \u2032\u2215uR k \u2032) Applying the rotations involves use of the original \u03b8: uT j \u2032 = |VT j| cos(\u03b8T j \u2032 + \u03b8); vT j \u2032 = |VT j| sin(\u03b8T j \u2032 + \u03b8) uT k \u2032 = |VT k| cos(\u03b8T k \u2032 + \u03b8); vT k \u2032 = |VT k| sin(\u03b8T k \u2032 + \u03b8) Post-impact velocity components are acquired by reversing the original transformation, that is, by adding pre-impact values of uk and vk: uj \u2032 = uT j \u2032 + uk; vj \u2032 = vT j \u2032 + vk uk \u2032 = uT k \u2032 + uk; vk \u2032 = vT k \u2032 + vk The integration time increment \u0394t during which collision occurs is reckoned from t0. Denoting time of collision by tc, molecules j and k move apart during the balance of \u0394t for length of time time d\u0394t, where d\u0394t = \u0394t \u2212 (tc \u2212 t0). The new locations at the end of the integration step are thus: xj = xc + uj \u2032d\u0394t; yj = yc + vj \u2032d\u0394t xk = xc + uk \u2032d\u0394t; yk = yc + vk \u2032d\u0394t If a framework can be set up for economically tracking, storing and retrieving velocity and location, then individual event descriptions can be dummies, to be embellished later ad lib. In this spirit, Figure 21.3 illustrated a collision with a containing surface: the molecule is 244 Stirling Cycle Engines provisionally assumed to rebound with the x-component of velocity reversed and with the y-component un-altered \u2013 except when the surface is in motion (horizontally at uw or vertically at vw), in which case the appropriate velocity component is increased by uw or vw. Proper functioning of the numerical algorithm should be independent of number of molecules. It can therefore be explored with a small number (some hundreds) of molecules"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002097_09544062jmes1228-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002097_09544062jmes1228-Figure2-1.png",
        "caption": "Fig. 2 The off-line of action contact for a tooth pair in approach (pinion is fixed)",
        "texts": [
            " The starting point of the off-line of action contact should be confirmed in order to determine ta. If the pinion is held stationary, the elastic deformation of preceding tooth pair caused the gear to rotate an angle of \u03c6k. Then, the two gears should be rotated backward to a position at \u03b81 and \u03b82 (\u03b81Rb1 = \u03b82Rb2, Rb1 and Rb2 denote the base radii of pinion and gear, respectively) to make the tooth pair contact at the starting point of the off-line of action contact, which is just the point D in Fig. 2. Theoretical start and end of contact are represented as points A and B in Fig. 2, while \u03b1A and \u03b1D are the pressure angles for the pinion tooth at points A and D, respectively. From Fig. 2, it is shown that the gear should be rotated forward to a position at (\u03b82 + \u03c6k) to make the tooth pair contact from point D to A. If the gear is assumed to be rotated with constant speed ng (units: r/min) in the pre-mature contact region, then ta (units: s) could be obtained as ta = (\u03b82 + \u03c6k)60 ng2\u03c0 (3) Thus, the approaching contact time ta could be determined using equation (3) once \u03b82 and \u03c6k are known. The rotation angle \u03c6k could be gained using \u03c6k = \u03b4a/Rb2, in which \u03b4a denotes the elastic deformation of JMES1228 \u00a9 IMechE 2009 Proc",
            " 223 Part C: J. Mechanical Engineering Science at RICE UNIV on April 13, 2015pic.sagepub.comDownloaded from preceding tooth pair along the line of action. \u03b4a is equal to the deformation of the tooth pair at the end of the single contact zone and could be calculated using the fitting formula [24], here it will not be repeated again. The determination for \u03b82 will be mainly introduced in the following analysis. The coordinates of point D is set to be (xD, yD), and from the geometrical relationships shown in Fig. 2 we have{ xD = Ra2 sin(\u03b32 + \u03c6k + \u03b82) yD = a0 \u2212 Ra2 cos(\u03b32 + \u03c6k + \u03b82) (4) where a0 denotes the center distance of the gear pair. Thus O1D = \u221a x2 D + y2 D (5) The pressure angles \u03b1D and \u03b1A are written as \u03b1D = arccos ( Rb1 O1D ) (6) \u03b1A = arctan \u239b \u239c\u239d (Rb1 + Rb2) tan \u03b1 \u2212 \u221a R2 a2 \u2212 R2 b2 Rb1 \u239e \u239f\u23a0 (7) in which \u03b1 is the pressure angle for the reference circle of the pinion. For \u03b31 and \u03b32, we have \u03b31 = \u03b1 \u2212 \u03b1A (8) \u03b32 = arccos ( Rb2 Ra2 ) \u2212 \u03b1 (9) Thus, an identical equation exists as follows arcsin ( xD O1D ) \u2212 inv\u03b1D \u2212 \u03b31 + inv\u03b1A = Rb2 Rb1 \u03b82 (10) where inv(\u2022) is the involute function",
            "comDownloaded from me equilibrium mass of the gear pair system ng rotational speed for the gear q relative variable part of the periodically time-varying mesh stiffness Ra1, Ra2 radii of the addendum circle for the pinion and gear, respectively Rb1, Rb2 base radii of the pinion and gear, respectively t time instant ta approaching contact time td time for meshing teeth in the lowest point of single tooth contact tr recessing contact time Tg operating torque of the gear Tm mesh cycle x relative displacement across the mesh action line zg teeth number of the gear \u03b1 pressure angle for the reference circle of the pinion \u03b1A, \u03b1D pressure angles for the pinion tooth at points A and D in Fig. 2 \u03b4a deformation of the tooth pair at the end of single contact zone \u03b4pi, \u03b4gi deformations of the single tooth for the pinion and gear under unit load (1 N) \u03b5 contact ratio \u03b81, \u03b82 back rotated angle for the pinion and gear due to the defection of the preceding tooth pair, respectively \u03b8p, \u03b8g rotational vibratory components of the pinion and gear, respectively \u03bb percentage of approaching contact time (or recessing contact time) on the mesh cycle \u00b51, \u00b52 characteristic exponents of the transformed system \u00b5\u2032 1, \u00b5\u2032 2 characteristic exponents of the original system \u03be mesh damping ratio \u03c6k rotation angle for the gear caused by the elastic deformation of the preceding tooth pair (pinion is fixed) + transition matrix for the positive sloping sawtooth waveform \u2212 transition matrix for the negative sloping sawtooth waveform L transition matrix for the lower constant part of rectangular waveform U transition matrix for the upper constant part of rectangular waveform \u03c8(t) varying pattern for the periodically time-varying mesh stiffness \u03c9ave average natural frequency \u03c9m mesh frequency JMES1228 \u00a9 IMechE 2009 Proc"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001369_j.jsv.2008.09.045-Figure3-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001369_j.jsv.2008.09.045-Figure3-1.png",
        "caption": "Fig. 3. Free body diagram of chain link interaction.",
        "texts": [],
        "surrounding_texts": [
            "The chain links in the chain CVT are modeled as kinematically decoupled planar rigid bodies which are connected to each other by force elements. Each chain link represents a rigid body with 3-degrees-of-freedom (dof) in a plane (i.e. (xc, yc, y)): two translations of the center of mass of the link and one rotation about an axis passing through the link\u2019s center of mass. The force elements take into account the elasticity and damping of the links and joints. In addition to these interconnecting forces, the links also experience contact forces in normal and tangential directions whenever they come in contact with the pulley sheaves. The chain in the CVT is modeled link by link to account for its discrete structure. Figs. 2 and 3 illustrate the free-body diagram of a chain link. It is to be noted that in Fig. 2, the dotted arrows represent the forces (fx and fy), which only arise when a chain link comes in contact with the pulley sheave. Using Newton\u2013Euler formulation [8\u201312] and the Theory of Unilateral Contacts [2,21], the equation of motion for all links under unilateral contact conditions with the pulleys can be written as M\u20acq h \u00f0WN \u00feWSl\u0302SjWT\u00de kN kT ! \u00bc 0, where l\u0302S \u00bc f mi\u00f0 _gi\u00de sign\u00f0 _gi\u00deg. (1) kN, kT are the normal and the sticking constraint forces of the links that are in contact with the pulley and _gi is the relative velocity between a link, i, and the pulley in the sliding plane (refer to Fig. 5). WN represents a matrix with coefficients of relative acceleration (in the normal direction) between the links and the pulleys in the configuration space. WS represents a matrix with coefficients of relative acceleration (in the slip direction) between the links and the pulleys in the configuration space when the links are slipping on the pulley sheave, whereas WT represents a matrix with coefficients of relative acceleration (in the slip direction) between the links and the pulleys in the configuration space when the links are sticking to the pulley sheaves. It is to be noted that these relative velocities and accelerations are computed at the points where the links contact the pulleys. ARTICLE IN PRESS N. Srivastava, I. Haque / Journal of Sound and Vibration 321 (2009) 319\u2013341324"
        ]
    },
    {
        "image_filename": "designv11_12_0000802_1.2908921-Figure3-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000802_1.2908921-Figure3-1.png",
        "caption": "Fig. 3 Circle-in-circle contact",
        "texts": [
            "org/about-asme/terms-of-use t t t a b o b a t c w h 0 Downloaded Fr he combined inertia of the bearing. For the ball bearing used in his study, the combined inertia is more than 31% larger compared o a case where only the inertia of the inner ring of the bearing is ccounted for. 2.3 Model of the Contact. A complete model of the contact etween the rotor and the retainer bearings includes descriptions f contact forces and friction. Contact between the rotor and the earing can be modeled using a nonlinear circle-in-circle contact, s depicted in Fig. 3. The radial contact force Fr is a function of he contact penetration and the penetration velocity. The radial ontact force, which affects the rotor, can be written as follows: Fr = K p 1 + 3 2 \u0307 , er cr and Fr 0 0, er cr or Fr 0 20 here K is the stiffness of the contact and is the parameter that as a value between 0.08 and 0.2 for a steel 22 . In Eq. 20 , p is Fig. 4 Diagram of the electric motor millimeters\u2026 21102-4 / Vol. 130, APRIL 2008 om: http://tribology.asmedigitalcollection.asme.org/pdfaccess.ashx?url=/d the parameter that is dependent on the type of contact, er represents the radial displacement of the rotor, cr is the radial clearance between the rotor and the inner ring of the retainer bearing, that is, an air gap, is the depth of penetration of the contact, and \u0307 is its derivative with respect to time. The X- and Y-components of the radial contact force Fr can be calculated using the geometry presented in Fig. 3. Equation 20 is based on the Hertzian contact theory for two spheres and limited impact velocity below 500 mm /s 23 . The penetration between the rotor and the inner ring of the bearing can be expressed as follows: = er \u2212 cr 21 As shown in Fig. 3, the radial displacement between the rotor and the inner ring of the bearing with respect to the center of the bearing can be obtained from the displacements along X- and Y-axes as follows: er = ex,r 2 + ey,r 2 22 The radial clearance in the contact can be obtained using the radii of the rotor rr and the inner ring ri as follows: cr = ri \u2212 rr 23 The magnitude of the friction force, which acts at the center of the rotor and is perpendicular to the radial contact force, can be calculated as follows: F = Fr 24 where is the coefficient of friction between the rotor and the inner ring of the bearing"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002319_j.jsv.2012.05.009-Figure7-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002319_j.jsv.2012.05.009-Figure7-1.png",
        "caption": "Fig. 7. Snapshots of the flapping motion at w\u00bc0.6 with the time interval of 0.25: (a) t\u00bcT0; (b) t\u00bcT0\u00fe0.25; (c) t\u00bcT0\u00fe0.5; (d) t\u00bcT0\u00fe0.75; (e) t\u00bcT0\u00fe1.0; (f) t\u00bcT0\u00fe1.25. (Re, G , rr, l)\u00bc(500, 50, 16, 30), Lz\u00bc2 (BBC).",
        "texts": [
            " 6, the average amplitude of the free end Aa increases with increasing plate width, which is partly caused by the constraint effects of the walls, as discussed later. For (Re, G, rr, l)\u00bc (500, 50,16,30), we can conclude that a relatively narrow plate flaps asymmetrically, whereas a wide plate flaps symmetrically. However, difference between the amplitudes A1 and A2 for w\u00bc0.4 is smaller than that for w\u00bc0.6, as shown in Fig. 6, indicating a very narrow plate should not exhibit apparent asymmetrical flapping motion. Figs. 7\u20139 show the snapshots of the flapping motion with the time interval of 0.25 for w\u00bc0.6, 0.8 and 1.4, respectively. Fig. 7 shows the asymmetrical flapping in the spanwise direction at w\u00bc0.6, and the plate often exhibits the spanwise warping deformation. For w\u00bc0.8 and 1.4, the flapping is symmetrical, but the stronger local deformation can be observed. For w\u00bc0.8, the strong deformation appears at the center region in the spanwise direction, as shown in Fig. 8, whereas for w\u00bc1.4, the deformation mode becomes complex: two spanwise waves can be observed at the free end, but only one spanwise wave is observed in the middle region (Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001575_j.automatica.2010.10.036-Figure4-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001575_j.automatica.2010.10.036-Figure4-1.png",
        "caption": "Fig. 4. A translational oscillator with a rotational actuator (TORA).",
        "texts": [
            " Moreover, consider the compact set Br\u2206 := {Ea \u2208 \u2126r\u2206|L(Ea) \u2264 Lm} as depicted in Fig. 2(c). One can conclude that if an error trajectory starts from a point inside Br\u2206 (i.e. Ea(0) \u2208 Br\u2206 ), then according to the standard Lyapunov theorem extension, the error trajectory Ea(t) is ultimately bounded (Ge & Zhang, 2003; Lewis, Yesildirek, & Liu, 1996; Yesildirek & Lewis, 1995). The block diagram of the closed-loop system is depicted in Fig. 3. A TORAmodel is considered to illustrate the performance of the proposed controllers (Karagiannis et al., 2005; Lee, 2004), see Fig. 4. The system\u2019s dynamics is governed by the following differential equations: (M + m)x\u0308 + ml(\u03b8\u0308 cos \u03b8 \u2212 \u03b8\u03072 sin \u03b8) = \u2212kx (J + ml2)\u03b8\u0308 + ml cos \u03b8 x\u0308 = \u03c4 where \u03b8 is the angle of rotation, x is the translational displacement, and \u03c4 is the control torque. The positive constants k, l, J,M and m denote the spring stiffness, the radius of rotation, the moment of inertia, the mass of the cart, and the eccentric mass, respectively. Define the states and the input variables as \u03b71 = x + ml sin \u03b8/(M + m), \u03b72 = x\u0307 + ml\u03b8\u0307 cos \u03b8/(M + m) z1 = \u03b8, z2 = \u03b8\u0307 , u = \u03c4 "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000399_s10999-008-9077-z-Figure3-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000399_s10999-008-9077-z-Figure3-1.png",
        "caption": "Fig. 3 Two bodies in contact: (a) unilateral contact model, and (b) schematic of kinematic contact condition",
        "texts": [
            " Unfortunately, the resulting solution algorithm suffers from the same disadvantages as those outlined in the traditional penalty approach. A few other publications have devoted attention to the practical implementation of variational inequalities in static contact problems (see e.g., Bogomolny 1984; Bischoff 1984; Qin and He 1995; El-Abbasi 1999). The current research study overcomes these difficulties by developing and implementing a new variational inequalities methodology to treat dynamic frictional contact in engineering structures. In the unilateral (Signorini-type) problem, the contact constraints are shown in Fig. 3, and can be expressed as follows: before/after contact uN g\\0) rN \u00bc 0 & _uN is unconstrained \u00f01a\u00de during contact uN g \u00bc 0) rN 0 & _uN 0 \u00f01b\u00de These contact constraints state that: (i) the magnitude of the normal contact stress is less than or equal to zero, and (ii) the displacement of the contacting surfaces must not allow for interpenetration. Note that no constants are required to describe the contacting surfaces in the normal direction. Coulomb\u2019s friction condition can be expressed as follows: uN\\g) rT \u00bc 0 \u00f02\u00de uN\u00bcg) jrTj lrN jrTj\\ lrN) _uT\u00bc0; jrTj\u00bc lrN)9k 0; _uT\u00bc krT: 8 < : \u00f03\u00de where l is the coefficient of friction and is assumed to be independent of velocity"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001512_s11668-010-9398-8-Figure4-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001512_s11668-010-9398-8-Figure4-1.png",
        "caption": "Fig. 4 First kind of gear tooth configuration: (a) geometrical dimension, (b) solid model, and (c) cooling model",
        "texts": [
            " There is one ratchet-coupling zone in the vicinity of the pitch diameter of the gears, which indicates that the maximum tension existed at that location. In the previous operational test, the maximum temperature was in the vicinity of theFig. 3 Standard spur gears pitch diameter because of the effects of the ratcheting pressure. The configuration used in this study attempted to reduce the accumulation of heat at the surface of the tooth. Therefore, a 3-mm hole was drilled at the pitch diameter, which was perpendicular to the tooth axis and went through the width of the tooth. The holes shown in Fig. 4 were drilled at an angle oriented along the gear axis. Fresh, cooling air flowed from one surface to the other as the gear rotated at a constant speed. The size of the holes that were drilled in the vicinity of the pitch diameter affected the quantity of air mass and also reduced the mechanical properties of the gear. However, the holes were carefully positioned to minimize the effect of the holes on mechanical behavior. Second Type of Configuration for the Reduction of the Plastic Gear Tooth Surface Temperature Figure 5 shows the geometry of the gears with two different cooling holes",
            " The material properties of the tooth became worse because of the continuously rising temperatures, and the gear did not perform adequately for the loadings and was damaged. Owing to the repeated loading and unloading of the gear tooth, severe gear tooth deflections occurred, and a large amount of heat was generated because of internal friction. An increase in the tooth load also contributes to an increase in the gear temperature. First Type of Gear Tooth Configuration The first type of gear tooth configuration was created such that the hole diminishes the temperature of the tooth surface, and the heat generally accumulates in the pitch region (Fig. 4). The thermal damage shown on the first gear tooth generally occurred as a result of accumulated heat in the pitch region. To expel the accumulated heat, a cooling hole was drilled at the pitch point of the gear tooth. The damage in the first type of configuration of the gear tooth was less than the damage in the standard gear tooth under the same operating conditions (Fig. 10). As a result of the cooling hole that was drilled in the pitch region, the accumulated heat in this region was expelled via air circulation that was generated as a result of the rotating motion"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002129_s0263574709005426-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002129_s0263574709005426-Figure1-1.png",
        "caption": "Fig. 1. Four bar planar closed chain.",
        "texts": [
            " The simplest technique is based on the Euler integration method; given an integration interval t , if the joint positions and velocities at time tk\u22121 are known, the joint positions at time tk = tk\u22121+ t can be computed as qs(tk) = qs ( tk\u22121 ) \u2212 N\u22121 s ( tk\u22121 ) Np ( tk\u22121 ) q\u0307p(tk) t. (5) By calculating the secondary joint positions using Eq. (5), a cumulative error in qs is introduced. Therefore, Eq. (1) is not satisfied and an opening in the closed-loop chain is introduced. To illustrate this, consider the four-bar mechanism in Fig. 1. The cumulative error in q\u0307s opens the closed chain, as depicted in Fig. 2. To solve this problem, we present in Section 4 a new method to model the robot differential kinematic equation, where the position is obtained using numerical techniques in which the closure error converges exponentially to zero. To describe it, first we present the fundamental kinematics tools used in this study. Our approach is based on the method of successive screw displacement,8 on the screw representation of differential kinematics, on the Davies method and on the Assur virtual chain concept, which is briefly presented in this section",
            " The matrix Re corresponds to errors measured in rex , rey and rez virtual rotative joints considering their structural conception. The \u2018position\u2019 error (which is a posture error involving position and orientation) is given by the position error vector qe = [rex rey rez pex pey pez]T . Applying the Euler integration method in Eq. (20) we obtain qs(tk) = qs(tk\u22121) \u2212 N\u22121 s (tk\u22121)Np(tk\u22121)q\u0307p t + \u00b7 \u00b7 \u00b7 + N\u22121 s (tk\u22121)Ne(tk\u22121)Keqe t. (26) The method presented is illustrated by solving the position kinematic of a planar four bar mechanism (Fig. 1.) In this example, joint 1 is considered primary while the others are secondary. Joint 1 moves from the initial position \u03c0/4 to the final position \u03c0/2 according to qp(t) = \u03c0/4 + \u03c0/4 sin(\u03c0 t/8) from t = 0 to 4 s. The magnitude of q\u0307p(t) can be obtained by differentiating qp(t) : q\u0307p(t) = \u03c02 32 cos ( \u03c0t 8 ) . The kinematic parameters are l1 = 0.5 m, l2 = 1.0 m, l3 = 0.5 m and l4 = 1.0 m. The initial position vector is qs(0) = [\u2212\u03c0 4 \u22123\u03c0 4 \u2212\u03c0 4 ]T rad. The integration interval is t = 0.01 s. Considering the reference frame attached to the last link of the error chain, the network matrices result in: Np = \u23a1 \u23a2\u23a3 1 l3s4 + l2s34 + l1s123 l4 + l3c4 + l2c34 + l1c123 \u23a4 \u23a5\u23a6 , Ns = \u23a1 \u23a2\u23a3 1 1 1 l3s4 + l2s34 l3s4 0 l4 + l3c4 + l2c34 l4 + l3c4 l4 \u23a4 \u23a5\u23a6 , Ne = \u23a1 \u23a30 0 1 0 1 ey 1 0 \u2212ex \u23a4 \u23a6 , where: ex , ey are the prismatic displacement of the Assur virtual error chain; sij and cij are the sine and cosine of qi + qj + \u00b7 \u00b7 \u00b7"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000870_iros.2007.4399336-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000870_iros.2007.4399336-Figure2-1.png",
        "caption": "Fig. 2. Schematic of ex-vivo probe test rig",
        "texts": [
            " Ovine liver has been chosen as test sample, and a stainless steel probe of diameter 6mm as the instrument interacting with the liver. A testing facility was required to ensure that the probe was under accurate and repeatable control at all times. To facilitate this, the probe was attached to the distal tip of a Mitsubishi RV-6SL 6-DOF robotic manipulator. An ATI MINI40 Force/Torque sensor (calibration SI-20-1, resolution 0.01N with 16-bit DAQ) was mounted at the interface between the probe and the manipulator end-effector as shown in Fig.2. This allows for the measurement of the interaction force imparted by the tissue onto the probe. B. Static Indentation Tests and Curve Fitting For measuring the viscoelastic properties of ovine liver, static indentation tests were conducted. In order to obtain consistent results without pre-conditioning the tissue, each test was completed in a different position. It was ensured (as much as was possible) that the boundary conditions were identical for each test. Three ovine livers overlaid on top of each other and a homogenous region (approximately 20mm in diameter, average thickness 62mm) on the left lobe of the top liver was chosen as the test site"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002335_detc2010-28926-Figure4-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002335_detc2010-28926-Figure4-1.png",
        "caption": "Figure 4. Example of decomposition a determinate truss into Assur Graphs. a) The determinate truss. b) The decomposition",
        "texts": [
            " The composition rule for constructing a determinate truss from its components (Assur Graphs) is done as follows: Let G1 and G2 be two Assur Graphs. G1 is defined to be preceding G2 if at least one ground vertex of G1 is connected to an inner vertex of G2. The decomposition process can be presented by a directed graph in which an edge e=<u,v> indicates that the Assur Graph corresponding to vertex u is preceding another Assur Graph, presented by vertex v. This means that in order to decompose Assur Graph \u2018v\u2019, Assur Graph \u2018u\u2019 has first to be removed, thus this graph is termed in the paper \u2013 decomposition graph. For example, in Figure 4.b the graph presents the order in which the determinate truss in Figure 4.a can be decomposed. We start with the initial vertex - a vertex to which no edge is incident. In this example the initial vertex 'F' corresponds to the dyad with the inner vertex 'F' and the two edges (F,G) and (F,J). Once this dyad is removed it is possible to remove, independently the dyads G or J, and so forth. graph. From the above it follows that once we have all the Assur Graphs it is possible to construct all different determinate trusses by composing different Assur Graphs, each time in a different order"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003751_icuas.2013.6564678-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003751_icuas.2013.6564678-Figure1-1.png",
        "caption": "Fig. 1. Pure pitching motion",
        "texts": [
            " Finally, section VI presents the conclusions and future work. In order to obtain the model equations, by omitting any flexible structure of the UAV, the fixed-wing UAV is then considered as a rigid body. Also we do not consider the curvature of the earth, it is considered as a plane, because we assume that the UAV will only fly short distances. With the previous considerations, we obtain the model by applying the Newton\u2019s laws of motion. The parameters involved in the longitudinal dynamic model (1)-(5) are shown in Figure 1. These parameters allow to analyzing the movement toward the front of an airplane [12], particularly the altitude control. V\u0307 = 1 m (\u2212D + T cos\u03b1\u2212mg sin \u03b3) (1) \u03b3\u0307 = 1 mV (L+ T sin\u03b1\u2212mg) sin \u03b3) (2) \u03b8\u0307 = q (3) q\u0307 = M Iyy (4) h\u0307 = V sin(\u03b8) (5) where V is the magnitude of the airplane speed, \u03b1 describes the angle of attack, \u03b3 represents the flight-path angle and \u03b8 denotes the pitch angle. In addition, q is the pitch angular rate (with respect to the y-axis of the aircraft body), T denotes the force of engine thrust, h is the airplane altitude [12] and \u03b4e represents the elevator deviation"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000715_s11249-007-9287-9-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000715_s11249-007-9287-9-Figure1-1.png",
        "caption": "Fig. 1 Schematic diagram of 4-roller tester",
        "texts": [
            " As the final step, optimal microtexture specifications were selected on the basis of the sensitivity analysis results and applied to the traction surfaces of an actual toroidal CVT. The effect of the surface microtexture on improving the traction coefficient and on ensuring durability was examined, and the results obtained are described here. 2 Experimental Setup and Procedure 2.1 4-roller Tester A schematic diagram of the 4-roller tester used to conduct tests under high contact pressure and high rolling speed conditions is given in Fig. 1. The rollers used all had an outer diameter of 60 mm and were made of SUJ2 (SAE 52100) bearing steel having a hardness of 60 HRC. A flat roller was placed in the center, with three rollers having an axial radius of curvature of R = 30 mm placed at 120 positions around it. Slip was applied between the center flat roller and the other three rollers, and the resultant shear force was measured with a torque meter attached to the center shaft supporting the flat roller. The measured result was divided by the applied loading force to calculate the traction coefficient"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002401_j.scient.2011.05.004-Figure5-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002401_j.scient.2011.05.004-Figure5-1.png",
        "caption": "Figure 5: Swarm of robots with different sensors.",
        "texts": [
            " According to Wei, Beard and Atkins in [19], \u2018\u2018Consensus algorithms are designed to be distributed, assuming only neighbor to neighbor interaction between vehicles. Vehicles update the value of their information state based on the information states of their neighbors\u2019\u2019. Using a consensus law, the objective is to converge, to a common value, the states of all agents in the network. Consensus algorithms have been studied to solve rendezvous problems, formation control problems, flocking, and sensor networks. Figure 5 depicts a group of robots; each one carrying a different type of sensor. The formation architecture can be described through the use of graphs (Figure 6). Graph G is a pair (V, E), where V = {V1, . . . , Vn} is a finite nonempty vertex set and E \u2286 V \u00d7 V is the edge set of ordered pairs of nodes (Figure 7). The rendezvous problem [20] for robots states that: \u2018\u2018Given a group of N robots dispersed in a plane, how should theymove to gather around a specific location?\u2019\u2019 Different approaches during the last few years addressed a solution to this problem [21\u201323]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000886_iros.2007.4399168-Figure3-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000886_iros.2007.4399168-Figure3-1.png",
        "caption": "Fig. 3. Locomotion Model of Rover",
        "texts": [
            " 1228 terrain, it is needed to decide whether the emphasis is on adhesion or stability depending on the terrain type, and at the same time to ensure some margin for them. The past researches were not enough for accomplishing these aims. In this paper, a new algorithm for solving these issues is proposed. In the next session, dynamic equation is introduced and some important concepts are shown. In this section, firstly the assumption of a locomotion model is discussed. And then, the dynamic equation of a locomotion model is discussed. A locomotion model of a rover is shown in Fig.3. For the locomotion model of a wheeled robot, the following assumptions are made in this paper. \u2022 Rover moves by 4 wheel drive system (4WD) \u2022 4 wheel motors can generate driving force indepen- dently \u2022 All the wheel motors and wheels are the same \u2022 Movable property of the rover is restricted in x-y plane and not considering steer motion \u2022 Rover moves very slowly \u2022 Movement of payload is done by a linear actuator \u2022 Center of mass position of payload and body part is located in midle for z-direction \u2022 Terrain is perfectly symmetrical \u2022 Angle of terrain gradient is estimated by camera \u2022 Driving force is generated by static frictional force [ [ \u2022 Static friction coefficient between wheel and terrain is constant value \u2022 Wheels contact with terrain by single point (not consid- ering ditch like terrain) To solve the dynamic effect of a rover, dynamic equation should be introduced"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002717_0954406212454390-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002717_0954406212454390-Figure2-1.png",
        "caption": "Figure 2. Coordinate systems used in the double envelope process: (a) profile of rack-cutter; (b) generation of pinion; (c) generation of internal gear.",
        "texts": [
            "15\u201317 During the derivation, the mathematical model of the pinion is firstly regarded as an envelope to the family of the rack-cutter surfaces. Then, the obtained pinion is assumed as the generating surface and defined as the second envelope to generate internal gear. In this way, the profiles of pinion and internal gear can be easily obtained. at SIMON FRASER LIBRARY on June 16, 2015pic.sagepub.comDownloaded from The profile of rack-cutter for involute gears The involute pinion can be easily generated by a rackcutter tool. Figure 2(a) shows the profile of rackcutter for involute gears, including line ab, arc bc, line cd and line de. To illustrate the profile, a coordinate system St(Ot, xt, yt) is built and rigidly connected to the rack-cutter. The xt axis has a distance of xm from the pitch line of rack-cutter while the yt axis coinciding with the pitch line of rack-cutter. Here, x is defined as shifting coefficient. Then, the profile can be represented in St Rab t \u00bc xabt yabt zabt 2 664 3 775 \u00bc xt hf\u00fexm 1 2 664 3 775, 04xt4xb \u00f01\u00de Rbc t \u00bc xbct ybct zbct 2 6664 3 7775 \u00bc xt yOc ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi r2o \u00f0xt xOc \u00de 2 q 1 2 66664 3 77775, xb 5 xt4xc \u00f02\u00de Rcd t \u00bc xcdt ycdt zcdt 2 6664 3 7775 \u00bc xt \u00f0xt p 4\u00de cot a\u00fe xm 1 2 6664 3 7775, xc 5 xt4xd \u00f03\u00de Rde t \u00bc xdet ydet zdet 2 664 3 775 \u00bc xt ha\u00fexm 1 2 664 3 775, xd 5 xt4xe \u00f04\u00de where xt is the design parameter of rack-cutter. Here, xb\u00bc xOc\u00bc 0.25 p tana[hf ro(1 sin a)] ro cos a, yOc\u00bc (h a\u00fec n)m\u00fero\u00fexm, xc\u00bcxOc\u00ferosin( /2 a), xd\u00bcp/4\u00fehatana, xe\u00bcp/2. The profile of pinion with involute tooth To determine the first envelope, the relationship between the rack-cutter and the pinion during the generation is depicted by Figure 2(b). Coordinate systems Sf(Of, xf, yf) and S1(O1, x1, y1) are built, which are rigidly connected to the ground and the pinion respectively. The origins of Sf and S1 coincide with the pinion center, and the pitch line of rack-cutter is tangent to the pitch circle of pinion at point Pt. During the generation, St moves with the rack-cutter while S1 rotates with the pinion. When St has a displacement of s, S1 rotates over an angle of \u20191, both of which are related by equation (5). s \u00bc r1\u20191 \u00f05\u00de According to the theory of gearing, the profile of pinion can be obtained by applying coordinate transformation from St to S1, as represented by equation (6)",
            " The profile of the conjugated involute internal gear According to the double envelope concept, the profile of the conjugated involute internal gear is regarded as an envelope to the family of the obtained pinion surfaces. Now, the mathematical model of the pinion profile becomes the generating surface and the profile of internal gear is considered as the generated surface. Another coordinate system S2(O2, x2, y2) is built to demonstrate the relationship between the pinion and the internal gear during the generation (Figure 2(c)). Here, S2 is rigidly connected to the internal gear and the origin of S2 coincides with the gear center. The pitch circles of pinion and internal gear are tangent at point P. The rotation angles of S1 and S2, \u20191 and \u20192, are related by equation (12). r1\u20191 \u00bc r2\u20192 \u00f012\u00de Then, the profile of internal gear can be obtained by equations (13) and (14). Ri 2\u00f0\u20191, \u20192,xt\u00de \u00bcM21\u00f0\u20191, \u20192\u00deR p 1\u00f0\u20191, xt\u00de \u00f013\u00de Np v \u00f012\u00de 1 \u00bc 0 \u00f014\u00de where Ri 2(\u20191, \u20192, xt) is the envelope to the family of the pinion surfaces, M21 (\u20191, \u20192) is the matrix for coordinate transformation from S1 to S2 as represented by equation (15), Np is the normal to the pinion surface and v \u00f012\u00de 1 is the relative velocity between the pinion and the internal gear",
            " V2 \u00bc Z tAC \u00bdQinv\u00f0 finv\u00de \u00feQfil dt \u00f035\u00de where tAC is the lasting time for the contact point moving from point A to point C along curve AC, and Qfil is the instantaneous flowrate when contact point moves along the root fillets. Here, the volume between the involutes can be obtained by equation (36). Z tAC Qinv\u00f0 finv\u00dedt \u00bc !1 2 Z fMP fAP r2a1 r1 r2 r2a2 r1\u00f0r1 r2\u00de 1 r1 r2 f 2inv dfinv \u00f036\u00de Based on equation (24), Qfil can be represented by equation (37). Qfil \u00bc Q\u00f0xff, y f f \u00de \u00bc 1 2 r2a1 r1 r2 r2a2 r2 r1 r2 \u00f0xff \u00de 2 \u00f0 yff yO1 f \u00de 2 \u00fe r1 r2 \u00f0 yff yO2 f \u00de 2 !1 \u00f037\u00de where (xff, y f f) is the coordinate of the contact point on curve AC, which is represented in Sf. According to Figure 2(a), equation (2) can be represented with parameter : Rbc t \u00bc xbct ybct zbct 2 664 3 775\u00bc ro sin \u00fexOc r1\u20191 ro cos \u00feyOc r1 1 2 664 3 775, 05 4 2 \u00f038\u00de Equations (6), (18) and (38) yield xff \u00bc ro sin \u00fe xOc r1\u20191 yff \u00bc ro cos \u00fe yOc \u00fe r1 8< : \u00f039\u00de Substituting equation (39) into equation (37), Qfil can be represented with parameter by equation (40). Qfil \u00bc Qfil\u00f0 \u00de \u00f040\u00de Equations (11) and (39) yield \u20191\u00f0 \u00de \u00bc yOc tan \u00fe xOc r1 \u00f041\u00de The derivative of equation (41) yields r1d\u20191 \u00bc yOc cos2 d \u00f042\u00de Then, a relation can be obtained by equation (43)",
            " With illustrations mentioned above, some relations can be obtained yE1 \u00bc tan xE1 \u00fe r1 \u00f0xE1 xA1 \u00de 2 \u00fe \u00f0 yE1 yA1 \u00de 2 \u00bc p2b ( \u00f045\u00de yI1 \u00bc tan xI1 \u00fe r1 \u00f0xI1 xB1 \u00de 2 \u00fe \u00f0 yI1 yB1 \u00de 2 \u00bc p2b ( \u00f046\u00de ffAO1I \u00bc \u2019A\u00f0 A\u00de \u2019I\u00f0 J\u00de cos ffAO1I \u00bc O1A O1I O1Aj j O1Ij j O1A \u00bc \u00f0x A 1 xO1 1 \u00dei1 \u00fe \u00f0 y A 1 yO1 1 \u00dej1 O1I \u00bc \u00f0x I 1 xO1 1 \u00dei1 \u00fe \u00f0 y I 1 yO1 1 \u00dej1 8>>>>< >>>>: \u00f047\u00de yM1 \u00bc tan xM1 \u00fe r1 ffAO1M \u00bc \u2019A\u00f0 A\u00de \u2019M\u00f0 C\u00de cos ffAO1M \u00bc O1A O1M O1Aj j O1Mj j O1M \u00bc \u00f0x M 1 xO1 1 \u00dei1 \u00fe \u00f0 y M 1 yO1 1 \u00dej1 8>>>>< >>>>: \u00f048\u00de where (xO1 1 , yO1 1 ), (xA1 , y A 1 ), (xB1 , y B 1 ), (xE1 , y E 1 ), (xI1, y I 1) and (xM1 , yM1 ) are the coordinates of points O1, A, B, E, I and M represented in S1, \u2019A, \u2019I and \u2019M are the positions of the pinion when the gears mesh at points A, I and M, and I and M are the corresponding parameters to points I and M. Obviously, curve AC is generated by arc bc. And, points A and C are generated by points c and b, respectively. Then (Figure 2(a)) A \u00bc c \u00bc =2 , C \u00bc b \u00bc 0 \u00f049\u00de Finally, the total volume of trapped fluid can be represented by equation (50). V \u00bc V1, tA 5 t5 tI V2, tI 5 t5 tM \u00f050\u00de where tA, tI and tM are the time when the second tooth pair meshes at points A, I and M. Flowrate characteristics under different design parameters With a program written by Matlab codes, the functions obtained above can be solved in at SIMON FRASER LIBRARY on June 16, 2015pic.sagepub.comDownloaded from Matlab environment. Then, influences of design parameters of gears, including shifting coefficient, pressure angle, tooth number and fillet radius, on flowrate characteristics of the pump can be investigated"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002481_tpas.1969.292344-Figure7-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002481_tpas.1969.292344-Figure7-1.png",
        "caption": "Fig. 7. Short-circuit armature current in phase c.",
        "texts": [],
        "surrounding_texts": [
            "This paper presents a mathematical analysis of unbalanced operations of synchronous machines with additional field circuits. The proposed method is aimed to be more rigorous than that initially proposed by Doherty, Nickle, and Concordia [5] and should offer solutions with a higher degree of accuracy. The derived expressions for short-circuit currents, torque, and openphase voltage for a double-line-to-ground fault apparently are new and of immediate practical use. The proposed method"
        ]
    },
    {
        "image_filename": "designv11_12_0003713_978-1-4614-2419-2-Figure61.7-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003713_978-1-4614-2419-2-Figure61.7-1.png",
        "caption": "Fig. 61.7 Frequency based substructuring results using raw measured data",
        "texts": [
            " The results shown here are for the frequency based system model developed from experimentally measured components using raw measured data and after VIKING data smoothing techniques are applied. The same BUH structure was used as in the case mentioned above. The measured frequency response functions (FRFs) for several locations are shown in Fig. 61.6 to illustrate the common difficulty when dealing with directional modes and making tri-axial and cross measurements to develop the connection FRF 606 L. Thibault et al. matrix. Using the raw measured FRFs to develop the frequency based system model did not produce acceptable results by any means. Figure 61.7 clearly shows that the results obtained are not acceptable. One of the problems is that the measurement of the tri-axial FRF is very difficult due to the directional nature of the modes of the components; obtaining noise-free, high quality measurements when the structure contains very directional modes is close to impossible. However, once VIKING is used to smooth the data impurities from the data, the measurements can clearly be observed to show significant improvement, as seen in Fig. 61.8. Once the measured FRFs are processed, the system model is reformulated and the resulting system characteristics are vastly improved, as seen in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000874_s10440-007-9155-5-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000874_s10440-007-9155-5-Figure2-1.png",
        "caption": "Fig. 2 A depiction of the Ehresmann connection on \u03c0TQ : TQ \u2192 Q associated with an affine connection on Q",
        "texts": [
            " (Note that we make an abuse of notation here by using \u201chlft\u201d both for the horizontal lift on \u03c0B : Q \u2192 B and on \u03c0TQ : TQ \u2192 Q.) We can also define the vertical lift map by vlft ( \u2202 \u2202qi ) = \u2202 \u2202vi . This definition can be written intrinsically as vlftvq (wq) = d dt \u2223 \u2223 t=0 (vq + twq), defining an isomorphism from TqQ to Vvq TQ. Note that we have S(vq) = hlftvq (vq). Also note that this splitting Tvq TQ TqQ \u2295 TqQ extends the natural splitting T0q TQ TqQ \u2295 TqQ that one has on the zero section away from the zero section. We depict the situation in Fig. 2 to give the reader some intuition for what is going on. In this section we \u201clift\u201d the preceding construction to construct an Ehresmann connection on \u03c0TTQ : TTQ \u2192 TQ. This requires an affine connection on TQ. It turns out that there are various ways of lifting an affine connection on Q to its tangent bundle, and the one suited to our purposes is defined as follows [29]. Lemma 3 If \u2207 is an affine connection on Q, then there exists a unique affine connection \u2207T on TQ satisfying \u2207T XT Y T = (\u2207XY )T for vector fields X and Y on Q"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003450_09540911311309077-Figure4-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003450_09540911311309077-Figure4-1.png",
        "caption": "Figure 4 Finite element model for creep and relaxation by temperature cycling analysis of nano-silver",
        "texts": [
            " The temperature accuracy of the chamber was ^18C and had been verified using a thermo-couple. Since the variation in bondline thickness (maximum 50mm) was very small compared to the total thickness (more than 1mm) of the assembly, it was assumed that therewas nodifference in temperaturewithin the assembly for all of the samples. At least three samples for each condition were taken out from the thermal cycling chamber and their curvatures weremeasured at room temperature after 50, 100 and 300 cycles (Figure 4). Finite element analysiswasemployed to simulate thebehaviour of the joint subjected to temperature cycling from 2408C to 1258C by ANSYS. Due to the symmetry, the finite element mesh of a quarter model is shown in Figure 1, which is made up of Solid45 element with a total of 1,076 elements and 1,080 nodes. Also, the lower corner at the substrate (alumina) side was restrained in three-directions to prevent rigid rotation. Thematerial properties of thedie, adhesive and substrateused for the simulation are presented in Tables II and III"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003878_1.4031025-Figure3-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003878_1.4031025-Figure3-1.png",
        "caption": "Fig. 3 Typical bevel gear flanks; area within dashed-border lines will be measured by CMM",
        "texts": [
            " Actual surfaces should be measured against its intended theoretical surface. The theoretical surface can be spherical involute, any modifications to it, or other bevel gear surface profiles used Journal of Mechanical Design SEPTEMBER 2015, Vol. 137 / 093302-3 Downloaded From: http://mechanicaldesign.asmedigitalcollection.asme.org/ on 01/17/2016 Terms of Use: http://www.asme.org/about-asme/terms-of-use by different manufacturers depending on their manufacturing process. An area within the tooth borders needs to be specified for measurement against theoretical surface. Figure 3 shows an example bevel pinion tooth border with dashed-line area specified for measurement. This area can be meshed with certain number of rows and columns of points, each point coordinates, and unit normal vectors are calculated. A file containing coordinates and normals for both sides of the tooth is generated and supplied to a CMM machine to measure the actual parts against this theoretical file. The CMM machine approaches each point along the unit normal direction and reports the error at that point against the theoretical surface. It should be noted that CMMs measure an offset of the given theoretical surface along local normals for as much as CMM probe radius. Usually the deviation at the middle point of the grid (O0 in Fig. 3) is set to zero and deviations of the other points are calculated and reported individually with respect to middle point. The tooth thickness is specified with a difference angle a that represents the angle between two flanks at a specific R and L at the flank (usually at R and L associated with the middle point). Figure 4 shows a sample CMM measurement chart for a gear set against its intended theoretical surfaces. All the numbers in this chart are in microns, \u201cLFl\u201d stands for left flank and \u201cRFl\u201d stands for right flank",
            " The charts show the difference between theoretical and measured surfaces at each grid point along the normal direction (if the error at the middle point is set to zero). The difference angle a between two sides of tooth is also measured and reported. Depending on the number of teeth on the gear usually three to four teeth are measured and the results are averaged. The proposed approach in this study is to capture the surface deviations of both flanks of both members in the form of a 2D third-order polynomial. To simplify the representations of the deviation surfaces, coordinate system XYZ (of Fig. 3) is defined. In this coordinate system, Z is in the normal to the surface direction and X is along lengthwise (pointing toward heel) and Y is along profile direction (pointing toward tip). The origin O0 of the XYZ coordinate is located at the middle of the face width and along the pitch line, as shown in Fig. 3, and is calculated as O0 \u00bc ro FW 2 sin dp ro FW 2 cos dp 0 T (21) where ro is gear outer cone distance, FW is gear face width, and dp is gear pitch angle. With this, the difference between actual < and the theoretical < surfaces can be represented up to third-order as Z \u00bc a1 \u00fe a2X \u00fe a3Y \u00fe a4X2 \u00fe a5XY \u00fe a6Y2 \u00fe a7X3 \u00fe a8X2Y \u00fe a9XY2 \u00fe a10Y3 (22) independently for each side of the tooth for each member (pinion or gear). After measuring the tooth, the goal is to modify the theoretical surface such that it matches with the actual measured surface"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003973_s12206-012-0811-y-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003973_s12206-012-0811-y-Figure1-1.png",
        "caption": "Fig. 1. The assembly draft and dimensional chain.",
        "texts": [
            " In this study, the deformation due to gravity effect is determined using finite element analysis and it is suitably incorporated in the tolerance stack up equation of tolerance design, thereby loosening tolerance requirement of critical components. Since the deformation is included in the early stages of design, the optimal tolerance values of some critical components of the assembly obtained are higher than that of those vales obtained by conventional method, resulting in reduction of the total manufacturing cost of the assembly. The piston-cylinder assembly [18] is the application example for the proposed tolerance design. Fig. 1 shows a graphic representation of the assembly along with dimensions. The piston-cylinder assembly consists of four components, which are piston, cylinder, connecting rod and crank shaft. The piston is made of aluminum alloy and the crank shaft and connecting rod are made of alloy steel. In order to maintain the compression ratio at a constant value, the tolerance of the clearance has to be maintained at a particular value. So the objective is to allocate appropriate tolerance in order to maintain sufficient clearance between the piston top surface and the cylinder top surface. First, a 3D model of the assembly is created using Pro/E wildfire 5.0 software as shown in Fig. 2. To determine the features which have an effect on the clearance measurement, a vector loop model of the assembly has been created as shown in Fig. 1. The assembly function that describes the quality value is: (0.7) (0) (0) (0) (50) (174) (69) (292.3) 2 4 6 3 5 7 1x x x x x x x x .\u2211 = + + + + + \u2212 (1) Once a three-dimensional model of the assembly is created, the next step is to develop a finite element model of the same to determine deformation of various components and their effect on the clearance. The assembly consists of four components and there are three contact pairs. The elements used in the analysis are SOLID 92 (Fig. 3), which has quadratic displacement behavior and is well suited to model irregular meshes such as produced from CAD data"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001301_s00466-009-0394-3-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001301_s00466-009-0394-3-Figure1-1.png",
        "caption": "Fig. 1 Rod seal in the deformed configuration",
        "texts": [
            " for hollow-cylinder rods in high-pressure hydraulic systems, cf. Nikas [18]. Modelling of the solid part is rather standard. Two configurations are introduced\u2014the stress-free initial configuration and the deformed configuration \u03c9. The boundary \u2202 is divided into three non-overlapping parts \u2202l , \u2202p and \u2202c associated with the hydrodynamic lubrication, hydrostatic sealed pressure and contact interaction with the housing, respectively. The deformed-configuration counterparts are \u2202l\u03c9, \u2202p\u03c9 and \u2202c\u03c9, cf. Fig. 1. This division is not given a priori and, in general, depends on the deformation of the seal. This is commented more in Sect. 3. The deformation from to \u03c9 is given by a continuous mapping x = \u03d5(X), where X \u2208 and x \u2208 \u03c9. In the absence of body and inertia forces, the weak form of the equilibrium equation reads \u222b P \u00b7 Grad \u03b4\u03d5 dV \u2212 \u222b \u2202 T \u00b7 \u03b4\u03d5 dS = 0, (1) where P is the first Piola\u2013Kirchhoff stress tensor and T is the nominal traction on the boundary \u2202 . A hyperelastic material model is assumed for the elastomeric seal",
            " Importantly, the dimension of the problem is reduced so that the Reynolds equation is two-dimensional in the general case and one-dimensional if the flow does not dependent on one spatial variable. Furthermore, the Reynolds equation is formulated in the Eulerian frame so that, in the case considered in this work, the Reynolds equation and all the quantities involved refer to the lubricated boundary \u2202l\u03c9 in the deformed configuration. As in the present model the rod is assumed to be rigid, it is convenient to introduce a domain, denoted by \u03b3 , which is a projection of the deformed boundary \u2202l\u03c9 onto the rigid surface of the rod, cf. Fig. 1. The Reynolds equation is then formulated on the domain \u03b3 . For an incompressible fluid, the Reynolds equation takes the form div\u03b3 q + \u2202h \u2202t = 0, q = u\u0304h \u2212 h3 12\u03b7 grad\u03b3 p, (11) where p is the hydrodynamic pressure, h the film thickness, q the lubricant flux, u\u0304 the average velocity of the contacting surfaces, and \u03b7 the viscosity. The gradient and divergence in Eq. (11) are defined on the surface \u03b3 , hence a subscript introduced in the respective operators. The essential and the natural boundary conditions are enforced on the respective parts of the boundary \u2202\u03b3 , namely p = p\u2217 on \u2202p\u03b3, q \u00b7 n\u03b3 = q\u2217 n on \u2202q\u03b3, (12) where p\u2217 is the prescribed pressure and q\u2217 n is the prescribed flux",
            " As isothermal conditions are only considered, the important temperature dependence of viscosity is not introduced. Having in mind the numerical examples presented in Sect. 5, the Reynolds equation can be simplified significantly. Axial symmetry is assumed, so that the Reynolds equation becomes one-dimensional, and the term \u2202h/\u2202t vanishes in steady-state conditions. The corresponding one-dimensional Reynolds equation reads d dx\u0304 ( u\u0304h \u2212 h3 12\u03b7 dp dx\u0304 ) = 0, (14) where u\u0304 = U/2, U is the rod speed (positive for outstroke, as indicated in Fig. 1, and negative for instroke), and the spatial variable x\u0304 is a local variable which parameterizes \u03b3 . Equation (14) is accompanied by the boundary conditions p(x\u0304s) = ps, p(x\u03040) = 0, (15) which are enforced on the sealed-pressure side (x\u0304 = x\u0304s) and on the air side (x\u0304 = x\u03040). It is clear that the Reynolds equation in the form (14) states that the flux q given by the term in brackets is constant along the lubricated boundary. It is also seen that at the points of zero pressure gradient (e.g. at the point of maximum pressure) we have q = u\u0304h\u2217, h\u2217 = q u\u0304 = 2q U , (16) where h\u2217 is the characteristic thickness of the lubricant film",
            " 5.2. Finally, the effect of pressure dependence of viscosity (piezo-viscous effect) is illustrated in Sect. 5.3. 5.1 Hydrodynamic lubrication in reciprocating O-ring seal The analysis below is carried out for a reciprocating rod seal. Hydrodynamic lubrication in steady-state conditions is studied during both outstroke (the rod moves towards the air side, U > 0) and instroke (the rod moves towards the sealedpressure side, U < 0). The general arrangement of the sealrod-housing system is shown in Fig. 1, and the geometrical and process parameters are provided in Table 1. The elastic properties of the elastomeric seal (NBR rubber, 70 ShA hardness) and the viscosity of the hydraulic fluid (Shell Tellus 46 oil) at the working temperature of 30\u25e6C are also given in Table 1. Five densities of the structured finite-element mesh have been used in the convergence studies presented in Sect. 5.2. Based on these studies, the finest mesh density 16 and pressure interpolation order m p = 4 have been used in the computations reported below in order to ensure reliable results in the whole range of process parameters covered by the present study"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001487_j.rehab.2009.12.001-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001487_j.rehab.2009.12.001-Figure2-1.png",
        "caption": "Fig. 2. Sche\u0301ma du prototype avec la technique alternative de propulsion (a\u0300 leviers).",
        "texts": [],
        "surrounding_texts": [
            "Les objectifs sont de\u0301crits comme suit : identifier les points forts et les points faibles du prototype pour les diffe\u0301rents terrains et dans les diffe\u0301rentes conditions d\u2019utilisation. solliciter la participation des utilisateurs de\u0301pendants du fauteuil roulant pour contribuer a\u0300 l\u2019ame\u0301lioration du prototype. e\u0301valuer la satisfaction globale des utilisateurs apre\u0300s la propulsion du prototype."
        ]
    },
    {
        "image_filename": "designv11_12_0001345_robot.2008.4543604-Figure6-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001345_robot.2008.4543604-Figure6-1.png",
        "caption": "Figure 6. (a) The grasping position; (b) The grasping parameters",
        "texts": [
            " Position Control We consider that the initial state of the system is given by Tqs 000 ,,0 (15) Ts 0,0,00 (16) where s,00 , , (17)sqq ,00 Ls ,0 corresponding to the initial position of the arm defined by the curve 0C sqsC 000 ,: , Ls ,0 (18) The desired point in is represent by a desired position of the arm, the curve that coils the load,dC T ddd q, (19) T d 0,0 (20) sqsC ddd ,: , Ls ,0 (21) In a grasping function by coiling, only the last mn 1 elements Nm are used. Let be the active grasping length, gl n mi ig ll (22) Let be the curve defines the boundary of the load and we denote by the origin of the coiling function, when is the intersection between the tangent from origin O and the curve (Figure 6.b). This curve can be expressed in the coordinates bC bO bO LC q, . sqsC bbb ,: , bLs ,0 (23) where is the length of the coiling measured on the boundary and bL bC slLs g (24) We define by te p the position error L lL bbp g dssqtsqstste ,, (25) It is difficult to measure practically the angles , for all q Ls ,0 . These angles can be evaluated or measured at the terminal point of each element. In this case, the relation (25) becomes N mi biibiip qtqtte (26) The error can also be expressed with respect to the global desired position Cd N i diidiip qtqtte 1 (27) (28) N i qiip tetete 1 The position control of the arm means the motion control from the initial position to the desired position in order to minimize the error"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000597_978-3-540-73129-0-Figure6.48-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000597_978-3-540-73129-0-Figure6.48-1.png",
        "caption": "Fig. 6.48. Two wireless micromotors. a Standard form, b rotating form",
        "texts": [
            " Several standing-wave ultrasonic motors (SWUMs), have been designed at Cedrat [68] (Figs. 6.48 and 6.49). A linear motor is a self-moving silicon plate including magnetostrictive film. It is submitted to a 10mT dynamic field produced by an external coil, which may be placed at some centimeters distance from the motor. At resonance, this field excites a flexure mode, producing vibrations in the plate, which in turn induces by friction a motor motion at 10 . . . 20mm/s. A rotating version has been also created (Fig. 6.48b) that uses a slightly different principle [68]: the vibrating rotor is based on a 100\u00b5m thick by 20mm diameter plate with 10\u00b5m deposited magnetostrictive films, which are wireless and excited by a small coil. Typical performance is a rotating speed of 30 rpm and a torque of 1.6\u00b5Nm, with a 20mT excitation field. These examples demonstrate some of the special advantages of magnetostriction, especially the fact that the moving parts are wireless. The disadvantage is the coil, which is difficult to miniaturise because of the field requirements"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000418_j.engfracmech.2007.01.016-Figure5-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000418_j.engfracmech.2007.01.016-Figure5-1.png",
        "caption": "Fig. 5. Moving contact loading configurations in respect to the initial crack position.",
        "texts": [
            " The lubricant pressure in the crack can be simply approximated with a uniform pressure distribution along the crack faces [10], where its level equals the pressure level determined at the crack mouth. Fig. 4 illustrates crack face pressure determination and distribution for two consecutive contact loading configurations. For more realistic simulation of the fatigue crack propagation on gear teeth flanks it is necessary to consider the influence of moving gear teeth contact in the vicinity of initial crack. The moving contact can be simulated with different loading configurations as it is shown in Fig. 5. Five contact loading configurations have been considered, each with the same normal p(x) and tangential q(x) contact loading distributions, but acting at different positions in respect to the initial crack. Experimental evidence suggests that the crack lengths associated with the process of surface pitting are very short, i.e. they span only several material grains [14]. Cracks are defined as being short when their length is small compared to relevant microstructural dimensions (a continuum mechanics limitation) or when their length is small compared to the scale of local plasticity (a linear elastic fracture mechanic limitation) [15]",
            " Thorough investigation of the crack growth dependence on position and size of various initial defects shall be performed in future investigations with further parametric simulations based on the proposed model. The crack increments were chosen to be equal to Da = 1.5 lm both for the VCE and the SED methods for easier results comparison. Five different loading configurations have been considered in each crack extension increment computation for the purpose of simulating the effect of the moving gear teeth contact (see Section 2.2 and Fig. 5). The hydraulic pressure mechanism of trapped fluid in the crack was simulated with distributed pressure loading along the crack faces (see Section 2.1). The pressure distribution was constant along the crack faces, apart from the faces of the crack tip elements, where triangular pressure distribution was adopted with pressure linearly dropping to zero at the crack tip. The same triangular isoparametric special crack tip elements have been employed in the first row of elements around the crack tip for both the VCE and the SED methods",
            " Configuration of the final surface pit. Crack propagation simulations were in both applied methods stopped after the fourth increment, when the combined stress intensity factor K became so large that the full fracture and appearance of the surface pit was imminent. The crack was then projected to the free surface in direction corresponding to the Kmax obtained in the fourth increment. From Table 1 it can be seen that the largest combined stress intensity factors K in both applied methods always corresponds to load case 3 (see Fig. 5), which can clearly be observed also from Fig. 14 for K\u2019s obtained with the VCE method. Study of results for SED criterion returns the same conclusion. This was expected, since in this loading configuration the pressure at crack mouth and along crack faces is the largest, which results in increased mode I and II separation at the crack tip. Fig. 13 summarises the crack propagation paths for both methods. From Table 1 and Fig. 13 one can observe steeper crack path for the VCE method, while the SED criterion produces somewhat shallower crack path, see also Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003713_978-1-4614-2419-2-Figure32.1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003713_978-1-4614-2419-2-Figure32.1-1.png",
        "caption": "Fig. 32.1 Three-story spatial frame structure (a) side view, (b) isometric view and (c) plan view",
        "texts": [
            " Once all properties for each damage case were determined, various damage indicators were applied. Finally, after the application of several damage detection methods, the effectiveness of each technique was evaluated and compared. Final recommendations for the three-story spatial frame were then made based on the results of all damage indicators. A three-story spatial frame, single bay structure was constructed from steel and aluminum in order to test and determine the relative effectiveness of several damage detection algorithms (Fig. 32.1). The columns are comprised of four continuous 0.3175 cm (1/8-in.) aluminum angles with 2.54 cm (1 in.) flanges, measuring 60.96 cm (24 in.) in height. The foundation of the structure consists of two stainless steel 12.70 cm (500) flanges bolted to each column and fastened to a massive granite table. Three 0.0508 cm (0.0200) thick steel shim sheet squares measuring approximately 10.16 cm by 10.16 cm (400 by 400) are fastened to the columns 15.24, 35.56, and 55.88 cm (600, 1400, and 2200) from the base"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003467_gt2013-95424-Figure14-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003467_gt2013-95424-Figure14-1.png",
        "caption": "Figure 14: Baseline CFD domain for heat transfer analysis",
        "texts": [
            " In this section the validated methodology is used to predict the wear for all three seal configurations shown in Figure 4. The first step was to determine the heat transfer coefficients for the seal fins. Subsequently these results were used in the FEA to simulate the wear an the seal strips. As for these calculations the same approach as described in section WEAR SAMPLE - CFD HEAT TRANSFER is used, only the differences will be highlighted in this section. For all seal configurations an axis symmetric fluid domain containing only one respectively two seal strips was used (cf. Figure 14 - Figure 16). In the FEA simulations the results for these seal strips were applied on all other seal strips of the corresponding configuration. In all cases the wall temperature of the seals was set to be constant at 800K whereat all other walls were set to be isothermal at 300K. A rotor surface speed of 165 m/s was applied, inducing a swirl flow in the CFD domain. Since there will be no axial flow when seals are in contact with the rotor, the fluid domains feature no in- or outlet. The resulting trends for airh on the surface of the seals are outlined together with the result for the wear sample in Figure 17"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001193_j.mechrescom.2009.07.007-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001193_j.mechrescom.2009.07.007-Figure1-1.png",
        "caption": "Fig. 1. Types of non-cylindrical helical spring and impulsive load.",
        "texts": [
            " The infinitesimal length element of the non-cylindrical helix is defined as c\u00f0/\u00de \u00bc ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi R2\u00f0/\u00de \u00fe h2\u00f0/\u00de q ; ds \u00bc c\u00f0/\u00ded/; cos a \u00bc R\u00f0/\u00de=c\u00f0/\u00de; sin a \u00bc h\u00f0/\u00de=c\u00f0/\u00de \u00f02\u00de where a and R\u00f0/\u00de are pitch angle and centreline radius of the helix, respectively. The curvatures of a non-cylindrical helical spring are v\u00f0/\u00de \u00bc R\u00f0/\u00de=c2\u00f0/\u00de; s\u00f0/\u00de \u00bc h\u00f0/\u00de=c2\u00f0/\u00de \u00f03\u00de The horizontal radius of any point on the axis of helix for barrel and hyperboloidal types can be determined by R\u00f0/\u00de \u00bc R1 \u00fe \u00f0R2 R1\u00de 1 / np \u00f04\u00de and for conical type can be determined by R\u00f0/\u00de \u00bc R1 \u00fe \u00f0R2 R1\u00de / 2np \u00f05\u00de where n is the number of active turns (Fig. 1). The relationship between the moving axis (t, n, b) and the fixed reference frame (i, j, k) is fVgT tnb \u00bc \u00bdB fVg T ijk Vt Vn Vb 8>< >: 9>= >; \u00bc \u00f0R\u00f0/\u00de=c\u00f0/\u00de\u00de sin / \u00f0R\u00f0/\u00de=c\u00f0/\u00de\u00de cos / \u00f0h\u00f0/\u00de=c\u00f0/\u00de\u00de cos / sin / 0 \u00f0h\u00f0/\u00de=c\u00f0/\u00de\u00de sin / \u00f0h\u00f0/\u00de=c\u00f0/\u00de\u00de cos / \u00f0R\u00f0/\u00de=c\u00f0/\u00de\u00de 2 64 3 75 Vi Vj Vk 8>< >: 9>= >; \u00f06\u00de Let the displacement of a point on the rod axis be Uo(s, t), and the rotation of the cross-section about an axis passing through the geometric centre G be Xo(s, t). Assuming the displacements and the deformations are infinitesimal, and that the material of the rod is homogenous, linear elastic and isotropic the governing equations of a space rod are obtained in vectorial form as R2/R 1",
            " Both the dynamic stiffness matrix and fixed end forces are derived by the complementary function methods in Laplace domain (for more detail in depth see \u00c7al\u0131m, 2003; Temel and \u00c7al\u0131m, 2003). The Durbin\u2019s inverse Laplace transform (Durbin, 1974; Narayanan, 1979; \u00c7al\u0131m, 2003; Temel and \u00c7al\u0131m, 2003) is applied for the transformation from the Laplace domain to the time domain. In order to validate the developed computer program, the free vibration frequencies of a non-cylindrical helical spring are compared with the results available in the literature. Different types of non-cylindrical helical springs considered in this paper are portrayed in Fig. 1a\u2013c. Example 1. In this example, hyperboloidal and barrel type springs fixed at both ends are considered in Fig. 1a and b. For this particular example, the pitch angle and the number of active turns are a = 4.8 and n = 6.5, respectively. In addition, the material properties are E = 2.06 1011 N/m2, m = 0.3, and q = 7900 kg/m3. Free vibration frequencies calculated for non-cylindrical helical spring types by using the present computer program are given as in Tables 1 and 2. The first six natural frequencies obtained are compared with the theoretical and experimental data given in the literature. It can be seen from the tables that the result of the present model demonstrates a good agreement with the previous results. As R2/R1 ratios of hyperboloidal and barrel helices increase, natural frequencies decrease. Example 2. In this example, a triangular impulsive load (Fig. 1d) is applied vertically at the point at arc-length mid-point of the hyperboloidal type spring fixed at both ends as seen in Fig. 1a. The spring has a circular cross-section with a diameter of d = 1 mm. The pitch angle and the number of active turns are chosen as a = 8.5744 , and n = 7.6, respectively. Assuming, the material properties are E = 2.06 1011 N/m2, m = 0.3 and q = 7900 kg/m3. Moreover, for this example R1, and c are taken as 5 mm and 0.0025 s, respectively. A time increment of 0.00005 s, Dt, is used in the calculations. Vertical displacement Uz at the point at arc-length midpoint of the spring and Mz moment at fixed end of the spring are compared with the results of the ANSYS (2008) in Fig",
            " These results using the present method have shown perfect agreement with those using ANSYS. The variation of vertical displacement at the arc-length midpoint with respect to time is shown in Fig. 3 based on the different R2/R1 ratios. As seen in Fig. 3, as the R2/R1 ratios increase, both vibration period and displacement amplitude increase. Example 3. A triangular impulsive load is applied vertically at the point at the arc-length mid-point of the barrel and conical type springs fixed at both ends as in Fig. 1b and c. Both types have the same geometrical and material properties, as similar to the previous example, but R1 in Fig. 1b\u2013c and c in Fig. 1d are taken as 25 mm and 0.025 s, respectively. A time increment of 0.0005 s, Dt, is used in the calculations. Vertical displacement at the arc-length mid-point of the spring and bending moment at the fixed end for barrel and conical types with respect to time are illustrated in Figs. 4a, b and 6a, b, respectively. Results obtained in this study agree well with the ones obtained from ANSYS. Vertical displacements at the arc-length midpoint for various R2/R1 ratios are shown in Figs. 5 and 7 for barrel and conical type springs, respectively"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003895_ls.171-Figure5-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003895_ls.171-Figure5-1.png",
        "caption": "Figure 5. Detailed configuration of the bump foil.",
        "texts": [
            " The steady-state film thickness can be found using the cosine rule of triangle Ob Oj0 M : R\u00fe h0\u00f0 \u00de2 \u00bc R\u00fe C \u00fe U0\u00f0 \u00de2 \u00fe e20 \u00fe 2 R\u00fe C \u00fe U0\u00f0 \u00dee0 cos\u03b8 (1) Expanding this equation, dividing by R2 and neglecting second-order terms in h0/R, C/R, e0/R and U0/R yields h0 \u00bc C 1\u00fe e0 cos\u03b8\u00f0 \u00de \u00fe U0 (2) As a first approximation, the corrugated subfoil is modelled as a simple Winkler elastic foundation, i.e. the stiffness of a bump is uniformly distributed throughout the bearing surface (isotropic stiffness) as depicted in Figure 4. With this consideration, the static radial deformation of a bump is proportional to the pressure difference (p0 pa), i.e. U0 \u00bc L0 p0 pa\u00f0 \u00de (3) where p0 and pa are the static gas-film and ambient pressures, respectively.L0 represents the compliance of the bump foil, inversely proportional to the bump foil stiffness Kb. With the geometry specified in Figure 5, L0 is defined as L0 \u00bc 2s E l tb 3 1 s2 (4) where s is the bump pitch, l is half of the bump length, tb is the bump foil thickness and E and s are the Young\u2019s modulus and Poisson\u2019s ratio of bump foil material, respectively. Accordingly, the static film thickness is written as h0 \u00bc C 1\u00fe e0 cos\u03b8\u00f0 \u00de \u00fe L0 p0 pa\u00f0 \u00de (5) In dimensionless form, Equation (5) reads ~h0 \u00bc 1\u00fe e0 cos\u03b8\u00fe a ~p0 1\u00f0 \u00de (6) where a is the dimensionless compliance operator defined as Copyright \u00a9 2012 John Wiley & Sons, Ltd. Lubrication Science 2012; 24:95\u2013128 DOI: 10"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000226_978-1-4020-6500-2-Figure5-11-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000226_978-1-4020-6500-2-Figure5-11-1.png",
        "caption": "Figure 5-11. ( s\u20131 for aliphati with identical GGA C-3\u2019 (to number of me",
        "texts": [
            " (Adapted ported by H ortho-quino R ual helices ecially at n more than 4 can-rate cyc all of the co 0 s\u20131) are re , the results omologous intervening llustrated in ly with n. sfer rate co ith distance eory. Nota essentially i alkylthiol b chrome c (\u03b2 alue of 1.0 s to an ext comparable ly to metal s f DM-DNA c s (\u03bd = 1 V s\u20131 on sites: 5\u2019-AT 3\u2019 (botton). (B from ref. 64). eller and as ne probe m FACE within the D egative app 5 \u00c5 in the s lic voltamm njugates are asonable va suggest tha DM-DNAmethylene u Figure 5-11 Using Lavir nstants, ks, , consistent bly, the d dentical to t ridge to bo = 1.0 per CH per bond for rapolated, z to those fo urfaces.66 onjugates at \u03bd = ) of DM-DNA C CTC AAT ) Plot of ln(k sociates.41 T olecule, P 147 NA lied eries etry the lues t ET thiol nits , the on\u2019s it is with ecay hose und 2)67 the eround 1 V films CAT s) vs. heir QQ, Electrochemistry at the DNA/Electrode Interface 148 covalently attached to the top of 12 base-pair duplexes similarly selfassembled onto gold electrodes via alkylthiol linkers"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003603_0954406212466479-Figure6-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003603_0954406212466479-Figure6-1.png",
        "caption": "Figure 6. Tooth shapes (tooth space) of: (a) rack-cutter and (b) internal gear.",
        "texts": [
            " Using the same method, the modified profile of the internal gear is demonstrated in Figure 5(b). To get a brief description, the term Table 1. Parameters of the internal gear pair. Parameter Value z1 28 z2 38 s 5 mm 25 h a 1 h d 1 at University of Bristol Library on January 6, 2015pic.sagepub.comDownloaded from profilemeans the modified profile in the following part of this article. As mentioned above, shapes of the tooth spaces of the rack-cutter and the internal gear can be obtained, as shown in Figure 6(a) and (b), respectively. With Figures 2 and 6, the basic geometric dimensions of the rack-cutter, the pinion, and the internal gear can be calculated by the parameters of z1, z2, s, , h a, and h d, as listed in Table 2. Based on the kinematic relationships shown in Figure 3, Figure 7 demonstrates the generating processes of the pinion and the internal gear by their at University of Bristol Library on January 6, 2015pic.sagepub.comDownloaded from corresponding cutting tool. Here, the rack-cutter is located outside the pinion, while the shape-cutter is located inside the internal gear"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000224_1-4020-2933-0_13-Figure9-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000224_1-4020-2933-0_13-Figure9-1.png",
        "caption": "Figure 9. A secondary point defined on a) two and b) three point systems",
        "texts": [
            " (27) For a body with 2 primary points we have 6 coordinates and one constraint; i.e., 5 DoF; 3 primary points provide 9 coordinates, 3 constraints, 6 DoF; and 4 primary points not being in the same plane yield 12 coordinates, 6 constraints, and 6 DoF. Non-primary or secondary points can also be defined on a body. The coordinates of a secondary point are computed from the coordinates of the primary points. For a body defined by two primary points, the coordinates of the secondary point A, as shown in Figure 9a), can be computed as: 211221 ,/)rr(r laaA \u2212= , (28) where a1 and a2 are directional constants. For a body with three primary points as shown in Figure 9b), the coordinates of the secondary point A, which is not necessarily in the plane of the primary points, are computed as: )rr)(r~r~()rr()rr(rr 131231321211 \u2212\u2212+\u2212+\u2212+= aaaA . (29) The three constants can be determined in several ways. For example if r1, r2, r3, and rA are known initially, we compute s2,1 = r2 - r1, s3,1 = r3 - r1, sA = rAr1, 1312 ,, ss~s = . We then describe ssss ,, 3132121 aaaA ++= , [ ]sssS 3,12,1\u2261 , { }T aaa 321\u2261a and sA =Sa. Solving a = S-1 sA yields the three constants. Kinematic joints can be described between bodies by either allowing primary points to be shared between bodies, or describing simple conditions between vectors that connect primary points"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001123_j.neucom.2006.05.018-Figure4-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001123_j.neucom.2006.05.018-Figure4-1.png",
        "caption": "Fig. 4. (a) PC-based experimental cont",
        "texts": [
            " A friction force takes rol system for bw \u00bc 0.02, bm \u00bc bv \u00bc br \u00bc 0.002: (a)\u2013(c) due to periodic step ARTICLE IN PRESS Y.-F. Peng, C.-M. Lin / Neurocomputing 70 (2007) 2626\u201326372634 place at the surface between the spacer and the table. The face of the spacer pressed against the table is designed to move away from the table during part of the cycle when the spacer is moving opposite to the direction of motion applied to the table. The configuration of the personal computer (PC)-based LUSM experimental system is shown in Fig. 4. The adopted LUSM is a HR4 motor manufactured by Nanomotion with the specifications listed in Table 1. A servo control card is installed in the control PC, which includes multi-channels of digital to analog (D/A) converter, analog to digital (A/D) converter, PIO and encoder interface circuits. The measured analog signals are converted to digital values using the A/D converter with a 12- bit resolution. The position of the moving table is fed back using a linear scale. Digital filters and frequency multiplied by 4 circuits are built into the encoder interface circuits to increase the precision of position feedback"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000410_0094-114x(80)90020-8-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000410_0094-114x(80)90020-8-Figure1-1.png",
        "caption": "Figure 1.",
        "texts": [],
        "surrounding_texts": [
            "Introduction ONE OF THE major tasks facing design-oriented kinematicians today is that of synthesising true spatial linkages, both for general industrial purposes and for more specific applications, such as manipulators. The considerable problems in planar linkage synthesis, manifested by limiting configurations, branching, ordering of design positions and poor transmission, are greatly magnified in spatial chains, both due to the extra dimension and because analysis of 3- dimensional loops is far from fully developed.\nThe resolution of these difficulties is ultimately dependent on relationships among locations and orientations of joint axes in a specific linkage, together with an account of the physical nature of each joint involved. The relationships cited are known as closure equations, or displacement equations, which describe the geometry of the spatial chain and may be presented in a variety of forms. The characteristics of a joint are most concisely defined by screw coordinates, for which there is also some variation in expression; relationships among these coordinates have been dubbed, collectively, \"kinematic geometry\" because of the nexus between the study and classical line geometry.\nIt has already been claimed[I-3] that screw system algebra, as developed through Ball[4], Everett[5] and Waldron[6-8], is a natural, effective and wide-ranging tool in this context, assisted by closure equations in the form adopted by Waldron[7, 9]. It is hoped to strengthen that contention by further application of the theory in this paper. Opportunity will be taken to remedy some unfortunate looseness of terminology used in [1-3]. No originality is claimed herein for the fundamental concepts discussed and applied, nor are any numerical results presented; hence, the paper may best be regarded as tutorial in nature.\nThe basic references for the phenomena underlying the abovementioned obstacles to linkage synthesis are undoubtedly Hunt's works [10, 11] on screw system geometry. (The latter of these references, in fact, may be said to include the former, since it provides a revised and more detailed treatment of the subject.) The matters of stationary positions and uncertainty configurations are therein fully delineated, their place in the geometry described and their physical implications discussed. These sources are strongly recommended to the reader, and any exposition here of the phenomena is unwarranted. We shall summarise the principal features below, as appropriate. Some other clarifying remarks are made in the contribution to discussion of [12].\ntSenior Lecturer, Department of Applied Mechanics, The University of New South Wales. P.O. Box 1, Kensington, N.S.W. 2033, Australia.\n255",
            "1. Preliminaries The form of screw system algebra applied here is described and/or used in several places[l-3, 6--8, 13], and so it should not be necessary to repeat it in detail. With reference to Fig. l, $2 is the instantaneous screw axis (ISA) by which body 23 screws about body 12, and is identifiable, in a kinematic chain, with the joint which connects the two bodies. It is therefore representative of a relative motion but, superimposed on other relative motions, it gives rise to an 'absolute' motion of body 23.0 is the origin of some fundamental frame of reference and P2 is the normal vector from 0 to $2.\nThe screw motor\nS: = (o~2, #2)\nis a 6-element vector made up as follows. to2 is the angular velocity vector of $2. ju2 is the velocity relative to 12 of a point in 23 instantaneously at O, and\nlu2 = h2t~2 + P2 X (02-\nh2 is the pitch of $2. For the purpose of superimposing relative motions, it is more convenient to use an ISA vector\n$2 = (d,2, ti2),\nin which\ntb2 = to2, unless o~: = O, o.) 2\nand\n/i 2 = iu_--~2; tO 2\nif\nto 2 = 0, iti2 = -- . #2\nNow there is a close relationship between the ISA vector and the unit line vector (or the line coordinates) as used by Milne[14], for example, and which, in turn, is directly related to Pliicker coordinates. For the line coincident with $2 we may write the unit line vector\nI~ = (A2, ~'~),",
            "where\nHence,\nti2 = h2cb2 + v~.\nOf significance in what follows will be the algebraic condition for two lines, given by Ii and 12 say, to intersect. We may state it as\n(o1\" v2 + cb2\" vl =0. (1.1)\nThis condition is equivalent to that for two screws of zero pitch to be reciprocal, the more general case being governed by\n&l\"/12 + d~2\" ill = 0. (1.2)\nThe form of the closure equations to be used is well-known, the standard symbols being most conveniently defined by means of Fig. 2. In addition, we use the abbreviations s for sine and c for cosine. For the linkage we choose as our leading example, the 6-bar closure equations are the relevant ones. A basic set of twelve equations is given in the Appendix. It is now a common ploy to obtain alternative equations by advancing subscripts in those listed.\nA second, less common, example will afterwards be used to further illustrate the flexibility and power of the theory.\n2. Discussion We choose to explain the detailed application of screw system algebra to special configurations by means of the R-S-C-R- chain, for two reasons. First, Hunt uses this linkage as an example in [10, 11], and therefore comparisons of algebraic and geometric considerations of special positions will be facilitated. Second, this loop does seem to have a reasonable possibility for actual use. The purpose of what follows is to demonstrate an application of the theory to very important and barely investigated phenomena. No attempt is made to solve in detail specific, dimensioned cases. Such is here regarded as belonging more to the realm of numerical analysis than algebraic kinematics."
        ]
    },
    {
        "image_filename": "designv11_12_0000726_s11044-007-9072-4-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000726_s11044-007-9072-4-Figure2-1.png",
        "caption": "Fig. 2 Element e",
        "texts": [
            " The relation between the transformation matrix A and Tait\u2013Brian angle is given by [12] A = [ c\u03b2c\u03b3 \u2212c\u03b2s\u03b3 s\u03b2 s\u03b1s\u03b2c\u03b3 + c\u03b1s\u03b3 \u2212s\u03b1s\u03b2s\u03b3 + c\u03b1c\u03b3 \u2212s\u03b1c\u03b2 \u2212c\u03b1s\u03b2c\u03b3 + s\u03b1s\u03b3 c\u03b1s\u03b2s\u03b3 + s\u03b1c\u03b3 c\u03b1c\u03b2 ] , (14) where c\u03b1, s\u03b1, c\u03b2, s\u03b2, c\u03b3, s\u03b3 are denoted as cos\u03b1, sin\u03b1, cos\u03b2, sin\u03b2, cos\u03b3, sin\u03b3 , respectively. Differentiating (10) yields r\u0307 = r\u03070 \u2212 A(\u03c1\u0303 + \u03be\u0303 ) \u03c9 + A(\u03c1\u0307 + \u03be\u0307 ), (15) where \u03c9 can be expressed as \u03c9 = D\u0398\u0307, D = [ c\u03b2c\u03b3 s\u03b3 0 \u2212c\u03b2s\u03b3 c\u03b3 0 s\u03b2 0 1 ] , (16) and \u03c1\u0303, \u03be\u0303 represent the skew-symmetric matrices corresponding to \u03c1, \u03be , respectively. Differentiation of (15) leads to r\u0308 = r\u03080 \u2212 A(\u03c1\u0303 + \u03be\u0303 )\u03c9\u0307 + A(\u03c1\u0308 + \u03be\u0308 ) + 2A\u03c9\u0303(\u03c1\u0307 + \u03be\u0307 ) + A\u03c9\u0303\u03c9\u0303(\u03c1 + \u03be), (17) where \u03c9\u0307 can be written as \u03c9\u0307 = D\u0398\u0308 + D\u0307\u0398\u0307. (18) The beam is divided into n elements. As shown in Fig. 2, for element e, le is the element length, and Oe\u2013XeYeZe is the element frame, and x\u0304 is the axial coordinate of k0 defined in the element frame. Let j1, j2 be two nodes of the element, the position vector \u03c1 and the torsion angle \u03b8 can be written as where Se,Se4 represent the shape function matrices, which can be written as Se = [ Se1 Se2 Se3 ] , Se1 = [0 s1 0 0 s2le 0 0 0 s3 0 0 s4le 0 0], Se2 = [0 0 s1 0 0 s2le 0 0 0 s3 0 0 s4le 0], Se3 = [0 0 0 s1 0 0 s2le 0 0 0 s3 0 0 s4le], (21) Se4 = [1 \u2212 x/le 01\u00d76 x/le 01\u00d76], (22) where s1 = 1 \u2212 3(x\u0304/ le) 2 + 2(x\u0304/ le) 3, s2 = x\u0304/ le \u2212 2(x\u0304/ le) 2 + (x\u0304/ le) 3, s3 = 3(x\u0304/ le) 2 \u2212 2(x\u0304/ le) 3, s4 = (x\u0304/ le) 3 \u2212 (x\u0304/ le) 2, (23) and the element nodal coordinate vector reads pe = [ \u03b8j1 (\u03c1j1) T (\u2202\u03c1/\u2202x\u0304)T j 1 \u03b8j2 (\u03c1j2) T (\u2202\u03c1/\u2202x\u0304)T j2 ]T "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003973_s12206-012-0811-y-Figure5-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003973_s12206-012-0811-y-Figure5-1.png",
        "caption": "Fig. 5. TARGE 170 element geometry.",
        "texts": [
            " The assembly consists of four components and there are three contact pairs. The elements used in the analysis are SOLID 92 (Fig. 3), which has quadratic displacement behavior and is well suited to model irregular meshes such as produced from CAD data. The element is defined by ten nodes having three degrees of freedom at each node. The element has plasticity, creep, swelling, stress stiffening, large deflection and large strain capabilities. The elements used to define the four contact pairs are CONTA173 (Fig. 4) and TARGE170 (Fig. 5). CONTA173 is normally used to represent contact and sliding between 3-D \"target\" surfaces defined by TARGE170 and a deformable surface, defined by this element. The element is applicable to 3-D structural and coupled field contact analyses. This element is located on the surfaces of 3-D solid or shell elements without mid side nodes. It has the same geometric characteristics as the solid or shell element face with which it is connected. Contact occurs when the element surface penetrates one of the target segment elements on a specified target surface"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003571_s0263574711000324-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003571_s0263574711000324-Figure2-1.png",
        "caption": "Fig. 2. Robotic mechanism.",
        "texts": [
            " The vector equation of motion of all the moments involved for each mode\u2019s tip of any link is H d2y dt2 + h + jT e \u00b7 Fuk + \u2666 \u00b7 \u00b7 \u03b5 + \u03b5 = 0 \u2223\u2223\u2223\u2223\u2223\u2223\u2223 \u2211 F=0( \u2211 M=0) at the tip of any mode of the link considered . (5) The mathematical model of all m motors can be written in a vector form as u = R \u00b7 i + CE \u00b7 \u02d9\u0304\u03b8 CM \u00b7 i = I \u00b7 \u00a8\u0304\u03b8 +Bu \u00b7 \u02d9\u0304\u03b8 \u2212 S \u00b7 (\u2666m \u00b7 \u03b5 + \u03b5m) \u2223\u2223\u2223\u2223\u2223\u2223 \u2211 M=0 about the rotation axis of the each motors . (6) Example. Let us analyze the behavior of the robotic pair consisting of elastic gear and flexible link in the presence of the second mode, as depicted in Fig. 2. The link has two modes (the lower one and the upper one) and each of them is considered as a mode of rectangular cross-section ci,j xbi,j . The presence of the second mode is introduced into the analysis of the robotic pair behavior. The relations between the important angles are defined in Fig. 3. q = \u03b8\u0304 + \u03be + \u03d11,1, \u03b3 = \u03b8\u0304 + \u03be, \u03b4 = \u03d11,2 + \u03c91,1, \u03c91,j = \u03d11,j 2 . (7) The dynamic model (both the model of flexible line and model of the motion of each mode\u2019s tip) is defined according to classical principles but with the previously introduced new DH parameters, using Lagrange\u2019s equations",
            " (8) The magnitude ri,j is the maximal deflection, i.e., the deflection at the each mode\u2019s tip. The bending angles are expressed in Fig. 3 via generalized coordinates, Eqs. (7) and (8) r1,1 = l1,1 \u00b7 (q \u2212 \u03b3 ), r1,2 = l1,2 \u00b7 ( \u03b4 \u2212 q \u2212 \u03b3 2 ) , (9) m\u0302el1,1 = 33 140 \u00b7 w\u03041,1 \u00b7 (x1,1 \u2212 x\u03021,1), m\u0302el1,2 = 33 140 \u00b7 w\u03041,2 \u00b7 (x1,2 \u2212 x\u03021,2), (10) J\u0302elzz1,1 = m\u0302el1,1 \u00b7 ( x1,1 \u2212 x\u03021,1 2 )2 , J\u0302elzz1,2 = m\u0302el1,2 \u00b7 ( x1,2 \u2212 x\u03021,2 2 )2 . (11) Equation (10) sources from ref. [22]. Kinetic and potential energies of the mechanism presented in Fig. 2 are denoted as \u02c6\u0302Ekm and \u02c6\u0302Ep. All the angles in the expression for kinetic and potential energies characterizing flexibility of the links should also be expressed via generalized coordinates. The bending moment is expressed at any point of mode in the form \u03b5\u0302i,j = \u03b2i,j \u00b7 \u22022(y\u0302i,j +\u03b7i,j \u00b7 \u02d9\u0302yi,j ) \u2202 x\u03022 i,j . So that total potential energy is \u02c6\u0302Ep = Epo + Epels + Epel\u03be . (12) So that total dissipative energy is = els + el\u03be . (13) Potential, dissipative energy as a result of elasticity of the first and second link is Epels = 1 2 \u00b7 Cs1,1 \u00b7 r2 1,1 + 1 2 \u00b7 Cs1,2 \u00b7 r2 1,2, els = 1 2 \u00b7 Bs1,1 \u00b7 r\u03072 1,1 + 1 2 \u00b7 Bs1,2 \u00b7 r\u03072 1,2 on the top of each link",
            " This is just the procedure for obtaining Euler\u2013Bernoulli equation by which the motion of any point on the flexible line of the first mode is performed [H\u03021,1 H\u03021,2 H\u03021,3 0] \u00b7 \u03c6\u0308 + h\u03021 + J\u03021,1 \u00b7 Fuk x + J\u03022,1 \u00b7 Fuk y \u2212 1 2 \u00b7 Cs1.2 \u00b7 l1,2 \u00b7 r1,2 \u2212 1 2 \u00b7 Bs1,2 \u00b7 l1,2 \u00b7 r\u03071,2 + \u03b21,1 \u00b7 \u22022(y\u03021,1 + \u03b71,1 \u00b7 \u02d9\u0302y1,1) \u2202 x\u03022 1,1 = 0. (19) Fuk (N) is the dynamic force of the contact (in this case). The component of the entire external force in the radial direction (see Fig. 3) is Fc = (me \u00b7 \u0308 + be \u00b7 \u0307 + \u00b7Fo c + ka1 \u00b7 ), whereas the friction force is Ff = \u2212\u03bc p\u0307s |p\u0307s | \u00b7 Fc, as in ref [23]. The friction coefficient is \u03bc. The velocity of the robot tip is p\u0307s . is the distance from the point \u201c0\u201d to the trajectory, marked with \u03bb on Fig. 2, and = l \u2212 , l = l1,1 + l1,2. me (kg) is the equivalent mass, be (N/(m/s)) is the equivalent damping, ka1 (N/m) is the equivalent rigidity. \u03c2 = C\u03be \u00b7 \u03be + B\u03be \u00b7 \u03be\u0307 is the elasticity moment of gear and \u03b5i,j = (Csi.j \u00b7 ri,j + Bs i,j \u00b7 r\u0307i,j ) \u00b7 li,j is the bending moment of each mode\u2019s tip motion \u03c6 = [q \u03b4 \u03b3 \u03b8\u0304]T, H\u03021,1 = m\u0302el1,1 \u00b7 (x1,1 \u2212 x\u03021,1)2 + (m + mel1,2) \u00b7 l2 1,1 + (m + mel1,2) \u00b7 l2 1,2 + 2 \u00b7 (m + mel1,2) \u00b7 l1,1 \u00b7 l1,2 \u00b7 cos \u03b4 + 9 4 \u00b7 J\u0302elzz1,1 + 9 16 \u00b7 (Jzz + Jelzz1,2), H\u03021,2 = . . . , H\u03021,3 = ",
            " (5) and (6)) y = a(xi,j , Tsti,j , Ttoi,j , \u03b8\u0304 , \u03b1, t), x = b(xi,j , Tsti,j , Ttoi,j , \u03b8\u0304 , \u03b1, t), z = c(xi,j , Tsti,j , Tsti,j , \u03b8\u0304 , \u03b1, t), \u03c8 = d(xi,j , Tsti,j , Ttoi,j , \u03b8\u0304 , \u03b1, t), \u03be = e(xi,j , Tsti,j , Ttoi,j , \u03b8\u0304 , \u03b1, t), \u03d5 = f (xi,j , Tsti,j , Ttoi,j , \u03b8\u0304 , \u03b1, t). (33) From Eq. (33), the motion of each mode\u2019s tip and link\u2019s tip can be calculated and finally of the robot tip\u2019s motion. Example. In order to define the shape and position of elastic line of the first and second mode link from Fig. 2, during the realization of robot\u2019s task in the space of Cartesian coordinates, it is necessary to find the solution Eqs. (19), (21), (24), and (25). The general solution of the dynamics movement of the observed model is given y\u0302 = a\u0302(x\u0302i,j , T\u0302sti,j , T\u0302toi,j , \u03b8\u0304 , \u03b1, t), x\u0302 = b\u0302(x\u0302i,j , T\u0302sti,j , T\u0302toi,j , \u03b8\u0304 , \u03b1, t), \u03c8\u0302 = d\u0302(x\u0302i,j , T\u0302sti,j , T\u0302toi,j , \u03b8\u0304 , \u03b1, t). (34) Furthermore, the position and orientation of any point on the link elastic line during the robot\u2019s task realization need to be defined",
            " (31) or the form of equation of motion solutions of robot tip Eq. (33)) and the procedure of the \u201cdirect kinematics\u201d Eq. (39) and \u201cinverse kinematics\u201d solutions Eq. (41) in Robotics In this way, the complete analogy between Eqs. (28), (39), and (41) is established. The analogue between the \u201cOriginal form of the Euler\u2013Bernoulli equation\u201d and its solution and modern knowledge from Robotics is presented in this way. Tip of robot started from the position \u201cA\u201d and moves directly to the point \u201cB\u201d in the predicted time of T = 2 (s), (see Fig. 2). The trapezoidal profile of velocity together with time of acceleration and deceleration from 0.2 \u00b7 T is adopted. The characteristics of stiffness and damping of gear in the real and reference regimes are not the same nor are stiffness and damping characteristics of the link C\u03be1 = 0.99 \u00b7 Co \u03be1, B\u03be1 = 0.99 \u00b7 Bo \u03be1, Cs1,1 = 0.99 \u00b7 Co s1,1, Bs1,1 = 0.99 \u00b7 Bo s1,1, Cs1,2 = 0.99 \u00b7 Co s1,2, Bs1,2 = 0.99 \u00b7 Bo s1,2. The only disturbance in the system is the ignorance of the rigidity characteristics and damping"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000094_jp073212s-Figure7-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000094_jp073212s-Figure7-1.png",
        "caption": "Figure 7. CVs of bare GCE (a), CaP-PDDA (b), Hb/CaP (c), and Hb/CaP-PDDA (c\u2032) film-modified GCE in PBS (pH 7.0). Scan rate: 100 mV/s.",
        "texts": [
            " For further comparison, we tested the UV-vis spectra of hemin, a free heme group, the Soret band of which was observed at 388 nm in solution (curve d). The obvious blue shift of the Soret band of hemin compared with Hb on CaP and CaP-PDDA-modified ITO glass slides suggested that the heme prosthetic group of Hb does not split out from Hb polypeptide matrix.50 Moreover, Hb/CaP-PDDA film showed a higher absorbency than Hb/CaP film, indicating that the composite film has a stronger ability to immobilize proteins. Direct Electron Transfer of Hb on CaP-PDDA-Modified GCE. In pH 7.0 PBS, neither bare (Figure 7 curve a) nor CaPPDDA-modified GCE (curve b) showed any redox peaks. However, Hb/CaP-PDDA (curve c) displayed a pair of welldefined redox peaks at Epc ) -0.403 V, Epa ) -0.303 V (vs SCE), which were in accordance with the characteristic of FeIII/ FeII redox couples of heme proteins. The shapes of the reduction and oxidation peaks were nearly symmetric. The formal potential, E0\u2032, as the average of oxidation and reduction peak potentials, was -0.353 V. When Hb was immobilized on CaP film, a pair of redox peaks was also observed (curve c\u2032), but the response was smaller than that of the Hb/CaP-PDDAmodified electrode"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001884_s11431-010-3100-y-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001884_s11431-010-3100-y-Figure1-1.png",
        "caption": "Figure 1 The PCM mechanism.",
        "texts": [
            " [17]: 1 1 ,      P L i j i j F f d (1) where fi is the number of freedoms of the ith kinematic pair, P is the total number of kinematic pairs, L is the number of independent loops, and di is the rank of the jth independent loop. Early in 1954, Moroskine gave a similar formula M = m i=1fir [18], where m is equal to P, and r =L j=1dj . In Gogu\u2019s recent papers [5, 16], eq. (1) was often considered not to be general. In ref. [5], Gogu analyzed the PMC mechanism with 12 lower-mobility pairs, as shown in Figure 1. For the two independent loops, he obtained that each rank was 5. So according to eq. (1), erroneous mobility of F=12-(5+5)=2 was obtained. In addition, he also set up the kinematic constraint equations, and got the rank of each loop from solving a 12\u00d712 matrix. The rank of each loop was 3. From eq. (1) again, another correct result of F=12-9 =3 was obtained. So he said, \u201cWe have to review the formulas and to consider the flaw to be that of the formula rather than of the mechanism\u201d. In ref. [14], we deduced eq",
            " All links in loop I cannot rotate about x-, y-axes, so dX I = dX I ( 0 0, x y z) = 4, FI = PI i=1fi dX I =64=2, d2,7 I ( 0 0, x y z)=4. In loop II, on account of the displacements caused by link 4, the rank of the link group EFG is dgz II (,x y z)=6, dX II =d2,7 I ( 0 0, x y z)+dgz II (, x y z)= dX II (, x y z)=6, FII = PII i=1 fi dX II =56=1. In loop III, the rank of the link group KH is dgz III (0 0 0, 0 y z) =2, d5,7 II (0 00, 0 y 0)=1. dX III = d5,7 II (0 0 0, 0 y 0)+dgz III (0 0 0, 0 y z)=dX III (0 0 0, 0 y z)=2, FIII =  PIII i=1fi dX III =22=0, F=FI +FII +FIII = 21+0=1. From eq. (4), we get F=13(4+6+2)=1. Example 3. Figure 1 illustrates the PCM mechanism with eleven links. In loop I, two RRRP chains form a link group. All R-, P-axes are parallel to plane x0O0y0. There exists a common constraint around z0-axis. Therefore, dX I =dX I (, x y z)=5, FI =  PI i=1fi dX I =85=3. In addition, the two revolutions in the moving platform 4 are always parallel to plane x0O0y0 , so the moving platform 4 has only a translating motion, the rank of the generalized pair including the moving platform and the base is dA,B I (0x y z)=3"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003761_9781118818428-Figure7.2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003761_9781118818428-Figure7.2-1.png",
        "caption": "Figure 7.2 represents a cylinder closed at its upper end and sealed from below by a piston. There are no ports, so total mass of gas processed, M, is invariant. The piston is driven at uniform angular speed by a \u2018real\u2019 mechanism, that is, motion is not confined to being simple-harmonic. The piston stroke is s, and axial clearance at inner dead centre position is c, so that compression ratio rc is 1 + s\u2215c. Allowing for horizontal offset (de\u0301saxe\u0301) e between axes of crankshaft and cylinder will permit eventual examination of asymmetry of expansion and compression processes.",
        "texts": [
            " This can be achieved by working in terms of fraction distance y\u2215Lwhere L is a length characteristic of local geometry. The obvious choice for internal flows is hydraulic radius rh. The resulting group is Nusselt number NNu: NNu = hrh\u2215k (7.10) The evolving treatment remains short of representing conditions in the variable-volume spaces. Cardinal features of the latter are: (a) cyclic variation of volume \u2013 and pressure \u2013 and (b) motion of the gas relative to the cylinder surface. A step in the right direction can be made with the aid of diagrams at the right-hand side of Figure 7.2. These show a slice through the axis of the vertical cylinder and an element of gas representative of conditions at radius r. All elements are subject to common instantaneous pressure p. The assumption of symmetry about the z axis means that, at any instant, temperature is uniform in the circumferential (\u03b8) direction. The unsteady energy equation for an element is: q\u2032 \u2212 w\u2032 = u\u2032 Net rate of energy accumulation due to thermal diffusion in the radial direction is: qr \u2032 = kr\u0394\u03b8\u0394z{r(\u22022T\u2215\u2202r2) + (\u2202T\u2215\u2202r)}\u0394r In the z direction: qz \u2032 = kr\u0394\u03b8\u0394r(\u22022T\u2215\u2202r2)\u0394z For any given element, dm calculated from initial conditions p0 and T0 is invariant"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000856_bjsm.2007.035915-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000856_bjsm.2007.035915-Figure1-1.png",
        "caption": "Figure 1 (a) Joint angles and (b) segment angles used for biomechanical analyses.",
        "texts": [
            "22 Allometric equations describe curvilinear relationships between physiological values, which take the general form: y = a 6xb, where x is lean body mass (LBM) and y VO2peak at submaximal running speed.19 20 Values of a and b can be derived from linear regression on logarithmic transformation of data by the formula lny = lna + b 6 lnx. When visualising the process of scaling by plotting oxygen uptake against LBM, the slope of the curve should be close to zero (parallel to the x axis).33 Biomechanics Before each test, 11 landmarks were placed on the lateral part of the body (ear, neck, shoulder, elbow, wrist, hip, knee, ankle, heel, metatarsal V and thoracal XII; fig 1). Running was filmed laterally at 50 Hz. The following kinematic variables were calculated in one stride: shank angle at contact, trunk angle, maximum plantar flexion angle, maximum knee flexion in support, minimum knee velocity, wrist excursion, vertical oscillation and minimal knee angle during swing phase (fig 1).25 Mass moment of inertia (I) of the leg in swing phase was calculated according to Steiner\u2019s Law. Partial mass and position of partial centres of gravity were corrected for age30 and used to calculate the mass moment of inertia. Stride frequency (SF) and stride length (SL) were derived from the video recordings. SL, the length from one foot contact to the next contact of the same foot, was calculated by SL = treadmill speed/SF. Statistics All data were stored and analysed using SPSS-PC V.12.0 (SPSS Inc, Chicago, Illinois, USA)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002469_s11804-012-1144-z-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002469_s11804-012-1144-z-Figure2-1.png",
        "caption": "Fig. 2 Body-frame and earth-fixed reference coordinate system",
        "texts": [
            " Therefore, the high frequency components should be filtered in the estimation of the location of the platform, and it is only necessary to consider the low frequency motion caused by wind, current, and second-order waves. Dynamic positioning is concerned primarily with control of the platform in the horizontal plane involving surge, sway, and yaw. In order to illustrate the motion of the platform in a complex ocean environment, two coordinates systems should be established: the body-frame coordinates O-XYZ and earth-fixed-reference coordinates OE-XEYEZE, see Fig. 2. O-XYZ coordinate system is located in the horizontal plane, axis OX points to the platform heading, axis OY points to starboard and axis OZ is vertical to the horizontal plane and points to the center of the earth. OE-XEYEZE coordinate system is located in static horizontal plane, axis OEXE points to the north, axis OEYE points to the east and axis OEZE points to the center of the earth. The three axes of the body-frame coordinates correspond to three main rotation axes of the platform. The origin of body-frame coordinates is usually found in the symmetry axis of the platform"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001465_robot.2010.5509554-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001465_robot.2010.5509554-Figure2-1.png",
        "caption": "Figure 2. A general case of the three stages for the secondary trajectory to",
        "texts": [
            " A secondary subtask representing the best position and orientation of the wheelchair is represented as a secondary set of trajectories for the wheelchair to follow. An optimal position/orientation combination of the non-holonomic motion of the wheelchair can be achieved if the secondary trajectory is divided into 3 stages. The first one is to orient the wheelchair facing its desired linear trajectory. The second stage is to proceed with a linear motion along the secondary trajectory to approach the final planar coordinates. Once the wheelchair reaches its final position, the third stage will be to orient the wheelchair to its final desired orientation. Figure 2 shows the three stages implemented for the secondary trajectory. be followed by the wheelchair The three stages to be applied for the secondary trajectory will only involve the position \u201cX\u201d and orientation \u201c \u03d5 \u201c variables of the wheelchair. As shown in figure 3, knowing the initial and final transformations of the wheelchair base, the trajectory angle \u03b1 can be defined as: [ ]f i f ia tan 2 (y y ) , (x x )\u03b1 = \u2212 \u2212 (10) That defines the amount of motion needed for the three stages to be followed in the following order: 1) Rotation by the amount of 1 i\u03b2 = \u03b1 \u2212 \u03d5 2) Translation by the amount of 2 2 f i f itr (x x ) (y y )= \u2212 + \u2212 3) Rotation by the amount of 2 f\u03b2 = \u03d5 \u2212 \u03b1 The above three wheelchair motion values can be utilized in the weight matrix as criteria to enforce the wheelchair motion"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002871_pime_conf_1966_181_306_02-Figure6.3-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002871_pime_conf_1966_181_306_02-Figure6.3-1.png",
        "caption": "Fig. 6.3. Co-ordinate system and finite dterence mesh",
        "texts": [
            " Finite difference representation of the energy equation The relaxation method is not well suited for the solution of the energy equation when temperature variation across (z) and along (x) the film has to be considered. In this work a marching procedure developed for the solution of thrust-bearing problems (equations (6.2) and (6.3)) has been employed. The only difference from the previous procedure is that the programme allows for a variable mesh size in the x-direction. The mesh and co-ordinate system used to establish the finite difference form of the energy equation is shown in Fig. 6.3. The increments in the .%direction are constant and equal to 62. If the number of increments is J, then St = l/J. The increments in the ?-direction are not constant and must be defined in relation to the values of f at surrounding stations, thus Sf\" = [%In- [TI.- [6fIn+l = [%]n+l-[f]n, etc. Proc Insrn Mech Engrs 1966-67 With this notation the energy equation (6.2) can be written in the following finite difference form : + 1 T,\",' +Bjn+ 1 T;,+,' +Cjn+l T j n + 1 + o j n + 1 Tjn + E j n + 1 Tjn- +Fin + 1 = 0 (6",
            "17) Proc Instn Mech Engrs 1966-67 The terms representing convection across the film, conduction along the film, and adiabatic compression turn out to be relatively small for the bearing geometry and operating conditions considered in this paper, but they are retained for completeness. In an overall sense the heat balance for the bearing indicates the relative importance of the convection and conduction terms, but this cannot be applied locally. In dimensionless form, . (6.2) where - T T = - T, The temperature gradients in equation (6.2) can be written in finite difference form with reference to the mesh shown in Fig. 6.3 as follows : 7 p r [ 1 +xl Y,\" + Pl - \" + 1c;; 1 w,n + '1 + T,=+r[ 1 -x, Y,\" + l j r n + 14; 1w,n + '1 + Trn+ \" - 2 + x 4 y2-+ 1 \u20ac1 - n + I,,, + 1 +X3Y,\" + 1 2 5 Pr P1 u1 1 -x2y3a+ 1 PI -.+lo+ P f I u1 - n + 11 + Tjn[x3 y 7 n + 1 + x 2 y 4 n + 1 - n + 10 + 1 - n + 1 + S ; n - 1[x3 y e n + 1 - X y n + 1 - n + lo+ 1 - n + I P9 Pj u, 1 +X, - . . . . . . . (6.18) since and Sf\" 62\" + 1 H2(2Sfn+' + 6R\") p + 1 SZ.\" + l( Sf\" + 1 + 8%\") y3n + = K, L, and M being functions of the increments in the x-direction and the dimensionless thermal conductivity"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000587_2007-01-2232-Figure12-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000587_2007-01-2232-Figure12-1.png",
        "caption": "Fig. 12 Torsional Damper Characteristics",
        "texts": [],
        "surrounding_texts": [
            "As a result of these countermeasures, the hybrid luxury sedan achieves a level of quietness vastly superior to that of ordinary gasoline-powered vehicles in the same class (Figs. 5 and 6).\nGenerally, in vehicles with longitudinal engines, vertical and lateral vibrations of the vehicle body are generated during engine start in conjunction with the roll and yaw resonance of the power plant. The same sort of vibration occurs in the hybrid luxury sedan with THS II. Also, because there is no clutch mechanism between the engine and the drive train, the torque fluctuation of the engine excites torsional resonance in the drive train, generating longitudinal vibration in the vehicle body. The compelling forces of these forms of vibration are the motor torque reaction force and engine compression pressure during motoring, as well as the combustion pressure during explosion. The following section describes technologies for reducing these compelling forces, as well as technologies for reducing vibration in the power plant mounting system and drive train.\nA variety of countermeasures that were developed for the THS II for FWD vehicles have been implemented in the hybrid luxury sedan.(1)-(3) The compelling forces were reduced by changing the intake valve closing timing, by controlling the piston stop position to reduce the engine compression pressure, and by adjusting the injected fuel volume and ignition timing.\nFig. 7 Motor Control System for Vibration Reduction\nThe lateral vibration of the vehicle body was reduced by shortening the distance between the principal elastic axis and the center of the gravity of the power plant. The excitation of torsional resonance in the drive train and resonance in the power plant mounting system was inhibited by operating MG1 at high torque during engine start. In the hybrid luxury sedan, in addition to the countermeasures described above, vibration-reducing motor control and the torsional damper, which has twostage hysteresis characteristics, were adopted to reduce the drive train torsional vibration. (4)-(6)\nVibration-reducing motor control is implemented in the form of feed-forward control by MG1 and feedback control by MG2 (Fig. 7). Because the THS II power split device uses a planetary gear, the revolution speeds of MG1, the engine, and the drive train axis can be expressed by the nomographic chart in Fig. 8. The MG1 feed-forward control estimates the torque fluctuation generated in the crankshaft based on the crank angle. Fluctuation in the revolution speed of the drive train axis is reduced by applying feed-forward compensation to the torque that is in phase with MG1 (Fig. 7 and 8). MG2 feedback control estimates the torsional angular velocity of the drive shaft based on the revolution speeds of the wheels and MG2 in relation to the torque fluctuation that is generated abruptly at the time of explosion in the engine. The longitudinal vibration of the vehicle body is reduced by feeding back the out-of-phase torque to MG2 (Fig. 9).\nFig. 9 Vibration Reduction with Motor Control\n-0.5 0 0.5 1 1.5 Time (s)\nGasoline engine\nTDC signal\nMG2 feedback control\nCrank angle calculation Engine torque fluctuation estimation\nConventional controller\nFilter\nMG1 feed-forward control\nMG1\nMG2\nAngle Angle\nWheel speed estimationGain\nMG2 Speed 1200\n800\n400\n0\nVibration control: OFF\nMotoring start Ignition start\nF lo\nor a\ncc el\ner at\nio n\n(m /s\n2 )\nE ng\nin e\nsp ee\nd (r\npm )\nVibration control: ON\nTorsional angular velocity",
            "The torsional vibration level of the drive train can also be reduced by increasing the damping of the torsional damper. Increasing the hysteresis torque is an effective way to increase damping, but it raises the concern of worse booming noise of engine explosion first-order component because the torque fluctuation to the drive train increases. In consideration of the different torsional angles of these two phenomena, a damper with twostage hysteresis characteristics was adopted (Fig.10). During engine start, the hysteresis torque is high in the high-amplitude region, but the hysteresis torque is low in the very low-amplitude region where booming noise tends to be a problem (Figs. 11 and 12).\nIn a vehicle powered by a V6 engine, in addition to the third-order component of engine revolution, the booming noise of the second-order component of engine revolution sometimes becomes a problem. The secondorder excitation force of revolution is mainly a secondorder couple of the reciprocating inertia of the piston. The vibration is amplified by the bending resonance of the power plant and is transmitted to the body from the mounts at the rear of the transmission, generating booming noise. A two-speed reduction gear was adopted for the RWD hybrid sedan, which made it possible to reduce the motor diameter and weight. But the overall length of the THS II transmission is 50 millimeters greater than that of a 6-speed automatic transmission (A/T) in an ordinary gasoline-powered vehicle on the same platform (Fig. 13), and the mass is approximately 35 kilograms greater.\nThere was concern that the increase in mass and overall length would cause the power plant resonance to drop into the normal engine speed range. An increase in the transmitted force was also predicted, because the increase in the load allocated to the mounts would require that the spring constant for the mounts be raised to 1.5 times that of the mounts for a 6-speed A/T.",
            "The deformation mode of the power plant resonance is such as to cause large deformation in the transmission, as shown in Fig. 14. A study was undertaken of how to improve the resonant frequency so as to separate it from the normal engine speed range. FEM analysis was used to optimize the shape of the transmission case by smoothing the outline, reinforcing the ribs, and so on. As a result, the resonant frequency was raised to 180 Hz, equivalent to that of a 6-speed A/T. Also, the mounting position was shifted 80 millimeters farther forward than in the original plan in order to set it at a nodal point of the vibration mode (Fig. 15). Changing the mounting position raised new issues of installation space and the separation of principal elastic axis and the center of gravity, which strongly influences the engine start vibration performance. In order to ensure the adequate installation space, the mounting is embedded to the cross member, a major change from the structure used for a 6-speed A/T (Fig. 16). The distance of the principal elastic axis and the center of gravity was reduced by optimizing the lateral-to-vertical ratio of the mount spring constants. These countermeasures successfully addressed the issues of booming noise and engine start vibration (Fig. 17).\nThe overall length of the transmission of the hybrid luxury sedan is greater than that of the THS II for FWD vehicles (Fig. 18), so the resonance that deforms the entire transmission is generated at a comparatively low frequency. Also, since the MG2 reduction gear ratio is low (Table 1), the resonance is generated by the 24thorder component of MG2 speed in the low vehicle speed range where there is less background noise (Fig. 19), which means that the motor noise is readily audible. Moreover, the transmission is installed in the center tunnel, which makes the acoustic transfer function from the transmission to the occupants higher than in a FWD hybrid vehicles, where the transmission is mounted in the engine compartment (Figs. 20 and 22). The following section explains countermeasures that were employed."
        ]
    },
    {
        "image_filename": "designv11_12_0001963_j.triboint.2008.12.005-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001963_j.triboint.2008.12.005-Figure2-1.png",
        "caption": "Fig. 2. FE model of a thermal actuated slider constructed in ANSYS.",
        "texts": [
            " The high complexity of the air flow and heat transfer across the HDI makes it non-trivial to solve the system as a whole numerically. The coupled-field analysis includes a slider FE model, an air bearing model and a heat transfer model. Fig. 1 shows the diagram of the HDI and models used in the study. The details of these models are given in the following sections. A three-dimensional FE model of an entire femto-sized slider, which includes the read/write element, substrate, upper and lower poles, write coil, photoresist, upper and lower shield and heater, is created as shown in Fig. 2. The heater has a thickness of 80 nm and is located between the Al2O3TiC edge and the lower shield. The thermo-mechanical properties of these slider elements are listed in Table 1. These values are process-dependent and we used the data published in Ref. [6]. The temperature distributions and thermal protrusion profiles were obtained by performing thermal-structural coupled field FE analyses using commercial software ANSYS [14]. For thermal boundary conditions of the slider, we set heat transfer coefficient as a constant of 100 W/m2 K on non-ABS surfaces of the slider and the heat transfer coefficient on the ABS will be calculated by applying our air bearing and heat transfer models"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003465_isiea.2012.6496635-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003465_isiea.2012.6496635-Figure1-1.png",
        "caption": "Fig. 1 1-unit Cubesat with magnetic actuators.",
        "texts": [
            " Substituting (7) into (6) and the resultant projected torque, known as the optimal torque, can be obtained: ( ) ( )opt f f MT2 f 1 S S= \u2212U \u0392 \u0392 U \u0392 (8) The close proximity of the magnetic actuators would result in disturbance to the magnetic field readings obtained by the magnetometers. In a practical implementation of an attitude control system, the magnetic actuators need to be turned off for a duration to allow the magnetic flux to dissipate before taking the magnetic field readings. To determine the effects of the magnetic actuators on the magnetometer, four magnetic actuators consisting three air coils and a torque rod are mounted onto a 1-unit Cubesat satellite fabricated in-house to complete a 3-axis magnetic actuation system as shown in Fig. 1. Fig. 2 shows an experimental result using the setup of Fig. 1. For this experiment, the magnetic actuators were turned on at 1s and turned off at 3s. It can be observed from Fig. 2 that when the magnetic actuators are turned on, there is an offset to the magnetometer readings denoted as \u2206Bf which are recorded in Table 1. This offset error is propagated into (8) and we have ( ) ( )f f opt f MT2 f f S S \u239b \u239e+\u239c \u239f= \u2212 \u239c \u239f+\u239d \u23a0 \u0392 \u0394\u0392 U \u0392 U \u0392 \u0394\u0392 (9) The error in the control action can be obtained using the difference between (8) and (9): ( ) ( ) ( )f f f opt f MT2 2 f f f S S S \u239b \u239e+\u239c \u239f\u0394 = \u2212 \u2212 \u239c \u239f+\u239d \u23a0 \u0392 \u0392 \u0394\u0392 U \u0392 U \u0392 \u0392 \u0394\u0392 (10) The averaged Earth geomagnetic field for LEO at an altitude of 600km is approximately 350 mGauss"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001724_tmag.2010.2072910-Figure6-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001724_tmag.2010.2072910-Figure6-1.png",
        "caption": "Fig. 6. Loci of . (a) Point A. (b) Point B. (c) Point C.",
        "texts": [
            " 5 shows the iron-loss distributions calculated from the vector relation between and . Usually, the iron loss is assumed to be in proportion with the square of the maximum magnetic flux density, . Actually, the iron loss cannot be determined with only . In this analysis, it was clarified the iron loss increased at parts where both and were large. Generally, the waveforms of the magnetic flux density and the magnetic field intensity in the stator yoke of the induction motors include secondly slot harmonic components. Fig. 6 shows the loci of at each point as an example. As shown in Fig. 6(a)\u2013(c), loci of were distorted, and the distortion was very large at the tooth top [Fig. 6(b)]. The influence of the secondly slot harmonic was very small at the core back part [Fig. 6(c)]. Fig. 7 shows the results of the fast Fourier transform (FFT) at point B indicated in the Fig. 6. From these results, it was shown that not only the magnetic flux density but also the magnetic field intensity was distorted. Moreover, it can be concluded that the magnetic flux and field intensity waveforms include about 30% of the harmonic components in the maximum at the tooth top parts near the air gap. As a result, we point out that the influence of the higher harmonics components cannot be neglected in the magnetic field analysis of induction motors. To accurately consider the higher harmonic components caused by influence of secondly slot construction, it is necessary to use dense mesh that can be neglected effects of the mesh size on numerical solutions or the sliding mesh and so on"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000439_s11047-008-9101-0-Figure5-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000439_s11047-008-9101-0-Figure5-1.png",
        "caption": "Fig. 5 PTRL-net model of a dynamic membrane system. For simplicity, all isolated empty places have been omitted. Transitions are labelled with the names of evolution rules from which they were derived. Lines with a small white (black) circle at one end denote range arcs with the weight [0,0] (resp. \u00bd1;1 )",
        "texts": [
            " In particular, we can drop those ti;r;~l;/;w for which there is a node j 2 ~l n fig such that no evolution rule r 2 R\u00f0j\u00de has the status rr = d. We may also omit an inhibitor arc adjacent to a status place if the latter is guaranteed never to be filled with a token. In particular, if no evolution rule in R can dissolve a membrane, then the reduced translation generates exactly one transition for each evolution rule of P. After adopting these simple observations, the translation for the dynamic membrane system from Fig. 3 is that shown in Fig. 5. As a first observation on this construction we note that the status places associated with the root remain empty during any run of NP, thus correctly reflecting that the outer (skin) membrane is never dissolved or thickened. This is an immediate consequence of Definitions 3 and 13. Proposition 3 If M is an m-reachable marking of NP for any execution mode m, then M(s,lroot) = M(d,lroot) = 0. To establish the behavioural equivalence of P and NP, we first capture the correspondence between configurations and markings and between the m-enabledness of vector multi-rules and steps"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003521_10402004.2011.639048-Figure7-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003521_10402004.2011.639048-Figure7-1.png",
        "caption": "Fig. 7\u2014Pressure distribution and film thickness profiles at different points along the line of action: (a) p \u2212 x curve and (b) h \u2212 x curve.",
        "texts": [
            " The maximum pressure corresponds to the film thickness bump in Fig. 5. These behaviors can be better understood by studying the pressure distributions and the film thickness profiles (see Fig. 6), which is caused by a squeeze effect. The central pressure in Fig. 4 was smooth but the maximum pressure and minimum film thickness was oscillating in Fig. 5 because the position of the pressure spike in the pressure distribution and the film thickness necking in film thickness profiles changed relative to the spatial mesh used for the problem. Figure 7 depicts the transient pressure distributions and film thickness profiles at five specific meshing points, the approach point A+, point B+ after the load is carried by two pairs of teeth transit into by only one pair, the pitch point C, point D+ after the load is carried by only one pair of teeth transit into by two pairs, and the recess point E\u2212. We can see that the minimum film thickness occurs at the approach point and the maximum film pressure occurs at point B+. The impact load has a strong influence on the gear transmission"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001523_tpas.1972.293335-Figure3-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001523_tpas.1972.293335-Figure3-1.png",
        "caption": "Fig. 3. The a -13 Reference Frame in Relation to the Natural Reference Directions.",
        "texts": [
            " It is a normal practice to transform the voltage and the flux linkage equations so that all variables are expressed in a common d - q reference frame. This has the merit of converting the standard steadystate a.c. problem into a d.c. time invariant problem. The present purpose - the derivation of the small perturbation equations - requires at this stage a more general transformation of the stator variables into an orthogonal freely rotating reference frame. Transformation to a Freely Rotating Frame (5) The a - A reference frame together with the fictitious a - stator windings is shown in relation to the other significant directions in Figure 3. The inductance matrix LI 2 is: -L -Lb cos29 Lb sin29 L cosO L cosg -L sinGad ad aq -Lbsin2g -La+Lbcos29 Lad sing Lad sinG cos -Lad cos -Lad sinG Lffd Lfkd 0 ad j ad s Lfkd Lkkd __O-L s -Lt sinIG L sing aq I cosG aq L + L with: L - a 2 Ld =L +L ador 0 Lkkq L -L Lb = 2 L = L + L q ar aq The electromagnetic torque opposing rotation results from the air-gap flux linking the stator windings while they are carrying currents i and i2: Te -=- - Oa2i + al i2 (6) It can be shown that the toruqe can also be expressed as: Te = i12(G12 + F12) j12 The transformation matrix Cop, which converts currents in the where: a -13 reference frame to currents in the 1 - 2 frame, is: (8) cos 0 sin ,0 -sin ,0 cos 0 0 0 0 0 0 0 0 0 o 0 0 0I 1 0 0 0 1 0 0 0 1J Power invariance requires that: vp = C V The transformation is more tractable if the voltage equation only is considered; after combining Equations (3) and (4) or (1) and (2), the voltage equation becomes: V12(Rm 1 2 +2 L12p i12 with G12 d L and Steady-State Operation If the quiescent operating point can be described by a set of constant RMS amplitude stator voltages and by a constant field voltage: v 2 V Cos (w t +G) v (17) (12) 0 then the freely rotating reference frame voltages become: V~3 = ~V cos ( X t + 0 - 0) ap V in (w t + G0- 0) Vfd _ (13) \u00b0 dm dt Pre-multiplying equation (13) by Ct and replacing i12 by C0 iv, the voltage equation becomes: ap = C, R1 2 C,0 p + m C0 G1 2 C0 iap +C L12 P (C0 ap i"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000680_jst.60-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000680_jst.60-Figure2-1.png",
        "caption": "Figure 2. Double pendulum model of the golf swing [7]. All movements occur in one plane and \u2018stop\u2019 prevents angle between upper and lower lever decreasing below 901. Reproduced by kind permission of Triumph Books.",
        "texts": [
            " The inverse dynamics solution typically involves the measurement of body movement and some forces (for example ground reaction forces), whereas input of the internal forces is required for the forward dynamics approach so that the output (body movement) can be calculated [2] (Figure 1). In order to gain insight into the basic mechanics involved in the golf swing and to explain the optimal coordination of the swing, Cochran and Stobbs [7] suggested a simple model of the downswing consisting of a double pendulum (Figure 2). They assumed that the two most relevant pivot points of the moving body segments were the wrist and a point \u2018roughly corresponding to the middle of the golfer\u2019s upper chest\u2019 [7]. This imaginary point is taken to be fixed in space and connected to an upper lever that is representative of the arms of the golfer. Another segment, representing the club, was connected to the upper lever via a hinge joint. This \u2018wrist\u2019 hinge was assumed to behave passively, restricted only by a stop that prevented the club segment from moving too far back at the top of the backswing"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001345_robot.2008.4543604-Figure4-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001345_robot.2008.4543604-Figure4-1.png",
        "caption": "Figure 4. (a) The backbone structure; (b) The backbone parameters",
        "texts": [
            " These sensors measure the contact with the load and ensure the distributed force control during the grasping. The sensor network is constituted by a number of impedance devices [11] (see Figure 3) that define the dynamic relationship between the grasping element displacement and the contact force. Nm B. Theoretical model The essence of the hyperredundant model is a 3- dimensional backbone curve C that is parametrically described by a vector 3Rsr and an associated frame whose columns create the frame bases (see Figure 4). The independent parameter s is related to the arclength from the origin of the curve C, , where: , where represent the length of the elements i of the arm in the initial position. 33Rs Ls ,0 N i ilL 1 il The position of a point s on curve C is defined by the position vector: srr , when . For a dynamic motion, the time variable will be introduced, Ls ,0 tsrr , . We used a parameterization of the curve C based upon two \u201ccontinuous angles\u201d s and [3-5] (Figure 4).sq The position vector on curve C is given by Ttsztsytsxtsr ,,,, (1) where s sdtsqtstsx 0 ,cos,sin, (2) s sdtsqtstsy 0 ,cos,cos, (3) s sdtsqtsz 0 ,sin, , ss ,0 (4) For an element dm, kinetic and gravitational potential energy will be: 222 2 1 zyx vvvdmdT , (5)zgdmdV where dsdm , and is the mass density. The elastic potential energy will be approximated by the bending of the element [10], N i iie qdkV 1 22 2 4 . III. DYNAMIC MODEL In this paper, the manipulator model is considered a distributed parameter system defined on a variable spatial domain L,0 and the spatial coordinate s"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002558_1.50800-Figure3-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002558_1.50800-Figure3-1.png",
        "caption": "Fig. 3 Object detection scheme: a) no detection and b) detection.",
        "texts": [
            " In this path planner, the obstacles and the target are modeled as ellipsoids of various sizes and orientations. Ellipsoids are also employed tomodel each UAV sensor range and communication range. These ellipsoids are centered upon the UAV center of mass and mimic the aircraft motion. 1. Object Detection Scheme Sensors onboard the UAV allow the aircraft to characterize its surrounding environment. An object (obstacle or target) is detected by the UAV when the ellipsoidal UAV sensor range intersects the ellipsoidal object [27] (Fig. 3, Table C1 in Appendix C). 2. Target Interception Scheme The UAVintercepts the target when its center of mass is contained within the ellipsoidal target [27] (Fig. 4, Table C1 in Appendix C). Themission is deemed completewhen all UAVs have intercepted the target. 3. Unmanned Aerial Vehicle Communication Scheme Information on objects (obstacles and target) detected by fellow UAVs that are in communication range is also relayed to the UAV. Two UAVs can communicate when the ellipsoids modeling their respective communication range encompass the other UAV center of mass [28] (Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000702_bf02441577-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000702_bf02441577-Figure1-1.png",
        "caption": "Fig. 1 Axisymmetric shell element",
        "texts": [
            " A radial section of the element only is Medical & Biological Engineering & Computing January 1977 67 necessary to describe the geometry by the use of local or natural co-ordinates, which have the values - 1 and 1 at the two ends. Shape functions, defined as giving unity at node i and zero at all other nodes, are used to determine the shell coordinates at any point in the element. Each node has three displacements (u, v, w along the shell or global co-ordinates) and a normal rotat ion (~). The element, with the two co-ordinate systems, is shown in Fig. 1. Global co-ordinates O, Z, R and local co-ordinates ~, ~/, ( are related by Z N , z, + E N , @ r RI where subscript i is used for nodes 1, 2 and 3, t is the thickness of shell wall, N is the shape function in ~/only and given as N1 = 89 N2 = 1 - r/2 N3 = 89 + e2) The displacement field inside the element is obtained from the nodal displacements and the shape functions. Displacements u, v and w at any point in the element are given by W WI P~, 17\"2 and 17\"3 are unit vectors along ~, ~/ and (, respectively, and can easily be related to unit vectors L ] and ~ along 0, Z and R"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002668_robot.2010.5509794-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002668_robot.2010.5509794-Figure2-1.png",
        "caption": "Fig. 2. Left: Leg Trajectories for Forward Alternating Tripod Gait, Right: RDK Hexapod Robot - www.sandboxinnovations.com",
        "texts": [
            " We constrain \u03b3(\u03be) to be a followable path for the given robot, for example a path for a Dubin\u2019s car is limited to some minimum turning radius. The controller type can change throughout the trajectory so we let the controller type being used at \u03be be represented as \u03ba(\u03be). The problem is to find a path and controllers to be used along that path (a trajectory) that minimize some cost function. We state the generic continuous problem here: minimize J(\u03b3(\u03be), \u03ba(\u03be)) (1) such that \u03b3(\u03be) \u2208C f ree for \u03be \u2208 [0,1] \u03b3(\u03be) is admissible \u03b3(0) = xstart \u03b3(1) = xgoal III. Experimental Setup We implemented our algorithm on the six-legged RDK, shown in Figure 2, which is similar to the RHex robot [10]. This system is a specific instance of the control structure in Figure 1. At the bottom layer, the six legs are controlled to track trajectories using PD control on the hip motors. At the next layer, different types of gaits described in Section III-B send leg trajectories to the motors. The controllers described in Section III-C utilize the gaits to follow line segments. At the highest level our planner generates the trajectories for the different controllers to follow. The alternating tripod gait is used as the starting point for all gaits. To move forward (or backward), the legs within each tripod rotate forward (or backward) along identical trajectories so that one tripod comes into contact with the ground as the other tripod leaves the ground. The exact trajectories followed by legs in this gait are determined by the four parameters shown in Figure 2 where \u03c6= 0 represents the straight down leg position [10]. Here \u03c6s is the stance angle, \u03c60 is the angular leg offset, tc is the period of the gait, and ts is the stance time. Turning while walking is achieved by changing the relative time legs are in the stance phase between the left and right sets of legs, ts(le f t) = ts(base)+\u2206ts and ts(right) = ts(base)\u2212 \u2206ts [10]. Here a positive \u2206ts causes the legs on the right side of the robot to move faster while in contact with the ground than the legs on the left which causes the robot to turn left while walking forward",
            " Using these parametrized trajectories we found a set of base gaits parameters (\u03c6s, \u03c60, tc, ts(base)) that works well at about 0.39 m/s, the slow gait, one that works well at about 0.62 m/s, the fast gait, and one that works well for turning in place. All of these parameters are shown in Table I and were found through testing on carpet. We fix \u03c60 to be constant across all gaits so the robot reaches the position where one tripod is at \u03c0+\u03c60 and the other is at \u03c60 twice during every cycle as shown in Figure 2. For this reason, we allow the controllers to switch to a different gait or change \u2206ts every half cycle. The body velocities and angular velocities achieved using several sets of gait parameters are shown in Figure 3. Here each point represents data captured from about 25 steps for a particular set of gait parameters. This plot shows that we were not able to achieve a turning radius of less than about 1 meter for turning while walking forward or backward. This unique set of achievable motions led to the creation of the hybrid controller for path following described in the next section"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000399_s10999-008-9077-z-Figure24-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000399_s10999-008-9077-z-Figure24-1.png",
        "caption": "Fig. 24 Shot-peening problem: (a) geometry and (b) discretized model",
        "texts": [
            "0%, respectively). Three different impact velocities were used: v = 50, 75 and 100 m/s. Coulomb\u2019s law, with friction coefficient l = 0.25, was assumed. The total integration time was tk = 10 ls with a maximum time step of Dtmax = 2 9 10-8s. Generalized-a time integration scheme was utilized with the following parameters c \u00bcffiffiffi 2 p 1=2; b \u00bc 1=2; aH \u00bc ffiffiffi 2 p 1; and aB = 0. Fournoded axisymmetric finite elements, with large strain and displacement capabilities, were used to discretize the target (Fig. 24). A fine mesh was used in the impact region where high stress gradients are expected. The rest of the target was discretized using larger elements. Convergence tests were conducted revealing only minor variations. Figure 25 shows the time variation of the velocity and total contact force during the entire impact process for an initial shot velocity v = 75 m/s and strain hardening H = 1000 MPa. The results show that contact with the target lasts for 1.70 ls. However, the shot rebound began at about 1"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002173_robio.2009.5420838-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002173_robio.2009.5420838-Figure2-1.png",
        "caption": "Fig. 2 Schematic diagram of the planar UVMS",
        "texts": [
            " \"3 ' (31) In a UVMS redundancy optimization problem, the optimal function (31) is used for joint limit constraints [10], drag force minimization [1] and restoring moment measure [2]. In this work, (31) is utilized in order to minimize the vehicle restoring moment and gravitational joint torque as it consumes more energy during end effector motion. To verify the performance of the proposed redundancy scheme, an ellipsoid AUV with a two-link manipulator is chosen for numerical simulation as shown in Fig. 2 [13]. Table I shows the parameter models used in the simulation, where the neutrally buoyant system is considered. In the two-dimensional task space, there is only surge, heave and pitch motions for vehicle movement. The attached manipulator gives additional DOFs to the system which leads to three redundant DOFs. Initially, the UVMS is in steady state motion and the end effector is required to move from a starting position, [2.57 0 -0.33] m to a final position, [2.77 0 -0.33] m within 10 seconds"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003353_iros.2012.6385499-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003353_iros.2012.6385499-Figure1-1.png",
        "caption": "Fig. 1. Experimental machine of combined rimless wheel with wobbling mass",
        "texts": [
            " We propose the use of a passive wobbling mass that vibrates up-and-down in the body frame. It is expected that the whole CoM orbit during motion becomes flattened by the wobbling mass which vibrates to an opposite direction to the up-and-down motion of the CRW body. This mechanism can be regarded as vibration of viscera and carrying loads [5][6][7]. We discuss the effect mainly from the viewpoint of CoM trajectory, and investigate the potentiality of speeding-up of PDW through numerical simulations and experiments. Fig. 1 shows the overview of our prototype CRW with a wobbling mass that moves up-and-down passively along the guide rail in the body frame. The two RWs are connected via a rigid rod so that they mutually synchronize or the phase difference is kept zero during motion. LEDs are attached on the wobbling mass and body frame for clearly observing the oscillation. In the following, we describe the mathematical model. Fig. 2 shows the ideal model. For simplicity, the rear RW is called \u201cRW1\u201d, and the fore one is called \u201cRW2\u201d",
            " In nonlinear science, such hysteresis is often found at resonances of coupled oscillators [8]. Fig. 12 shows the simulation results with Kc = 250 [N/m], where (a) and (b) correspond to the initial conditions set by increasing Kc from 1 [N/m] and by decreasing Kc from 1000 [N/m], respectively. These results strongly support that, even when Kc is fixed, the CRW can exhibit both anti-phase and in-phase oscillations with different walking speeds depending upon the initial condition. We conducted PDW experiment using the prototype machine shown in Fig. 1 with three springs whose elastic coefficients are Kc = 216.6, 237.2, and 401.6 [N/m]. Table II lists the physical parameters of the experimental machine. We set the slope angle of the treadmill for four values: 3.0, 3.5, 3.9, and 4.2 [deg]. Five experimental data were measured for each spring and slope angle. We could not obtain the walking speeds for some situations because the wobbling mass motion was highly excited. Fig. 13 plots the experimental results of the walking speed with respect to the slope angle"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003824_jas.2014.7004670-Figure11-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003824_jas.2014.7004670-Figure11-1.png",
        "caption": "Fig. 11. Time histories of actual uncertainty and estimated uncertainty.",
        "texts": [
            " System uncertainty is estimated with a linear-inparameter neural network. The following basis functions are used to estimate the uncertainty: \u03c3 (w1, w2) = [w1, w2, (w1) 2 , (w2) 2 , w1w2, (w1) 3 , (w2) 3 , (w1) 2 w2, w1 (w2) 2]T Identifier states and neural network weights are initialized as zero. Fig. 9 shows the system state histories with respect to time under the control action as shown in Fig. 10 (a). Fig. 10 (b) show the extra (adaptive) control needed to compensate for the estimated uncertainty. Fig. 11 (a) and (b) show the actual (modeled) and estimated uncertainties with respect to time, respectively. To see the importance of adaptive control, shown in Fig. 10 (b), the system dynamics was simulated with the nominal (offline) control architecture also. Fig. 12 shows that system is unstable when just the nominal control, designed using SNAC, is applied to the system with uncertainty. Fig. 13 shows the corresponding nominal control action. Figs. 14\u223c 16 compare the results of system trajectory and control action at different spatial points"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000698_j.ijmecsci.2008.02.003-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000698_j.ijmecsci.2008.02.003-Figure1-1.png",
        "caption": "Fig. 1. Model of a flexible rotor supported by two turbulent journal bearings.",
        "texts": [
            " 3, the radial impact force fn and the tangential rub force ft could be expressed as f n \u00bc \u00f0e d\u00dekc (9) f t \u00bc \u00f0f \u00fe bv\u00def n; if eXd (10) Then we could get the rub-impact forces in the horizontal and vertical directions: Rx \u00bc \u00f0e d\u00dekc e \u00bdX \u00f0f \u00fe bv\u00deY (11) C.-W. Chang-Jian, C.-K. Chen / International Journal of Mechanical Sciences 50 (2008) 1090\u20131113 1095 C.-W. Chang-Jian, C.-K. Chen / International Journal of Mechanical Sciences 50 (2008) 1090\u201311131096 Rx \u00bc \u00f0e d\u00dekc e \u00bd\u00f0f \u00fe bv\u00deX \u00fe Y (12) 2.3. Dynamic equations Fig. 1 shows a flexible rotor supported horizontally by two identical and aligned turbulent journal bearings with non-linear springs. Om is the center of rotor gravity, O1 is the geometric center of the bearing, O2 is the geometric center of the rotor, O3 is the geometric center of the journal. Fig. 2 shows the cross-section of the fluid film journal bearing, where (X,Y) is the fixed coordinate and (e, j) is the rotated coordinate, e being the offset of the journal center and j being the attitude angle of the X-coordinate"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001242_s12239-009-0050-0-Figure11-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001242_s12239-009-0050-0-Figure11-1.png",
        "caption": "Figure 11. Movement of turn center according to the virtual rigid axle.",
        "texts": [
            " Thus, the virtual rigid axle must be set up according to the angle of articulation first. Then, the turn center should be made coincident by considering the maximum steering angle of each axle. A flow chart for setting the virtual rigid axles is shown in Figure 10. Assuming that the rear axles are fixed, the value of the virtual rigid axles is zero and the turn center is located at the O position. However, when the rear axles are steered, the turn center moves from the O position to the O' position with the value of the virtual rigid axle, as shown in Figure 11. The maximum steering angle of the 1st axle in equation (7) was obtained by using the maximum articulated angle when the rear axles are fixed. The value of the virtual rigid axles in equations (8) and (10) were taken from the maximum steering angle of the rear axles. Equations (9) and (11) were obtained from the relation between virtual rigid axle 1 and virtual rigid axle 2. \u03b42= tan 1\u2013 \u2013 P1 tan\u03b41\u00d7 l1 P1\u2013 ----------------------- \u239d \u23a0 \u239b \u239e \u03b43= tan 1\u2013 \u2013 P2 tan\u03b11\u00d7 L2 L1\u2013( )+ L1 P1+ cos \u03b11 ----------------\u2212P2 --------------------------------------------------- \u239d \u23a0 \u239c \u239f \u239c \u239f \u239b \u239e \u03b4 n 1+ = tan 1\u2013 \u2013 P n tan\u03b1 n 1\u2013\u00d7 l n L n 1\u2013\u2013( )+ L n 1\u2013 P n 1\u2013+ cos \u03b1 n 1\u2013 --------------------------\u2212P n ------------------------------------------------------------------ \u239d \u23a0 \u239c \u239f \u239c \u239f \u239b \u239e , \u03b10=\u03b41, L0=P0=0 (8) (9) (10) (11) The maximum steering angle in Table 2 is the mean of the maximum steering angles of the right wheel and the left wheel"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000256_00124278-200711000-00011-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000256_00124278-200711000-00011-Figure1-1.png",
        "caption": "FIGURE 1. Two-LPT (linear position transducer) configuration used for the calculation of vertical and horizontal displacement. (a and b) Voltage output from LPT converted to displacement values through calibration to known distances; c constant distance between LPTs; x horizontal displacement; y vertical displacement.",
        "texts": [
            " Briefly, the system incorporates both displacement and force measurements to allow for the calculation of power output throughout movements. Athletes completed the entire lift while standing on a force plate (AMTI, BP6001200, Watertown, Massachusetts, USA). Two LPTs (PT5A-150, Celesco Transducer Products, Chatsworth, CA) located above anterior and above posterior to the subject were attached to the barbell, resulting in the formation of a triangle. This configuration allows for calculation of vertical and horizontal displacement through trigonometry involving constants and displacement measurements from the LPTs (Figure 1). From laboratory calibrations, the LPTs and force plate voltage outputs were converted into displacement and vertical ground reaction force, respectively. The analog signals were collected for every trial at 1,000 Hz using a BNC-2010 interface box with an analog-to-digital card (NI PCI-6014; National Instruments, Austin, TX). LabVIEW (version 7.1; National Instruments) was used for recording and analyzing the data. Signals from the 2 LPTs and the force plate underwent rectangular smoothing with a moving average half width of 12"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001247_0954406jmes321-Figure3-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001247_0954406jmes321-Figure3-1.png",
        "caption": "Fig. 3 Applied coordinate systems for generation of the face gear",
        "texts": [
            "comDownloaded from respectively x1 = rbs(\u00b1\u03b8ks cos(\u03b8ks + \u03b8os \u00b1 \u03c8) \u00b1 sin(\u03b8ks + \u03b8os \u00b1 \u03c8)) y1 = \u2212rbs(\u03b8ks sin(\u03b8ks + \u03b8os \u00b1 \u03c8) + cos(\u03b8ks + \u03b8os \u00b1 \u03c8)) z1 = p\u03c8 (7) The surface normal vector N 1 of the shaper is determined in coordinate system S1 by the matrix given in equation (8) N 1 = \u2202R1 \u2202\u03b8ks \u00d7 \u2202R1 \u2202\u03c8 = \u03b8ksrbs \u23a1 \u23a2\u23a3 \u2212p cos(\u03b8ks + \u03b8os \u00b1 \u03c8) \u2213p sin(\u03b8ks + \u03b8os \u00b1 \u03c8) rbs \u23a4 \u23a5\u23a6 (8) The helical gear tooth surface defined in equations (7) and (8) can be used as a shaping cutter to generate the tooth surface of a standard face gear. The process of generating a face gear by a shaper is illustrated in Fig. 3. The coordinate systems S1(x1, y1, z1) and S2(x2, y2, z2) are rigidly attached to the shaper and the face gear, respectively. The coordinate systems Sm(xm, ym, zm) and Sf (xf , yf , zf ) are rigidly attached to the frame of the cutting machine. The shaper and the face gear rotate about axes Z1 and Z2 with angles \u03c61 and \u03c62, respectively. The shaft angle \u03b3m is the angle between the shaper and the face-gear rotational axes, while parameter E is the shortest distance between the axes of rotation of the shaper and the face gear"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003924_1.4030612-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003924_1.4030612-Figure1-1.png",
        "caption": "Fig. 1 12DOF element and coordinates",
        "texts": [
            " The supporting systems, including fluid bearings and external torsional forces from motor, are treated as nonlinear. The discussion of coupled torsional and lateral analysis is a further development of prior lateral transient analysis [14]. To calculate the coupled lateral and torsional analysis, an element with 12 degrees-of-freedom (DOF) is applied to model the linear shaft of the complex rotor system based on the FE method with a fixed (nonrotating) frame as reference. The system of fixed global coordinates and a 12DOF element used for shaft modeling are shown in Fig. 1. The displacement vector of each node is (x, y, z, hx, hy, and hz). The shaft is assumed to be axially symmetric and modeled with Timoshenko beam elements for lateral directions (x, y, hx, hy) and a linear element for axial and torsional directions (z, hz). The built on components, such blades, disks, and impellers, are modeled as lumped masses plus the properties of polar and transverse mass moments of inertia. An additional node with 2DOF (xsd, ysd) is added for each SFD to indicate the location of bearing center, which whirls inside the SFD, as shown in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001704_acc.2009.5160384-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001704_acc.2009.5160384-Figure2-1.png",
        "caption": "Fig. 2. Voronoi interpretation",
        "texts": [
            " The above theorem shows that it is possible to use a broadcast control command to make two agents meet at the same location simultaneously for almost all initial conditions. However, the solution is also unique and hence the location of the meeting point cannot be chosen arbitrarily. One can also interpret this result by noting that the final meeting point is on the Voronoi edge (equidistant line) between the two initial positions of the agents. It can be shown that only one unique point on the Voronoi edge satisfies the requirement that the orientation change angle is the same for both the agents (see Fig. 2). The point p moves on the equidistant line from \u2212\u221e to +\u221e and the corresponding orientation angle change \u03b8 is plotted for the two agents. The intersection of the two curves is the unique control command point. It can be seen that when the number of agents is more than two, each pair gives rise to a different unique meeting point. Thus, there does not exist a common control command to be broadcast so that all the agents meet at a point. In the absence of such a command, the best that can be done is to determine a (\u03b8, d) which brings the agents in closest proximity with each other"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002930_978-3-642-31988-4_26-Figure5-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002930_978-3-642-31988-4_26-Figure5-1.png",
        "caption": "Fig. 5 Geometric parameter of the six-cable driven parallel manipulator",
        "texts": [
            " The idealization would be possible if the ends of the cable were subjected to tensions that are predominantly larger than the effect of the cable mass or the accuracy requirement is not high. In Fig. 3, line equation with elastic deformation of a cable can be easily derived as follows: l = ( h2 + L2 )1/2 (13) T = ( V 2 + H2 )1/2 (14) L = l H T (15) h = l V T (16) \u0394l = T l E A0 (17) where \u0394l is elastic deformation of the cable. In Fig. 4, for studying the largest radio telescope FAST, a similarity model of FAST is set up in Beijing. The related geometric parameters of the six-cable driven parallel manipulator in the similarity model are given in Table 1 [9]. In Fig. 5, two coordinates are set up for the six-cable driven parallel manipulator: an inertial frame : O \u2212 XY Z is located at the center of the reflectors bottom. Another moving frame \u2032 : O \u2032 \u2212 X \u2032Y \u2032Z \u2032 is located at the center of the moving platform. Bi (i = 1, 2, ..., 6) are the connected points of the cables and cable towers, and A j ( j = 1, 2, 3) are the connected points of the cables and moving platform. For analyses, the symbols used in this section are defined as: O \u2032 is the O \u2032 expressed in the inertial frame; B i the vector Bi expressed in the inertial frame; A j the vector A j expressed in the inertial frame; A \u2032 j the vector A j expressed in the moving frame; rb the radius of the cable towers distributed circle; ra the radius of the moving platform; h the height of the cable tower. According to Fig. 5, the vector of the cables can be expressed as: B i = [rb cos((i \u2212 1)\u03c0/3), rb sin((i \u2212 1)\u03c0/3), h]T , i = 1, 2, . . . 6 (18) A \u2032 j = [ra cos((4 j \u2212 3)\u03c0/6), ra sin((4 j \u2212 3)\u03c0/6), 0]T , j = 1, 2, 3 (19) A j = R \u00b7 A \u2032 j \u2212 O \u2032 (20) where R is the coordinate-axis rotation matrix. Assuming Li = A j \u2212 B i , ui = Li/ \u2016Li\u2016 , r i = A j \u2212 O \u2032 , static equilibrium equation of the six-cable driven parallel manipulator can be written as: F = JT\u03c3 (21) where \u03c3 is the cable tension vector; JT the tension transmission matrix of the cable driven parallel manipulator; F \u2208 Rn the wrench of the moving platform"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000303_j.jsv.2006.09.031-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000303_j.jsv.2006.09.031-Figure1-1.png",
        "caption": "Fig. 1. Meshing of pinion and gears with three contact lines and the friction forces on the contact zone.",
        "texts": [
            " The variation of coefficient of friction has been attributed to a number of parameters, such as tooth geometry, surface hardness, axial velocity, contact pressure and misalignment [16]; contact ratio, speed ratio, transmitted load and lubricating condition [17]; and viscosity and sliding velocity [18]. Though the sliding friction is very prominent, Michlin and Myunster [19] have found that the rolling friction cannot be neglected as it causes frequent spalling near the pitch point. This paper is devoted to determine variation in the number of contact lines in the meshing zone of a helical gear system. An example of the meshing action of a helical gear system is illustrated in Fig. 1. The pressure plane or contact zone of the gear meshing is limited by line of the intersection of the addendum circles of the gear and pinion with the tangent drawn through the two base circles. The contact lines are formed by the engaged teeth of the meshing gear and pinion. Since the width of the gear and pinion is not an integral multiple of the axial pitch, there is a continuous variation of the contact lines with the phase of engagement. The terminologies used in this paper is as given below: l1 \u00bc b1 tan ab l, l2 \u00bc 3pt l b1 tan ab, l3 \u00bc ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi r2a \u00f0r cos ft\u00de 2 q r sin ft, l4 \u00bc 2pt b1 tan ab, l5 \u00bc b1 2 tan ab l3, \u00f01\u00de C"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003713_978-1-4614-2419-2-Figure19.1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003713_978-1-4614-2419-2-Figure19.1-1.png",
        "caption": "Fig. 19.1 Electro-mechanic scheme of pitch system",
        "texts": [
            " The pitch system is also used as WTGs\u2019 main braking system where individual blade pitch offers redundancy. The mechanism of the pitch system can be either hydraulic or electro-mechanic and the focus of this work is on electromechanic systems. The blade bearing is a part of such a pitch system and its main role is to connect rotor blades to a hub that enables the blade to rotate around its longitudinal axis. An electro-mechanical pitch system also includes a drive control unit, motor and gearbox, as illustrated in Fig. 19.1. It operates in such a way that the WTG pitch controller sends a control signal to the drive control unit which then calculates a drive signal, i.e. torque demand for the pitch motor. The gearbox is connected to the blade bearing through a pinion-tooting mechanical system. Many different designs of bearing, hub and blades are present in the industry and they strongly interfere with the friction and blade response. Thus, a generalized model of pitch system, which includes all different configurations and effect, would be very complicated to make"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002263_j.proeng.2011.11.115-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002263_j.proeng.2011.11.115-Figure1-1.png",
        "caption": "Figure 1. Two types of samples manufactured by DMLS in this work: (a) H-geometry samples (b) Prototyped guide vanes",
        "texts": [
            " These residual stresses are specially important in some cases since they are added to the external stress fields leading to a real stress in the component higher (in the case of tensile residual stresses) than the nominal applied stress and therefore its fatigue life could be reduced resulting in a premature failure of the component. In this work, surface integrity of CoCr, Maraging Steel and Inconel 718 DMLS parts subjected to different thermal and finishing (surface polishing and shot peening) treatments has been characterized with the objective of determining the optimum stress relieving treatment. The technology used to manufacture the samples has been DMLS (Direct Metal Laser Sintering). Two different sample geometries have been studied: CoCr, Inconel 718 and Maraging Steel H-geometry samples (Figure 1 (a)) and Maraging Steel prototyped guide vanes (Figure 1 (b)). For the manufacturing of the samples a commercial DMLS system (EOS M270) with a 200 W Yttrium laser has been employed. In H-geometry samples it has been studied the effect on the final surface integrity of shot peening and different thermal treatments: aging in Inconel 718 and 2 hours of thermal treatment at 650, 850 and 1000 \u00baC under different atmospheres (air, vacuum, N2) in CoCr samples. Ceramic balls of 0.5 mm diameter have been propelled using air at 4 bar of pressure for the shot peening"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001660_1.3065533-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001660_1.3065533-Figure1-1.png",
        "caption": "FIG. 1. Schematic sketch for a physical model and coordinates, and b computed results on the transverse cross section at z=0.",
        "texts": [
            " Even though the transport phenomena have been extensively studied, no systematic investigation has been undertaken to understand the origin of the formation wavy fusion zone boundary. In this work, an unsteady two-dimensional model3 is used to simulate the quasisteady three-dimensional heat transfer and fluid flow during melting and welding. The distributed incident flux moving at a constant speed along the centerline and in the negative z coordinate can be considered as a time-dependent distributed incident flux with a half beam width X on the transverse cross section at z=0, as sketched in Fig. 1 a . Fluid flow in the molten pool is driven by thermocapillary force on the flat free surface at this cross section, as illustrated in Fig. 1 b . To simplify the analysis without loss of generality, the following major assumptions are made. 1 An unsteady two-dimensional heat transfer model with time-dependent incident flux can be applied to simulate the quasisteady three-dimensional heat transfer. This can be verified from scale analysis. The incident power in the three-dimensional model for a high scanning speed is of the same order of magnitude as the energy required to raising temperature of the incoming solid from the ambient temperature to the melting point, Q cp Tm \u2212 T Uwh , 1 where Q and U are, respectively, beam power and scanning speed of the incident flux, w and h are width and depth at the workpiece surface and centerline of the molten pool, respectively, and , cp, Tm, and T are density, specific heat at constant pressure, melting, and ambient temperatures, respectively",
            " Thermal conductivities and specific heats are averaged between solid and liquid k cp = f + 1 \u2212 f ks cps . 7 Thermocapillary force, which is the driving force for the fluid flow in the pool, is balanced by viscous stress on the flat free surface u y = Ma Pe T x . 8 Marangoni number in Eq. 8 is defined as Ma \u2212 d /dT Tm\u2212T / , where and are, respectively, thermal diffusivity and dynamic viscosity of the liquid. The incident flux of a Gaussian distribution at location z=0 on the top surface is given by T y = 3Q exp \u2212 3r 2 for x X t , 0 t 2, 9 where dimensionless beam power Q Q / k Tm\u2212T . Referring to Fig. 1 a , Pythagorean theorem see Fig. 1 a gives r2=x2+ \u2212Ut 2 at any time. Dimensionless radial coordinate therefore is related to x coordinate by r t = x 2 + 1 \u2212 t 2. 10 Substituting r t =1 and x =X t , Eq. 10 gives the timedependent half width of the incident flux appearing in Eq. 9 , X t = 1 \u2212 1 \u2212 t 2. 11 Equations 9 \u2013 11 indicate that half width of the incident flux is a function of scanning speed, energy distribution parameter, and time. The half width of the incident flux increases from zero to the beam radius as time increases from zero to t =1"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000265_978-3-540-71364-7_23-Figure22.4-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000265_978-3-540-71364-7_23-Figure22.4-1.png",
        "caption": "Fig. 22.4. (left) Experimental setup, (right) Geometric relation for the task \u201cpathfollowing\u201d",
        "texts": [
            " In the robotic procedure adopted by systems such as Robodoc [28], Acrobat [29] and CASPAR [30], a mill is used for the cutting procedure, and the blade is required to cut along the planned path on the bone. Meanwhile, the cutting edge of the tool should be kept perpendicular to the cutting plane in order to provide more efficient force. We model the femur cutting task as a task to guide the tip of a long straight tool following a 2D b-spline curve C1 in plane \u03a0 while keeping the tool shaft perpendicular to the plane. The geometric relation is shown in Figure 22.4. We assume that the path C1 and the cutting plane are known in the robot coordinate frame by using an appropriate registration method. During the procedure, the tip of the tool (task frame {t}) is allowed to move along the planned path C1. At the same time, a point, xp,s on the tool shaft (task frame {s}) is only allowed to move along the second path C2, which is a translation of C1 above the target plane. xcl p,t is the closest point to the tip of the tool on C1 and xcl p,s is the projection of xcl p,t on C2"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003531_j.piutam.2011.04.015-Figure5-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003531_j.piutam.2011.04.015-Figure5-1.png",
        "caption": "Fig. 5: The multibody model",
        "texts": [
            " The ASLIP is able to regulate its torso far more accurately (within 2e-16 radians, of the desired orientation) than the 1\u25e6 \u22122\u25e6 of torso sway that is typical of human gait [14] likely because its hip torques are not bandwidth limited, and its back segment is perfectly rigid. The sharp change in hip torque as the leg transitions from single to double stance (and vice versa) causes a subtle cusp in the horizontal ground reaction force profile (Fig. 4 C) that is not present in human ground reaction force profiles. An 11-dof anthropomorphic sagittal plane gait model (Fig. 5) was developed along with a control system for the legs to regulate the orientation of the torso of the multibody model and cause it to track the XY position of the target SLIP model. The multibody model is controlled using 6 joint torques applied at both hips, knees and ankles. The model interacts with the ground using a two-segment foot contact model \u2014 an extension of Millard et al. [24] \u2014 consisting of two spherical volumetric contact elements [25] to represent the heel and metatarsal pads. The midfoot is not rigid, but is allowed to flex slightly at a revolute joint that has a linear spring-damper in parallel with it",
            " Each of the feedback error terms (\u03bdx,\u03bdy and \u03bd\u03b8) take the form of a state feedback PD controller: \u03bdx = \u2212Kx(xM \u2212 xS ) \u2212 Dx(x\u0307M \u2212 x\u0307S ) (35) \u03bdy = \u2212Ky(yM \u2212 yS ) \u2212 Dy(y\u0307M \u2212 y\u0307S ) (36) \u03bd\u03b8 = \u2212K\u03b8(\u03b8M \u2212 \u03b80) \u2212 D\u03b8(\u03b8\u0307M). (37) Input-output feedback linearization [17, 18] is used to compute the hip, knee, and ankle torques required to accelerate the torso of the multibody model such that Eqns. 32-34 are satisfied. The input-output feedback linearization control expressions are not formulated using the multibody model (Fig. 5) \u2014 due to the difficulties the full foot model introduces \u2014 but with an approximate single stance model (Fig. 6A) that includes a rigid foot. To form the control law, we first begin with the equations of motion of the stance control model (Fig. 6A) in functional form (using square brackets to denote matrices). \u0308\u03b3S S = [MS S ]\u22121 4\u00d74 ( \u2212 CS S + [PS S ]4\u00d73 { \u03c4S S } 3\u00d71 + [QS S ]4\u00d73 { FS W \u03c4S W } 3\u00d71 ) (38) In Eqn. 38 \u03b3S S is the vector of joint angles (and respective derivatives) of the single stance (SS) control model (Fig",
            " The hip, knee and ankle states can be mapped directly from the multibody model to the stance control model. \u03b31 S S = \u2212(\u03b8M \u2212 \u03c0 2 + \u03b1M) (41) \u03b32 S S = \u2212\u03b2M (42) \u03b33 S S = \u2212\u03b3M (43) \u03b3\u03071 S S = \u2212(\u03b8\u0307M + \u03b1\u0307M) (44) \u03b3\u03072 S S = \u2212\u03b2\u0307M (45) \u03b3\u03073 S S = \u2212\u03b3\u0307M (46) The geometry of the foot of the control model (Fig. 6A) \u2014 the length of the link between the COM of the foot and the revolute joint attached to the ground \u2014 is adjusted so that the revolute joint attaches to the ground at a location that coincides with the COP of the foot of the multibody model (Fig. 5). The angular velocity of the stance model foot, and the translational velocity of the COP of the stance model (\u03b3\u03074 S S in Fig. 6A and x\u0307COP in Fig. 6E) are computed such that the translational velocity of the ankle joints of the stance control and multibody gait model match. Once the swing foot comes into contact with the ground, the controller changes its internal state from single stance to double stance (Fig. 7), and employs a completely different control model (Fig. 6C), for which a new control law must be derived",
            " 57 for the Lagrange multipliers that yield the desired joint reaction forces and substituting the result into Eqn. 56 we have [MDS ]9\u00d79 { \u0308\u03b3DS } + CDS + \u23a7\u23aa\u23aa\u23aa\u23aa\u23a8\u23aa\u23aa\u23aa\u23aa\u23a9 05\u00d71 DDS ,4\u00d71[BDS ]\u22121 { F1 DS F5 DS } 4\u00d71 \u23ab\u23aa\u23aa\u23aa\u23aa\u23ac\u23aa\u23aa\u23aa\u23aa\u23ad = [PDS ]9\u00d76 { \u03c4DS } 6\u00d71 (58) Since the two hip torques (\u03c41 DS and \u03c45 DS ) and forces ( f 1 DS and f 5 DS ) are known from the solution to Eqns. 50-52 and Eqns. 53-55, Eqn. 58 has embedded in it a set of four equations (the constraint equations) that are linear in four unknowns (\u03c42 DS , \u03c43 DS , \u03c46 DS , and \u03c47 DS ). After the state of the multibody model (Fig. 5) is mapped to the equivalent state of the double stance control model (Fig. 6) \u2014 using the same procedure detailed in Sec. 3.1 \u2014 Eqn. 57 can be solved for the remaining knee and ankle torques required to satisfy Eqns. 47-49. The swing phase has been a topic of robotics research for many years and has resulted in a number of standard approaches: active trajectory tracking [4, 30], passive swing [11, 31], and a combination of passive and active swing techniques [32]. Although a lot of research has been done on the topic of swing, little of it is directly applicable to formulating a control law that will yield a human-like swing phase"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002220_978-1-4020-9340-1_8-Figure8.9-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002220_978-1-4020-9340-1_8-Figure8.9-1.png",
        "caption": "Fig. 8.9. Temperature field of a nascent weld bead on sections along its plane of symmetry (a) and on the interface between the substrate (AISI1045 steel) and the Stellite 21 coating (b). The heavily printed curves enclose the molten region, the lateral width of which determines the bead width. The dashed circle corresponds to the focal spot of the top-hat laser beam. Parameters: laser beam power P : 900W, beam diameter: 2.6mm, feed rate: 10mms\u22121, mass flow: 0.1 g s\u22121, powder jet diameter at the focus of the coaxial nozzle: 5.0mm, absorptivity: 50%.",
        "texts": [
            " Starting from the temperature field of a laser beam moving over a plane substrate, we first calculate the bead shape on the basis of the modified ellipsoid model described above and then the corresponding temperature field. The latter requires corrections of the bead shape for which the temperature field has to be recalculated and so on until convergence is achieved. Figures 8.9 and 8.10 show the temperature distribution and bead geometry as calculated in this way for two different values of the laser beam power and otherwise equal process parameters. The optimised power of 900W (Fig. 8.9 and Fig. 8.10a) leads to a tight welding joint between substrate and bead with minimum dilution as shown in Fig. 8.10c, whereas an excess of power causes enhanced substrate melting and dilution (see Fig. 8.10b,d). Obviously, the shape of the interface between substrate and bead observed in Fig. 8.10d differs strongly from the calculated one shown in Fig. 8.10b. It cannot be explained by a purely conductive heat flow model and is probably caused by the convection in the melt pool. The effect of fluid flow in the melt pool on the temperature field was neglected in the preceding section for simplicity"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003751_icuas.2013.6564678-Figure3-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003751_icuas.2013.6564678-Figure3-1.png",
        "caption": "Fig. 3. Pure rolling motion",
        "texts": [
            " Fx = q\u0304SCx0 (13) Fy = q\u0304SCy (14) N = q\u0304SCn (15) where Cx0, Cy y Cn are the aerodynamic coefficients involved for the lateral dynamics. These coefficients are obtained by considering small angles and with low airplane speed. The following equations describe the dynamics for the roll motion: \u03c6\u0307 = p (16) p\u0307 = L\u0304 Ixx (17) V\u0307y = Fy m + pVx (18) V\u0307x = Fx m \u2212 pVy (19) where p denotes the roll rate, L\u0304 is the rolling moment, Ixx represents the inertia for the x-axis and \u03c6 describes the roll angle. The aerodynamic effects on the airplane are obtained as they have been obtained in the yaw motion. In the Figure 3, it is observed that \u03b4a represents the deviation of the ailerons. In the case of the roll moment, this corresponds the expression L\u0304 = q\u0304SbCL, where b is the wing span of the airplane and CL represents the aerodynamic coefficient of the roll moment [3]. In this section, we describe the nonlinear controllers that have been designed in order to control the fixed-wing UAV. In order to design the altitude control law, we consider the equations that define the longitudinal dynamics, except the equation (1) which defines the linear longitudinal velocity, because it is considered to be constant"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002218_j.ymssp.2010.12.001-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002218_j.ymssp.2010.12.001-Figure1-1.png",
        "caption": "Fig. 1. Schematic of dual-rotor system: (1) left bearing housing, (2) outer rotor, (3) inner rotor and (4) right bearing housing.",
        "texts": [
            " In the first part, this paper presents the whole-beat correlation method for the whole-machine balancing of the dual-rotor system based on correlation theory. In the second part, an optimized whole-beat correlation method is proposed based on error analysis. In the last part, a balancing experiment is conducted on the horizontal decanter centrifuge, validating the precision, efficiency, and applicability of the recommended method. & 2010 Elsevier Ltd. All rights reserved. There is a special dual-rotor system in industry. As shown in Fig. 1, this co-axial dual-rotor system comprises an outer rotor together with an inner rotor conforming to the profile of the outer rotor by bearings on the two sides, rotating at a slightly differential speed compared with the outer rotor. The outer rotor is equipped on the bearing pedestal [1,2]. When the inner and outer rotors both are in presence of unbalance mass, there will be a combined vibration described as Eq. (1), which can be tested from the outer bearing pedestal: x\u00f0t\u00de \u00bc x1\u00f0t\u00de\u00fex2\u00f0t\u00de \u00bc A1 sin\u00f0o1t\u00fej1\u00de\u00feA2 sin\u00f0o2t\u00fej2\u00de \u00f01\u00de where x1(t) and x2(t) are the vibration signals and Ai(i=1,2) the amplitude, oi(i=1,2) the rotating angular frequency, and ji(i=1,2) the vibration phase relative to the key-phase"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000857_j.mechatronics.2008.01.003-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000857_j.mechatronics.2008.01.003-Figure1-1.png",
        "caption": "Fig. 1. The electro-mechanical components of the supply kit.",
        "texts": [
            " During the course of the semester, numerous tools are provided for the students. These tools include a supply kit containing various electro-mechanical and pneumatic devices, including a stand-alone controller. The supplies allow the students to construct a reasonably complex and powerful machine. In addition, several machine tools and various hand tools are provided for the students to manufacture their designs. The electronic and electro-mechanical components of the supply kit issued to the students are shown in Fig. 1. The central element of the kit is a BASIC Stamp powered controller box, designed and built at Georgia Tech. The box is capable of driving two DC motors and one stepper motor at various speeds in two directions, in addition to three solenoids. The box has sensing capabilities via two micro-switches, an infrared distance sensor, an encoder, and a flex sensor, all of which are provided as part of the supply kit. To program the controller boxes, the PBASIC programming language is used. Where possible, the kit has been made \u2018\u2018plug-and-play\u201d with many of the difficult operations hard-wired into the board and code examples made available"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003903_tdei.2015.005053-Figure14-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003903_tdei.2015.005053-Figure14-1.png",
        "caption": "Figure 14. 3D model of XZP2-300 type porcelain insulator and its meshing result.",
        "texts": [
            "5cm to the height of exposed part of zinc sleeve based on the existing height. For organic materials, if its surface electric field intensity is higher than 5 kV/cm, surface discharge phenomena occur. It poses adverse impact on the mechanical and electrical characteristics of organic material and accelerates its aging. So, the calculation of electric field intensity is significant to the material selectionand structure optimize design of organic sleeve. The 3D model of XZP2-300 type porcelain insulator is shown in Figure 14a. In order to monitor the electric field intensity, a path from iron cap to pin on the insulator's surface is selected for observation, as illustrated in Figure 14a. The voltage applied on pin is 15 kV. The relative permittivity and conductivity of each element are listed in Table 3.The meshing resultis shown in Figure 14b. The surface electric field intensity along observation path under without organic material sleeve and with high temperature vulcanized silicone rubber sleeve (referred to as silicone rubber sleeve) conditions are shown in Figure 15. Figure 15 illustrates that, the installation of silicone rubber sleeve can decrease the surface electric field intensity near iron cap, but the maximum field intensity can reach 7.8 kV/cm at the pin side. When water drop existed on this area, surface discharge will occur"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003713_978-1-4614-2419-2-Figure31.2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003713_978-1-4614-2419-2-Figure31.2-1.png",
        "caption": "Fig. 31.2 Rendering of torsional vibration test bed with 30 hp drive motor and 70 hp load motor connected by 25 mm diameter shaft",
        "texts": [
            " The test stand was designed to: (1) have a scale more representative of an industrial installation; (2) capability to grow a fatigue crack in situ without disassembly and reassembly of the shaft line; (3) contain a comprehensive suite of health monitoring instrumentation including accelerometers, shaft proximity sensors and torsion vibration sensors; (4) continuous monitoring of health features. The following section will describe the test stand and the outcome of the baseline testing which yielded insightful data about the operational and environmental effects on the acquired machinery health features. 314 M.S. Lebold et al. Both torsional and lateral loads can be applied to a shaft during testing. The test rig is depicted in Fig. 31.2 and consists of a Figure 31.2 renders the torsional vibration test bed with 30 hp drive motor and 70 hp load motor connected by 25 mm diameter shaft. When a shaft is installed in the test stand, it is supported by bearings in pillow blocks at locations symmetric about the shaft midpoint. Both temperature and horizontal acceleration are collected from accelerometers installed on the top of both pillow blocks. To laterally load a test shaft, the hydraulic ram is actuated and provides a force on the test rig\u2019s load plate assembly"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003998_s0026261713040139-Figure3-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003998_s0026261713040139-Figure3-1.png",
        "caption": "Fig. 3. Current density generated by Shewanella oneidensis MR 1 (1), FRS1 (2), and FRB1 (3) on days 13 and 20 of cultivation in MFC1.",
        "texts": [
            " The maximal difference in voltage between the original strain and the mutants at days 10\u201325 of the experi ment was 200 mV. The maximal voltage for the mutants and the original strain was 620 and 600 mV, respectively. Importantly, the original strain main tained the maximal voltage for 4 days, with a subse quent drastic decrease. Unlike the original strain, the mutant FRS1 maintained 620 to 570 mV until day 25 of the experiment. After 10 days of incubation, the voltage generated by strain MR 1 decreased to 370 mV, while in the case of the mutants it was 570 mV. The diagrams presented on Fig. 3 show the current density calculated for days 13 and 20 of the experi ment. It can be seen that the current density generated by the mutants on day 13 was ~1.7 times higher com pared to the original strain. On day 20, the current density generated by FRS1 retained at the same high level, while in the case of FRB1 it decreased, remain ing, however, 1.2 times higher than the current density MICROBIOLOGY Vol. 82 No. 4 2013 INTENSIFICATION OF BIOELECTRICITY GENERATION IN MICROBIAL FUEL CELLS 413 generated by the original strain"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003479_bf02326326-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003479_bf02326326-Figure1-1.png",
        "caption": "Fig. 1--Specimen configuration",
        "texts": [
            " A fine wire was then placed th rough the hole to pro tec t it, while a jewelers file was used to cut a slot f rom the edge to the hole, thus forming a notch with a uniform root radius. The specimen was loaded in tension, using loading pins which assured very good ax ia lqoad appl ica t ion . The specimens were then subjec ted to a convent ional stress-freezing cycle to increase the i r sensit i v i t y and to enable more f reedom in handl ing. On complet ion of the stress-freezing cycle, the dimensions of the specimen were measured, including the width of bo th the nar row and wide por t ions of the shank, the hole location, the root rad ius and the dep th of the notches (Fig. 1). This app roach insured t h a t the dimensions would be effectiv- Experimental Mechanics ] 513 from the grinding was eliminated and any dimpling effect, al though small, would not obstruct the light path through the specimen. Using the above procedure and the appropriate magnification (from 70 X to about 300 \u2022 the number of fringes produced at the boundary was sufficient to obtain an accurate measure of the stress at the bo t tom of the notch. The focusing was done on an imaginary plane approximately half way between the outside faces of the specimen",
            " The mos t well known of these is Neube r ' s approximat ion 4 in which the s t ress-concent ra t ion factors for shallow el l ipt ical notches and deep hyperbol ic notches are combined to ob ta in an approx ima te fac tor for notches of a r b i t r a r y depth . Also, in a recent s tudy of the influence of notch dep th in edgeno tched plates, Dixon, 9 using a l imi t ing process of Neube r ' s solution, ob ta ined an a l t e rna te expression for the s t ress-concent ra t ion factor for small rad i i in the form 1 + 2(a/p) '/2 K = (1 -- a/b)l/~[1 -}- (Ir2/4 -- 1)a/b] '/2 (1) where a, b, and p are defined in Fig. 1. Fo r the range of var iables included in this s tudy , the stressconcent ra t ion factor ob ta ined f rom Neube r ' s m e t h o d or from the above expression differs by less t han 1 percent . A different approach to the problem of finding an approx imate s t ress-concent ra t ion factor, based on l inear f rac ture-mechanics concepts, was also demons t r a t ed by Dixon in the same paper . I rwin 18 had proposed t ha t the s t ress- in tens i ty factor, KI, can be obta ined, for the opening mode, from the l imit ing value of a known s t ress-concentrat ion factor by the re la t ion K I = lima,~a~(Trp)'/2/2 (2) p~O where the ~,/2 t e rm has been in t roduced in eq (2) to t ake into account the recent redefini t ion of Kt, and where area= is the m a x i m um stress a t the root of the notch"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003520_j.euromechsol.2011.11.003-Figure10-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003520_j.euromechsol.2011.11.003-Figure10-1.png",
        "caption": "Fig. 10. Maximally regular T2R1-type parallel manipulator with bifurcated planarspatial motion of the moving platform: constraint singularity (a), branch with planar motion (b) and branch with spatial motion (c); limb topology PkRtPtkRtRPassPatcstPtkPtRtRSP.",
        "texts": [],
        "surrounding_texts": [
            "The term of constraint singularity (CS) has been recently coined (Zlatanov et al., 2002) to characterize the configuration of lower mobility parallel manipulators in which both the connectivity of the moving platform and the mobility of the parallel mechanism increase their instantaneous values. From a constraint singularity, the mechanism can get out with or without branching (Gogu, 2008b). When branching occurs in a constraint singularity, the mechanism can reach different configurations, called branches, and have different independent motions of the moving platform. In this case, the constraint singularity is also called branching or bifurcation singularity. A branch refers to the free-of-singularity configurations of the mechanism inwhich each structural parameter keeps its value. For this reason, this value is called global of full-cycle value for a branch. Two types of branching in constraint singularity (BCS) have been defined in Gogu (2008b). Branching of type BCS1 occurs when a parallel mechanism F ) G1 . Gj . Gk gets out from a constraint singularity in different branches characterized by the same mobility and the same connectivity of the moving platform but with different bases of the vector space of relative velocities between the moving and fixed platforms. Branching of type BCS2 occurs when a parallel mechanism F ) G1 . Gj . Gk gets out from a constraint singularity in different branches characterized by distinct values of mobility and connectivity of the moving platform. The parallel manipulators presented in this paper have constraint singularities with branching of type BCS1. Branching of type BCS2 occurs in kinematotropic mechanisms (Gogu, 2009c). The finite displacements and the velocities in the actuated joints are denoted by qi and _qi, the linear velocities of the characteristic point H of the moving platform, by v1 \u00bc _x; v2 \u00bc _y and the angular velocity of the moving platform by u \u00bc ua \u00bc _a or u \u00bc ud \u00bc _d: In both branches, the moving platform undergoes two planar translations and one rotation but the rotation axis is different in the two branches. In the first case, the rotation velocity u \u00bc ua \u00bc _a is parallel to the x-axis and the moving platform undergoes a spatial motion. In the second case, the rotation velocity u \u00bc ud \u00bc _d is parallel to the z-axis and the moving platform undergoes a planar motion. With these notations, the linear mapping between the actuated joint velocity space and the end-effector velocity space for a branch the T2R1-type PMs is defined by: 2 4 v1 v2 u 3 5 \u00bc \u00bdJ 2 4 _q1 _q2 _q3 3 5 (8) where J is the Jacobian matrix. In the constraint singularity, the T2R1-type parallel manipulators with bifurcated motions has instantaneously iM \u00bc iSF \u00bc 4 and (iRF)\u00bc(v1, v2, ua, ud). The bifurcation in this constraint singularity can be used to change motion type of the moving platform. In the two distinct branches, the parallel mechanism is characterized by M\u00bc SF \u00bc 3 and (RF)\u00bc(v1, v2,ua) or (RF)\u00bc(v1, v2,ud). In the constraint singularity, both connectivity of moving platform and mobility of parallel manipulator increase their instantaneous values. More details about the behaviour of parallel manipulators in a constraint singularity and the associated structural parameters can be found in Gogu (2009c)."
        ]
    },
    {
        "image_filename": "designv11_12_0001643_09544062jmes1340-Figure3-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001643_09544062jmes1340-Figure3-1.png",
        "caption": "Fig. 3 Geometric conditions for the line contact arrangement of a cylindrical gear pair",
        "texts": [
            " The line of action is extended to a plane and the gear teeth can be engaged in line contact. The condition of line contact can be obtained from equations (7) and (8) by = \u03b2Cb2 + \u03b2Cb1 (30) d = rCb1 + rCb2 (31) The derived equations can be also explored via the geometrical relation. For the case of line contact, the Proc. IMechE Vol. 223 Part C: J. Mechanical Engineering Science JMES1340 \u00a9 IMechE 2009 at UNIV OF MICHIGAN on June 20, 2015pic.sagepub.comDownloaded from relation between the base cylinder of gear 1 and gear 2 (see Fig. 1(a)) must satisfy the conditions below (as shown in Fig. 3): (a) the shaft angle is the sum of the base helix angle of gear 1 and of gear 2; (b) the offset d is equal to the sum of the base radius of gear 1 and of gear 2. The line contact condition is also valid when the angle \u03b7 is equal to zero. By substituting \u03b7 = 0 into equation (17), the same result as for equation (30) can be obtained. Edge contact occurs when the line of action shifts from its theoretical position outside the face-width.This can occur due to assembly errors or manufacturing errors"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001931_iros.2010.5649323-Figure19-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001931_iros.2010.5649323-Figure19-1.png",
        "caption": "Fig. 19. The Flow Field of Fig. 16 with the new bj(z)",
        "texts": [
            " That is, the ellipse flow field is used in this simulation. The initial conditions of Fig. 16 is the same as those of Fig. 15. In comparison of Fig. 15, Fig. 16 and Fig. 17, the difference of their velocities is shown. It can be seen that the maximum velocity of the new method is a half maximum velocity of our previous method. As a result, the effectiveness of our method by using the ellipse flow field can be shown. Fig. 18 shows the simulation with applying the correction function bj(z) in Fig. 15. And Fig. 19 shows the simulation with applying correction function bj(z) in Fig. 16. Similarity, Fig. 20 also shows the difference of their velocities. It can be seen that the case of Fig. 19 is smoother than the case of Fig. 18. The blue line as shown in Fig. 17 has a discontinuous point at the maximum velocity point. However, the blue line as shown in Fig. 20 has no discontinuous point at all time. As a result, the effectiveness of our method by using the new correction function can be shown. Is the robot motion considered? If the robot is just in front of the obstacle, how does the robot move? These answers are very simple. That is, the robot can move while touching the avoidance circle of the obstacle",
            " 20 come from both the ellipse field and the correction function, and they are only influenced by the parameters of m and uj . However, the m and the uj are the same value, respectively in each simulation. The m denotes a sink value and the standard robot velocity comes from the m. That is, the m can not be changed. This can be seen by the vertical constant velocity lines shown in Fig. 17 and Fig. 20. On the other hand, the uj denotes the obstacle velocity. That is, the uj can not be changed too. All distances of many circles of the moving obstacles shown in both Fig. 15 and Fig. 16, and also both Fig. 18 and Fig. 19 are the same, respectively. Therefore, the changing velocities of these simulations come from the other factors. The factors are using both the ellipse field and the correction function. In this paper, we proposed the improved method by applying the Conformal Transformation and the new correction function to our previous Hydrodynamic Potential method for path planning of a mobile robot to avoid the moving obstacle smoothly. A mobile robot can gradually avoid a moving obstacle from further away, and can be safely guided without rapid acceleration"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001931_iros.2010.5649323-Figure15-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001931_iros.2010.5649323-Figure15-1.png",
        "caption": "Fig. 15. The Flow Field by using the Previous Method",
        "texts": [
            " That is, using an avoidance ellipse is more safety than using avoidance circle. Therefore, hereinafter, we decide to use the style of Fig. 14(d) for avoiding the moving obstacle. The flow field to the preceding chapter is drawn without actual move of the obstacle. The reason is that the characteristic of the flow field with the moving obstacle cannot be displayed simply. However, the moving obstacle will move itself in the real environment. Therefore, we investigate the usefulness of the new method in the simulation. Fig. 15 shows our previous simulation by using Hydrodynamic Potential where a mobile robot avoids one moving obstacle and can reach the goal. In this figure, many circles display the location of both the mobile robot and the moving obstacle respectively at the constant interval. Before the start of the simulation, the obstacle stands the right side, and the mobile robot stands the left side in the simulation display area. Moreover, the goal is set at the right side. The moving obstacle moves to the left direction and the mobile robot moves to the right direction after starting this simulation. If the distance between two circles is long, it is shown that the velocity is fast. In the intersection on near the center of display area, it can be seen that the mobile robot can avoid the moving obstacle. However, the velocity of the mobile robot accelerates rapidly when it avoids a moving obstacle. Fig. 16 shows the simulation with the new method by using CT. That is, the ellipse flow field is used in this simulation. The initial conditions of Fig. 16 is the same as those of Fig. 15. In comparison of Fig. 15, Fig. 16 and Fig. 17, the difference of their velocities is shown. It can be seen that the maximum velocity of the new method is a half maximum velocity of our previous method. As a result, the effectiveness of our method by using the ellipse flow field can be shown. Fig. 18 shows the simulation with applying the correction function bj(z) in Fig. 15. And Fig. 19 shows the simulation with applying correction function bj(z) in Fig. 16. Similarity, Fig. 20 also shows the difference of their velocities. It can be seen that the case of Fig. 19 is smoother than the case of Fig. 18. The blue line as shown in Fig. 17 has a discontinuous point at the maximum velocity point. However, the blue line as shown in Fig. 20 has no discontinuous point at all time. As a result, the effectiveness of our method by using the new correction function can be shown",
            " 20 come from both the ellipse field and the correction function, and they are only influenced by the parameters of m and uj . However, the m and the uj are the same value, respectively in each simulation. The m denotes a sink value and the standard robot velocity comes from the m. That is, the m can not be changed. This can be seen by the vertical constant velocity lines shown in Fig. 17 and Fig. 20. On the other hand, the uj denotes the obstacle velocity. That is, the uj can not be changed too. All distances of many circles of the moving obstacles shown in both Fig. 15 and Fig. 16, and also both Fig. 18 and Fig. 19 are the same, respectively. Therefore, the changing velocities of these simulations come from the other factors. The factors are using both the ellipse field and the correction function. In this paper, we proposed the improved method by applying the Conformal Transformation and the new correction function to our previous Hydrodynamic Potential method for path planning of a mobile robot to avoid the moving obstacle smoothly. A mobile robot can gradually avoid a moving obstacle from further away, and can be safely guided without rapid acceleration"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001901_1.3591479-Figure8-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001901_1.3591479-Figure8-1.png",
        "caption": "Fig. 8 Spatial four-bar mechanism wi th one turning pair and threa cyl indrical pairs",
        "texts": [
            " Spatial Four-Bar Mechanism With One Turning and Three Cylindric Joints The input-output (crank rotation) equations of the mechanism date back to Dimentberg [6] and can be found also in Denavit [5], Yang and Freudenstein [31] and Woerle [28]. A discussion of the coupler curve in a parametric form (which, however, is not rational) has been given by Worele [27], who has also given some diagrams of the coupler curves and has pointed out that the curve may have a branch extending to infinity. In the discussion, which follows, we shall limit ourselves to the determination of the order and genus of the curve described by a point M on the output link. The mechanism is shown schematically in Fig. 8. The turning joint is at A and the cylindrical pairs are at B, C, D. The perpendicular distances (\u00ab,,) between the axes of the joints and their angular offsets (a,-,) are fixed-linkage dimensions. The sliding along the axes of the joints (sf) and the relative rotation between the links ((?,\u2022) are variable, except for si, which is associated with the fixed link. The rotation, 91, of the turning joint is usually regarded as the input and the motion (9.1, s4) as the output. Motion of Output Link. F o r the m o t i o n of a p o i n t Mix, y, z) 011 the output link, we may choose x = d \u2014 st, y = a sin $,, z = a cos O't, where a is the radius of the cylinder on which the curve lies"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002112_robot.2009.5152261-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002112_robot.2009.5152261-Figure1-1.png",
        "caption": "Fig. 1. Description of the movement of the follower in the diagonal formation. The blue leader moves randomly and the red follower keeps the specified distance (ld) and angle (\u03b3) at each time step. The distance l and the bearing \u03b1 are computed through the image measurements.",
        "texts": [
            "00 \u00a92009 IEEE 351 monocular camera mounted on each robot, ii) we assume that the robots are not communicating with each other, and iii) we have no pre-specified marks for recognizing the leader. In the line formation the follower tries to keep the specified distance and bearing from the leader. In the diagonal formation the robot moves diagonally keeping a specified angle from the leader. The diagonal formation in our work is defined as a follower moving diagonally with a certain, variable angle from the measured relative bearing to the leader as shown in Fig. 1. Changing the follower\u2019s position diagonally to the leader allows a follower to estimate the leader\u2019s velocity unlike the line formation case. This is accomplished by applying input-output feedback linearization, a well known methodology in control theory. We also present the visibility issues for a monocular camera and the estimate of the leader\u2019s velocity. Without knowing or estimating the linear and angular velocities of the leader, the formation controls may result in missing the leader in the next time step"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000154_iros.2007.4399278-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000154_iros.2007.4399278-Figure1-1.png",
        "caption": "Fig. 1. Vehicle schematic for vertical flight mode of Bidule",
        "texts": [
            " In recent years, interest in Vertical Take-Off and Landing (VTOL) mini Air vehicles (mAVs) have increased significantly due to a desire to operate UAVs in an urban environment. Many concepts have been proposed globally [4]. The Bidule mAV was developed at the University of Sydney to explore design issues related to small flight platforms [5]. The latest version, the Bidule CSyRex, is a joint project between the University of Sydney and the University of Technology of Compie\u0300gne to develop a VTOL variant of the Bidule. The vertical flight schematic of this VTOL vehicle is shown in Figure 1, which is basically a fixed wing tailless aircraft with two propellers. In hover the altitude is controlled with the collective thrust, this means, the lift force is generated increasing the speed of the propellers. The pitch attitude angular displacement is achieved by moving the elevons in the same direction. The vertical yaw-attitude angular displacement is achieved through moving the elevons in opposing direction. The vertical roll-attitude angular displacement is controlled by changing the pitch angle of the Variable Pitch Propeller (VPP)",
            " The transformation of the components of the angular velocity generated by a sequence of Euler rotations from the body to the local reference system is written as follows: H (\u03a6) =   1 t\u03b8s\u03c6 t\u03b8c\u03c6 0 c\u03c6 \u2212s\u03c6 0 s\u03c6/c\u03b8 c\u03c6/c\u03b8   (6) where s and c are used to denote the sin and the cos respectively, then using (6), the kinematic equations (3) can be rewritten as: \u03c6\u0307 = P + tan \u03b8 (Q sin \u03b8 +R cos\u03c6) (7) \u03b8\u0307 = Q cos\u03c6\u2212R sin\u03c6 (8) \u03c8\u0307 = (Q sin\u03c6+R cos\u03c6) / cos \u03b8 (9) The term Jb in (5) represents the inertia matrix, and is defined by Jb =   Jx Jxy Jxz Jyx Jy Jyz Jzx Jzy Jz   If the Bidule CSyRex can be assumed to have the body axis xz-plane coincident with the plane of symmetry, then the products of inertia Jxy and Jyz vanishes. This tail-sitter configuration, also presents a plane of symmetry in the yzplane, then the product of inertia Jxz = 0. Then Jb and its inverse can be written by Jb =   Jx 0 0 0 Jy 0 0 0 Jz   and (Jb)\u22121 =   1 Jx 0 0 0 1 Jy 0 0 0 1 Jz   Note that the mass of the elevons is neglected. The aerodynamics and thrust moments can be denoted by M b A,T = [ \u2113 m n ]T , they are shown in the Figure 1, then using the matrix of inertia and the moment vector, the equation (5) yields: P\u0307 = (Jy \u2212 Jz)QR Jx + \u2113 Jx (10) Q\u0307 = (Jz \u2212 Jx)RP Jy + m Jy (11) R\u0307 = (Jx \u2212 Jy)PQ Jz + n Jz (12) This section presents three decoupled stability augmentation control systems for the roll, the pitch, and the yaw positions of the vehicle in hover flight. These subsystems will be obtained using only the kinematics and moment equations from the general model. Several aerodynamic factors will be taken into account to obtain the transfer function that represents the dynamic of each system"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001874_1.3103805-Figure3-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001874_1.3103805-Figure3-1.png",
        "caption": "Figure 3. Eletrochemical measurement cell for the biofuel cell (open-air type cell).",
        "texts": [
            " Characterization of the CF-biocathode We measured the catalytic current of the CF-biocathode by using an open-air type cell (Figure 2), which is exposed to air. In this cell, O2 can diffuse from air to the CFbiocathode efficiently. At the same time with the electrochemical measuremnt, the local pH values on the CF-biocathode was evaluated by contacting a pH meter (TYPE PCE308S-SR, Toko Chemical Laboratory Ltd., Japan) with a flat surface on the CFbiocathode. Characterization of the biofuel cell We constructed a passive-type biofuel cell (Figure 3) in which cellophane was sandwiched between CF-bioanode and open-air type CF-biocathode. A 1.0 M PBS (pH 7.0) and a 2 M IBS (pH 7.0) containing 0.4 M glucose were used as the fuel solution to characterize the biofuel cells. 11 ) unless CC License in place (see abstract).\u00a0 ecsdl.org/site/terms_use address. Redistribution subject to ECS terms of use (see 138.251.-85.28Downloaded on 2015-02-13 to IP Electrochemical measurement and local pH measurement of the biocathode The power densities of biofuel cells reported (1-5) are in the order of sub-mW/cm2 even with fuels- and O2-convective system such as mechanical stirring or a pump, resulting in a few \u00b5W per single biofuel cell",
            "5 mW/cm2 at 300 mV (black) (6, 7). We have further reported a 3.0 mW/cm2 at 500 mV (red) by introducing 2-amino-1,4-naphthoquinone, ANQ, which has lower potential than conventional VK3, as anode mediator (8) (Figure 5). These results were obtained by using the PBS (1 M, PBS) system. In the case of using the new electrolyte solution system of the high concentrated IBS (2 M, pH 7.0), the maximum power density has reached 5.0 mW/cm2 at 500 mV (blue). The cathodic current densities estimated at each potential of the cell (Figure 3) were almost identical with those with the electrochemical measurement cell for the biocathode (Figure 2). The current densities of the bioanode were also enhanced in the same way of biocathode (Figure 5). This implies that the IBS system is also effective to maintain the cell potential even under mA current density as well as the PBS system. 13 ) unless CC License in place (see abstract).\u00a0 ecsdl.org/site/terms_use address. Redistribution subject to ECS terms of use (see 138.251.-85.28Downloaded on 2015-02-13 to IP 14 ) unless CC License in place (see abstract)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003954_213548-Figure7-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003954_213548-Figure7-1.png",
        "caption": "Figure 7: In this picture, the Reynolds number (ratio between inertial forces and viscous forces) is plotted for the journal surface.",
        "texts": [
            " By creating the model in this way, the shaft is always centered and no further errors occur due to the geometrical tolerances that are introduced. Then, a cylinder with diameter \ud835\udc51 is attached to the external face of the lubricant domain.The axis of the cylinder passes through the centre of the bearing and lies on the \ud835\udc65-\ud835\udc66 plane. The cylinder is then rotated on \ud835\udf03 \ud835\udc60 (positioning angle on the inlet). There are two assumptions for the simulations: laminar flow and Reynolds cavitation model. The laminar flow assumption allows the avoidance of the use of turbulence models.The assumption is also verified in Figure 7 where the Reynolds number is shown (Reynolds number under 30 denotes full laminar regime). by guest on June 23, 2016ade.sagepub.comDownloaded from The boundary conditions are the following. (i) Inlet: the inlet condition is \u201cpressure inlet\u201d with gauge total pressure 140KPa and initial gauge pressure 0 Pa. The inlet temperature is set to \ud835\udc47in. (ii) Outlet: the outlet condition is \u201cpressure outlet\u201d with static pressure set to 0 Pa. It should be pointed out that the operating pressure is set to 101325 Pa, so all the pressures at the boundaries are added to the operating pressure (absolute pressure = relative pressure + operating pressure)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001751_icca.2010.5524421-Figure5-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001751_icca.2010.5524421-Figure5-1.png",
        "caption": "Fig. 5. Kinematic analysis of secondary projection system (a) physical model (b) Kinematic diagram",
        "texts": [
            " This passive manipulator can be installed with IntersenseTM wireless orientation tracker similar to the serial manipulator. The forward kinematics with respect to the primary projector can therefore be computed. The homogenous matrix relating the secondary projector to the primary projector can be represented as: ++\u2212\u2212\u2212+ ++\u2212\u2212\u2212 ++++ = 1000 22233234523452345234 212231323415234151523415152341 212231323415234151523415152341 5 0 sdsdcrcsscs csdcsdssrssccsssscccs ccdccdscrsccssccssccc T (12) where sijk=sin (qi+qj++qk), cijk=cos(qi+qj+ qk). Fig. 5 shows the kinematic analysis of the manipulator arm for the secondary projector. The derivation of the kinematic model for the needle insertion serial manipulator can be found in [2]. A. Intraoperative Visual Guidance While sophisticated medical imaging modules are available for comprehensive diagnostic and preoperative plans, the effectiveness of surgical treatment lies in the ability to execute the plan with consistency. As such, intraoperative guidance plays an important role in the success of image guided therapies such as ablation of large tumor"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003236_s11837-013-0614-3-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003236_s11837-013-0614-3-Figure1-1.png",
        "caption": "Fig. 1. (a) Scheme of multilayer microwire. SEM images of FeSiB/ CoNi multilayer microwires: (b) FeSiB nucleus and glass cover, (c) glass layer and CoNi outer shell, (d) CoNi outer shell, and (e) crosssection view of the multilayer microstructure.",
        "texts": [
            " They diversify the magnetic response of the MM at a high frequency involving interesting effects as asymmetric magnetoimpedance and multiabsorption ferromagnetic resonance. This article revises most relevant magnetic properties of bimagnetic microwires focusing on: (I) biphase behavior and interphase magnetoelastic and magnetostatic interactions, (II) asymmetric magnetoimpedance and multiabsorption FMR effects, and (III) multilayer biphase microwires as sensing elements in various devices. The biphase magnetic microwires are composed by a ferromagnetic nucleus, an intermediate glass layer, and a ferromagnetic outer shell (Fig. 1a). They are fabricated by the combination of ultrarapid solidification and electroplating techniques. The ferromagnetic nucleus covered by a Pyrex layer is obtained by the quenching and drawing method3,7 (Fig. 1b). The diameter of the nucleus and thickness of the glass cover range between 1 lm and 20 lm. Due to the strong quenching rate of approximately 105 K/s, the ferromagnetic nucleus typically exhibits amorphous structures conferring soft magnetic properties. Three different kinds of magnetic alloys with soft behavior have been considered according to its saturation magnetostriction value: Fe-based (ks = 10 5), Co-based (ks = 10 6), and CoFe-based alloys with vanishing magnetostriction (ks = 10 7). Several attempts were made to obtain microwires with harder magnetic behavior21\u201323 using different compositions as CoFeCr, CoNiCu, and CoFeMo after suitable temperature treatment to promote the crystallization of magnetic phases with strong crystalline anisotropy",
            " Others alloys (FeNdB, FePt, FePtNdB, FePd, and CoSm) were also explored; however, their saturation magnetization significantly decreases after the treatment at a high temperature. In this work, we show, as an example, results on FePt microwire after suitable thermal treatment: The transition from face-centered cubic (fcc) disorder to face-centered tetragonal (fct) ordered phase enhardens its magnetic behavior.19,24 An Au nanolayer (typically 30 nm thick) is grown on top of the glass surface using commercial sputtering system to serve as an electrode for the subsequent electroplating of the magnetic outer layer (Fig. 1c\u2013e). Typically, two different magnetic alloys have been selected as magnetic outer layer: FeNi10,24,25 and CoNi15,17 with soft and hard magnetic behavior, respectively. The nominal composition of the alloys is tailored through the current density of the electroplating, whereas the thickness is almost proportional to both time and current density. Other parameters affecting the quality and characteristics of the electroplating are the electrolyte temperature (optimal temperature of 40 C for these alloys) and the mechanical stirring"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003246_1.4004116-Figure6-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003246_1.4004116-Figure6-1.png",
        "caption": "Fig. 6 Geometric sketch of the left tooth profile of noncircular gear generated by shaper cutter",
        "texts": [
            " In theory, the tooth profile of noncircular gears is enveloped by a shaper when its pitch circle rolls on the pitch curves of noncircular gears without sliding. Many efforts are devoted to build the mathematical model for tooth profile of noncircular gears. In the sections below, the tooth profile formulas of deformed limacon gear pair are deduced on the basis of the mathematical model in literature [20]. 3.2.1 Tooth Profile Formulas. The geometric graphics of noncircular gears manufactured by shaper cutter is shown in Fig. 6. P1P2 is tangent to the pitch circle of the shaper and the pitch curve of the noncircular gear through A. The instantaneous meshing point of their tooth profile is the point N. Line NA is the normal of tooth profile. Point A1 is the crossover point of the tooth profile of noncircular gear and its pitch curve, while A2 is the crossover point of the tooth profile of the shaper and its pitch circle. The angle between lines AN and P1P2 is the profile angle a of the shaper cutter. The following vector equation is derived based on the geometric relationship in Fig. 6 ON \u00bc OA\u00fe AN (24) If the pitch curve formula of deformed limacon gear is r \u00bc r\u00f0u\u00de, the vector angle and module of OA is u and r, respectively. The module of vector AN equals to the distance between the instantaneous rotation center and meshing point of tooth profile. Figure 7 shows the meshing relationship of a circle gear and a rack. The pitch circle is tangent to the pitch line at the point B. Line BN is perpendicular to tooth profile of the rack. The tooth profile of the circular gear and the pitch circle intersect at B1, while the tooth profile of the rack and the pitch line intersect at B2",
            " 3 Pitch curves of two kinds of gear pairs Fig. 4 Transmission ratio curves of two kinds of gear pairs 061004-4 / Vol. 133, JUNE 2011 Transactions of the ASME Downloaded From: http://mechanicaldesign.asmedigitalcollection.asme.org/pdfaccess.ashx?url=/data/journals/jmdedb/27948/ on 03/26/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use The length of profile normal BN can be derived by jBNj \u00bc jB2Bj cos a \u00bc B1B _ cos a, where a is the profile angle of the rack. Similarly, the length of profile normal AN in Fig. 6 can be represented as jANj \u00bc A2A _ cos a \u00bc A1A _ cos a. If the intersection point of tooth profile and profile normal is outside the pitch curve of noncircular gear as shown in Fig. 6, the vector angle of AN is u\u00fe l a. On the contrary, if the point is inside the pitch curve as shown in Fig. 8, the vector angle of AN is represented as u\u00fe l a\u00fe p. Using Eq. (24), the formulas of left tooth profile of noncircular gears can be derived as follows xNL \u00bc r cos u6 A1A _ cos a cos\u00f0u\u00fe l a\u00de yNL \u00bc r sin u6 A1A _ cos a sin\u00f0u\u00fe l a\u00de ( (25) where tan l \u00bc r=r0\u00f0u\u00de. Similarly, the right profile formulas are represented as follows xNR \u00bc r cos u A1A _ cos a cos\u00f0u\u00fe l\u00fe a\u00de yNR \u00bc r sin u A1A _ cos a sin\u00f0u\u00fe l\u00fe a\u00de ( (26) The upper and lower signs in Eqs"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001927_nme.1620030405-Figure3-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001927_nme.1620030405-Figure3-1.png",
        "caption": "Figure 3. Central cap element",
        "texts": [
            " On assembling the elements into a representation of the over-all shell, compatibility must be maintained for all the displacement degrees of freedom occurring at the interzelement nodes, i.e. at i and j in Figure 2. The two displacement degrees of freedom at the internal nodes m and n are not required in the assemblage. Thus they must be removed by a process of static condensation prior to assemblage of the total structural stiffness matrix. 498 E. P. POPOV AND P. SHARIFI The displacement model (2) must be specialized for the case of the central cap, Figure 3. Here, if Wand U are the radial and tangential components of displacements at the apex, because of symmetry, the tangential component of displacement, Ui, and rotation, xi, vanish. The fulfilment of the symmetry conditions reduces the number of generalized co-ordinates from eight to six for the cap element. For these reasons it can be shown that the displacement pattern in 8 - q co-ordinates is ANALYTICAL PROCEDURE Element stiffness matrix The details of generating the element stiffness matrix for an axisymmetrical curved element of the type considered here have been discussed el~ewhere"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001888_0022-2569(71)90007-3-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001888_0022-2569(71)90007-3-Figure1-1.png",
        "caption": "Figure 1.",
        "texts": [],
        "surrounding_texts": [
            "Let M , M., ,4. ,A ~ (Fig. I ) be the sinlply ~,kew four-bar. M I M., = ~. MI ~A z = R1. M . , A ._, = R._,, A~A. , = b. The links M~A~ and M.>_A._, move in the planes I<\"~ and V.., respectively. passing through M~M. , ; the angle between I/~ and V., is a = 2/3.0 -< ~< ~ ~/2. InA~ and A2 the l i nkage has s p h e r i ca l j o i n t s . T a k e M~M.,. as the Y-axis of a C a r t e s i a n f r ame , the o r ig in O in the m i d p o i n t o f M ~ M 2 a n d let X O Y a n d Z O Y c o i n c i d e wi th the b i s ec t i ng p l a n e s o f Vt a n d V.,_. W e d e n o t e the ang le A t M t M . , . by ~ a n d A. , .MzMt by \u00a20_. I f A i = (x~. y~. zi we have x~ = Rt s i n ~ cos /3 , y~ = R~ cos ~ t - \u00bd g , z.t = R~ s i n ~ s in /3 , x2 = - R._, s in ~,._, cos /3 , y._, = - R.,_ cos ~z + \u00bdg. co = R., s in ,\u00a2., s in /3 , (2.1) or. if .~_ 2tti cos ~i -- 1 -- tt~ ~ 9 tg _ = tt t . s in ~i = 1 + l t t \" \" I -q- Ill\" (2.2) a n d m a k i n g use o f h o m o g e n e o u s c o o r d i n a t e s xt = 2Rt cos /3 . tq . 3'1 = Rt ( 1 - - lit 2) _ l g ( 1 + 16\"-'), Zt = 2Rt s i n / 3 . It t, w t = 1 + ltl 2, x._, = - 2R._, cos /3 . u._,. y., = - R2 ( 1 - u.,.\"- ) + \u00bdg ( 1 + u.,.\" ) . z., = 2R., s i n / 3 , u._,. w.z = 1 + u..,-."
        ]
    },
    {
        "image_filename": "designv11_12_0003001_00405001003696464-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003001_00405001003696464-Figure2-1.png",
        "caption": "Figure 2. Woven fabric: (a) geometrical model created by WiseTex, and (b) tensile diagram of the wires.",
        "texts": [
            " With the wires with non-deformable cross-sections (in contrast with fibrous textile yarns, which have easily deformable cross-sections), the natural way to handle the interpenetration is to define the contact pairs at places of contact with the yarn. Finally, the FE problem is solved using the nonlinear geometry option and an appropriate number of the load steps. The result is a stress\u2013strain diagram of the textile, which can be used in macro-modelling as a material model of the homogenised behaviour of the fabric, and details of the stress\u2013strain state of the wires. Material and geometrical model of woven fabric Table 2 and Figure 2(a) show input parameters of the woven fabric made of NiTi wires, the resulting parameters of the fabric calculated by WiseTex and the ready geometrical model. The input parameters were deliberately chosen in such a way that the wires in the fabric have a relatively high crimp ratio of 14.2 (ratio: ( l y \u2212 l f ) / l f , where l y is length of a wire and l f is the corresponding length of the fabric). A high crimp ratio will make the difference between the tensile resistance of the fabric and that of the wires more visible. The assumed tensile diagram of the wires is shown in Figure 2(b). Figure 2. Woven fabric: (a) geometrical model created by WiseTex and (b) tensile diagram of the wires. Approximate model of biaxial tension of woven fabrics The model of bi- and uniaxial tension of woven fabrics in its full form is applicable to complex weave structures and to compressible yarns. Here, it is formulated for the simple plain weave fabric described in the previous section, with uncompressible wires with nonlinear tension and bending. The reader is referred to Lomov (2007), Lomov and Verpoest (2006) and Lomov et al. (2003) for the full formulation of the model and discussion of the assumptions. The model for plain weave fabrics has been formulated long ago (Hearle & Shanahan, 1978; Kawabata, Niwa, & Kawai, 1973) and is given here for clarity of its application for SMA fabrics. Consider a fabric shown in Figure 2(a) under biaxial tension with the applied deformations \u03b5 1 and \u03b5 2 in the Table 1. \u201cRoad map\u201d for FE modelling of SMA textiles. Action Input Result Comments Creation of a geometrical model of the textile Diameter of the yarns, weave/knit pattern, yarn spacing, average bending rigidity of the yarns Geometrical model of the yarns in the unit cell of the fabric, defined in a format of the geometry modeller See Lomov, Mikolanda, et al. (2007), Lomov, Ivanov et al. (2007) and Verpoest and Lomov (2005) for WiseTex data format Building solid model of the yarns Geometrical model, dimensions of the unit cell, tensile diagrams of the wires Solid model (volumes, surfaces, points), confined in the unit cell \u201cbox\u201d, material properties assigned In the WiseTex package, use an FETex tool to create a solid model Meshing of the yarns (solid elements) Solid model Mesh Meshing engine of the FE package is used",
            " Equation (2) allows calculating the curvature of the wires, the average deformation of the wires, and the angle of inclination of the centre lines of the wires Equations (2)\u2013(7) have two unknowns \u2013 crimp heights h 1 and h 2 \u2013 related with the constraint Equation (3). They are calculated from the condition of minimum energy of the whole structure (de Jong & Postle, 1978; Hearle & Shanahan, 1978): where \u201cbend\u201d and \u201ctens\u201d refer to the energy of bending and tensile deformations, respectively. Assuming that Equation (6) represents the average tensile deformations of the wire, the tension energy is calculated as where \u03c3 ( \u03b5 ) is the tensile diagram shown in Figure 2(b). Note that this assumption neglects the distribution of strain over the wire cross-section. The bending energy is calculated as where B ( \u03ba ) is the bending rigidity (units [F][L] 2 ) of the wire corresponding to its curvature. z x h x p x p z x h y p y p 1 1 2 3 2 2 2 2 1 3 1 2 2 4 6 1 2 4 6 1 2 ( ) =       \u2212       +         ( ) =       \u2212       +         ; ( ) h h d1 2 2 3+ = ( ) p p p p l l p h l l p h l x z p 1 0 1 2 0 2 1 2 1 2 1 2 2 1 2 0 1 1 1 4 = +( ) = +( ) = ( ) = ( ) = + \u2032( )    \u222b \u03b5 \u03b5; , ; , ; ( / d ) \u03ba( ) ( / x p z z x p = \u2032\u2032( ) + \u2032( )      \u222b1 1 5 2 2 5 2 0 d ) \u03b5 \u03b51 1 0 0 2 2 0 0 0 0 0 6y yl l l l l l l l p h= \u2212( ) = \u2212( ) = ( )/ ; / ; , ( ) \u03b8 \u03b81 2 2 12 2 7= \u2032( ) = \u2032( )arctan / , arctan / ",
            " ( 2 2 2 2 2 15 d d d d ) D ow nl oa de d by [ U ni ve rs ity o f C al if or ni a, S an D ie go ] at 1 5: 12 1 3 M ar ch 2 01 5 The Journal of The Textile Institute 237 With the solution of the problem known, the bending stiffness is calculated from Equation (14) as It can be expressed as a dimensionless number, normalised by the elastic bending rigidity Bel (Equation (13)) The FE calculations were performed with ANSYS, using solid elements SOLID45, with the element size r/ 5, for \u03bb=5 and \u03bb=10, in the curvature range The value of the total energy W in Equation (16) is provided by ANSYS. The calculations for the elastic material of the wire result in the values of B \u2013 in the range 0.97\u20130.99, which gives an estimation of the accuracy of the FE model. For the superelastic material, with the tensile diagram shown in Figure 2(b), the bending stiffness depends heavily on the curvature of the wire (Figure 5). The average of the two curves shown in Figure 5 is used for the calculation of the biaxial tension of the SMA woven fabric. Figure 5. Bending rigidity of the superelastic wire (tension diagram shown in Figure 2(b)) as function of the curvature for two values of the elongation. Figure 6 shows the computed diagram of uniaxial tension of the woven SMA fabrics with two different values of the fabric density. The diagram is expressed as force per wire and is compared with the tensile diagram of the wire. The tensile diagram of the fabric is the result of the combination of the tensile resistance of the wires and their (de)crimping and hence the bending resistance. Several features of the fabric tensile diagram should be noted: Figure 6 also shows the computed diagram of biaxial tension of the woven SMA fabric in comparison with the diagram of uniaxial tension and with the tensile diagram of the wire",
            " This assumption takes into account the distribution of strain in the wire cross-section only approximately. By the integral effect, this distribution has on the overall bending rigidity. The FE modelling, discussed next, will reveal more \u201csmoothened\u201d tensile diagrams. Building and solving an FE model of woven fabric Figure 7(a) shows ANSYS FE model of the woven SMA fabric. Note that the fabric geometry is symmetrised for the transition to FE model from WiseTex. The material model is defined as multilinear elastic, with the tensile diagram shown in Figure 2(b). The elements used are 3D structural eight-node solid elements of the type SOLID45 (prismatic). Figure 7. FE model of woven SMA fabric: (a) the mesh and (b) a contact pair.Contact pairs are defined at the contact zones on the yarns, as shown in Figure 7(b). There is important interpenetration of the volumes of the yarns. This is a common difficulty for the FE modelling of textile structures. Geometrical models, like the models used here, use several simplifying assumptions. The shape of the yarn\u2019s central line prescribes the positions of the centres of the cross-sections"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001119_j.molcatb.2008.04.006-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001119_j.molcatb.2008.04.006-Figure1-1.png",
        "caption": "Fig. 1. Proposed mechanism of silyl ether cleavage by serin",
        "texts": [
            "006 ytic competence of the enzymes. \u00a9 2008 Elsevier B.V. All rights reserved. urantia was able to catalyse the polymerisation of silicates, in vitro, sing the xenobiotic alkoxysilane tetraethoxysilane, as a substrate, ia cleavage of the Si\u2013O bond of the molecule [7]. Mutation experments on the silicatein- of T. aurantia have demonstrated that i\u2013O bond cleavage by this enzyme is catalysed by a classical serne hydrolase triad, that features in a protein whose sequence is omewhat related to that of mammalian cathepsins (Fig. 1) [8]. Other experiments have also demonstrated Si\u2013O bond cleavage n microorganisms. Semprini and co-workers described the enrich- ent, in a chemostat, of a microbial consortium that was able to egrade tetraethoxysilane by apparent cleavage of the Si\u2013O bond 9] and Fattakhova et al. demonstrated that the yeast Rhodotorula ucilaginosa produced a substrate-inducible esterase that catalsed the cleavage of the Si\u2013O bond in a range of silatrane substrates 10,11]. Most recently, Bassindale et al. have shown that serine ydrolases such as trypsin and chymotrypsin appear to catalyse he formation of siloxane bonds at the active sites of the enzymes, et the hydrolytic reaction, whilst accelerated in the presence of he protein, is promoted only by non-specific interactions with the nzyme in use [12]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000963_ecc.2007.7068846-Figure7-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000963_ecc.2007.7068846-Figure7-1.png",
        "caption": "Fig. 7. Picture of the DC motor platform",
        "texts": [],
        "surrounding_texts": [
            "In this section the experimental platform and the experimental design are briefly explained. The identification method previously described is applied to a real DC motor. The experimental results obtained are described and validated."
        ]
    },
    {
        "image_filename": "designv11_12_0001465_robot.2010.5509554-Figure4-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001465_robot.2010.5509554-Figure4-1.png",
        "caption": "Figure 4. Gradient variable limits for the first wheelchair trajectory stage",
        "texts": [
            " Similar to the arm, mathematical representations can be obtained by treating the range of desired wheelchair motion as a motion limit. The upper limit in this case is set to be the current initial orientation (or position for the second trajectory stage) of the wheelchair. The lower limit is set to be double the rotation angle \u03b21 or \u03b22 (or double the translation distance tr for the second trajectory stage). In this case, the middle of that range will be the desired orientation/position of the wheelchair, and either limit will be avoided. Figure 4 shows an example of the limits for the first wheelchair trajectory stage. To generalize the representation of the objective function, let variable \u201cP\u201d be a representative for \u03b21, \u03b22 or tr. The objective function in this case is: 2 max min max current current min (P P )1L(P) 4 (P P ) (P P ) \u2212 = \u22c5 \u2212 \u22c5 \u2212 (13) and the gradient of the criterion function can be defined as: 2 max min current max min 2 2 max current current min (P P ) (2 P P P )L(P) P 4 (P P ) (P P ) \u2212 \u22c5 \u22c5 \u2212 \u2212\u2202 = \u2202 \u22c5 \u2212 \u22c5 \u2212 (14) For the first stage, when the wheelchair\u2018s angle is in the middle of its allowable range, (14) becomes zero, and when it is at its limit, (14) becomes \u201cinfinity\u201d, which means that the variable will carry an infinite weight that makes it impossible to move any further"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003501_2041302510394742-Figure8-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003501_2041302510394742-Figure8-1.png",
        "caption": "Fig. 8 Hybrid aircraft model",
        "texts": [
            " Two Servo-motor 18 drive a pair of threaded screw rods 15, respectively, and their rotating directions are different. Thus, threaded collars 16 move along threaded screw rods 15. Finally, free wings 2 can keep a given angle. The four tail fins are used to regulate yaw and pitch dynamics during forward-flight mode, and the angles of four tail fins remain zero during mode transition. Therefore, the authors ignore their effects during mode transition. 3 MODELLING OF HYBRID AIRCRAFT The forces and moments for the aircraft are shown in Fig. 8. In Fig. 8, C is the centre of gravity of the aircraft. V is the flying velocity of the centre of weight. M is the pitch moment generated by gearing arrangement 19 for the free wings 2. L1 and D1 are the lift and drag forces, respectively, generated by the wings, \u03b8 is the pitch angle of the fuselage, \u03b1 is the angle of attack for the fuselage. 3.1 Thrust by front counter-rotating propellers The thrust Tf is generated by co-axial counter-rotating propellers Tf = Tf0 + f (1) where Tf0 = kf 1\u03c9 2 f 1 + kf 2\u03c9 2 f 2 (2) \u03c9f 1 and \u03c9f 2 are, respectively, the two angular velocities of the front counter-rotating propellers",
            " Fyb, Fxb, and Fzb are the forces of the aircraft along three-axe directions in the body co-ordinate frame, h = |CK |. Let f1 = 4\u2211 i=1 fi,2 = f10 + 1r f2 = f1,1 + f2,1 = f20 + 2r f3 = f3,1 + f4,1 (11) Proc. IMechE Vol. 225 Part G: J. Aerospace Engineering 835 at The University of Manchester Library on April 23, 2015pig.sagepub.comDownloaded from Xinhua Wang and Hai Lin where f10 = 4\u2211 i=1 fi,2,0, 1r = 4\u2211 i=1 i,2 f20 = f1,1,0 + f2,1,0, 2r = 1,1 + 2,1 (12) 3.5 Motion equations of hybrid aircraft From Fig. 8, let = (Xg, Yg, Zg) denote the righthanded inertial frame, = (xb, yb, zb) denotes the frame attached to the body\u2019s aircraft whose origin is located at its centre of gravity (see Fig. 8). (X , Y , Z) denotes the translation co-ordinates relative to the inertial frame, and (\u03c8 , \u03b8 , \u03c6)T \u2208 3 describes the aircraft orientation expressed in the classical yaw, pitch, and roll angles (Euler angles).The orientation of the convertible aircraft is given by the orthonormal rotation matrix R = \u23a1 \u23a3c\u03b8 c\u03c8 s\u03c6s\u03b8 c\u03c8 \u2212 c\u03c6s\u03c8 c\u03c6s\u03b8 c\u03c8 + s\u03c6s\u03c8 c\u03b8 s\u03c8 s\u03c6s\u03b8 s\u03c8 + c\u03c6c\u03c8 c\u03c6s\u03b8 s\u03c8 \u2212 s\u03c6c\u03c8 \u2212s\u03b8 s\u03c6c\u03b8 c\u03c6c\u03b8 \u23a4 \u23a6 (13) where s\u03b8 = sin \u03b8 , c\u03b8 = cos \u03b8 . For the thrust force by front counter-rotating pro- pellers [ Tf 0 0 ]T in the body frame, one can obtain it in the inertial frame\u23a1 \u23a3c\u03b8 c\u03c8 c\u03b8 s\u03c8 \u2212s\u03b8 \u23a4 \u23a6 Tf (14) and for the thrust force by tail counter-rotating propellers [ Tt 0 0 ]T in the body frame, one can obtain it in the inertial frame\u23a1 \u23a3c\u03b8 c\u03c8 c\u03b8 s\u03c8 \u2212s\u03b8 \u23a4 \u23a6 Tt (15) For the force [\u2212f1 0 0 ]T in the body frame, one can obtain it in the inertial frame \u2212 \u23a1 \u23a3c\u03b8 c\u03c8 c\u03b8 s\u03c8 \u2212s\u03b8 \u23a4 \u23a6 f1 (16) For the force [ 0 0 f2 ]T in the body frame, one can obtain it in the inertial frame\u23a1 \u23a3c\u03c6s\u03b8 c\u03c8 + s\u03c6s\u03c8 c\u03c6s\u03b8 s\u03c8 \u2212 s\u03c6c\u03c8 c\u03c6c\u03b8 \u23a4 \u23a6 f2 (17) For the force [ 0 f3 0 ]T in the body frame, one can obtain it in the inertial frame\u23a1 \u23a3s\u03c6s\u03b8 c\u03c8 \u2212 c\u03c6s\u03c8 s\u03c6s\u03b8 s\u03c8 + c\u03c6c\u03c8 s\u03c6c\u03b8 \u23a4 \u23a6 f3 (18) The lift force and drag force in the body frame are, respectively\u23a1 \u23a3L1 sin \u03b1 0 L1 cos \u03b1 \u23a4 \u23a6 (19) and\u23a1 \u23a3\u2212D1 cos \u03b1 0 D1 sin \u03b1 \u23a4 \u23a6 (20) Therefore, the lift force and drag force in the inertial frame are, respectively\u23a1 \u23a3c\u03b8 c\u03c8s\u03b1 + (c\u03c6s\u03b8 c\u03c8 + s\u03c6s\u03c8)c\u03b1 c\u03b8 s\u03c8s\u03b1 + (c\u03c6s\u03b8 s\u03c8 \u2212 s\u03c6c\u03c8)c\u03b1 \u2212s\u03b8 s\u03b1 + c\u03c6c\u03b8 c\u03b1 \u23a4 \u23a6 L1 (21) and\u23a1 \u23a3\u2212c\u03b8 c\u03c8c\u03b1 + (c\u03c6s\u03b8 c\u03c8 + s\u03c6s\u03c8)s\u03b1 \u2212c\u03b8 s\u03c8c\u03b1 + (c\u03c6s\u03b8 s\u03c8 \u2212 s\u03c6c\u03c8)s\u03b1 s\u03b8 c\u03b1 + c\u03c6c\u03b8 s\u03b1 \u23a4 \u23a6 D1 (22) The equations of motion written in terms of the centre of mass C in the fixed axes of co-ordinate (X , Y , Z ) are then mX\u0308 = (Tf + Tt \u2212 f1)c\u03b8 c\u03c8 + f2(c\u03c6s\u03b8 c\u03c8 + s\u03c6s\u03c8) + f3(s\u03c6s\u03b8 c\u03c8 \u2212 c\u03c6s\u03c8) + L1(c\u03b8 c\u03c8s\u03b1 + (c\u03c6s\u03b8 c\u03c8 + s\u03c6s\u03c8)c\u03b1) + D1(\u2212c\u03b8 c\u03c8c\u03b1 + (c\u03c6s\u03b8 c\u03c8 + s\u03c6s\u03c8)s\u03b1) mY\u0308 = (Tf + Tt \u2212 f1)c\u03b8 s\u03c8 + f2(c\u03c6s\u03b8 s\u03c8 \u2212 s\u03c6c\u03c8) + f3(s\u03c6s\u03b8 s\u03c8 + c\u03c6c\u03c8) + L1(c\u03b8 s\u03c8s\u03b1 + (c\u03c6s\u03b8 s\u03c8 \u2212 s\u03c6c\u03c8)c\u03b1) + D1(\u2212c\u03b8 s\u03c8c\u03b1 + (c\u03c6s\u03b8 s\u03c8 \u2212 s\u03c6c\u03c8)s\u03b1) mZ\u0308 = \u2212(Tf + Tt \u2212 f1)s\u03b8 + f2c\u03c6c\u03b8 + f3s\u03c6c\u03b8 \u2212 mg + L1(\u2212s\u03b8 s\u03b1 + c\u03c6c\u03b8 c\u03b1) + D1(s\u03b8 c\u03b1 + c\u03c6c\u03b8 s\u03b1) Jzb\u03c8\u0308 = \u2212f3h Jxb\u03c6\u0308 = Mf 1 \u2212 Mf 2 + Mt1 \u2212 Mt2 Jyb \u03b8\u0308 = \u2212f2h + M + Mfw J1\u03b8\u03081 = \u2212M (23) where Jxb, Jyb, and Jzb are the three-axis moment of inertias, respectively",
            "3 Control of longitudinal dynamics The result of controlling the pitch angle \u03b8 is the longitudinal dynamics (\u03c6 = 0, \u03c8 = 0), which is described by mX\u0308 = (Tf + Tt \u2212 f1)c\u03b8 + f2s\u03b8 + L1(c\u03b8 s\u03b1 + s\u03b8 c\u03b1) + D1(\u2212c\u03b8 c\u03b1 + s\u03b8 s\u03b1) mZ\u0308 = \u2212(Tf + Tt \u2212 f1)s\u03b8 + f2c\u03b8 \u2212 mg + L1(\u2212s\u03b8 s\u03b1 + c\u03b8 c\u03b1) + D1(s\u03b8 c\u03b1 + c\u03b8 s\u03b1) Jyb \u03b8\u0308 = \u2212f2h + M + Mfw J1\u03b8\u03081 = \u2212M (28) Thus, the mathematical model of the longitudinal dynamics is shown in Fig. 11, where \u03b81 is the pitch angle of the free wing, \u03b11 is the attack of angle of the free wing, and \u03b3 is the track angle. Other parameters have been introduced in Fig. 8. Because \u03b3 = \u03b8 \u2212 \u03b1, one has mX\u0308 = (Tf + Tt \u2212 f1) cos \u03b8 + f2 sin \u03b8 \u2212 L1 sin \u03b3 \u2212 D1 cos \u03b3 mZ\u0308 = \u2212(Tf + Tt \u2212 f1) sin \u03b8 + f2 cos \u03b8 \u2212 mg + L1 cos \u03b3 \u2212 D1 sin \u03b3 Jyb \u03b8\u0308 = \u2212f2h + M + Mfw J1\u03b8\u03081 = \u2212M (29) Set u1 = (Tf 0 + Tt0)/mg, u2 = f20h/J \u03b5 = J /(hmg), \u03b51 = J1/J , u3 = M /J1 (30) x = X g , z = Z g , L = L1 mg , D = D1 mg (31) 1 = ( f + t \u2212 f1) cos \u03b8 + 2r sin \u03b8 mg 2 = \u2212( f + t \u2212 f1) sin \u03b8 + 2r cos \u03b8 mg 3 = Mfw J (32) The equations of motion finally read x\u0308 = u1 cos \u03b8 + \u03b5u2 sin \u03b8 \u2212 L sin \u03b3 \u2212 D cos \u03b3 + 1 z\u0308 = \u2212u1 sin \u03b8 + \u03b5u2 cos \u03b8 + L cos \u03b3 \u2212 D sin \u03b3 \u2212 1 + 2 \u03b8\u0308 = \u2212u2 + \u03b51u3 + 3 \u03b8\u03081 = \u2212u3 (33) The dimensionless parameter \u03b5 is constant"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001247_0954406jmes321-Figure4-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001247_0954406jmes321-Figure4-1.png",
        "caption": "Fig. 4 Fillet surface, contact lines, and working surface of face gear",
        "texts": [
            " If \u03c91rbs\u03b8ks is omitted, equations (13), (14), and (8) yield the following expression for the equation of meshing f1(\u03b8ks, \u03c8 , \u03c61) = prbs(m21 cos \u03b3m \u2212 1) cos2(\u03b8ks + \u03b8os \u00b1 \u03c8) + p sin(\u03b8ks + \u03b8os \u00b1 \u03c8)(Em21 cos \u03b3m \u00d7 cos \u03c61 + rbs(m21 cos \u03b3m \u2212 1) sin(\u03b8ks + \u03b8os \u00b1 \u03c8)) + m21p cos(\u03b8ks + \u03b8os \u00b1 \u03c8)(p\u03c8 cos \u03c61 sin \u03b3m + E cos \u03b3m sin \u03c61) + m21 sin \u03b3m(Erbs \u2212 p2\u03c8 sin(\u03b8ks + \u03b8os \u00b1 \u03c8) sin \u03c61 + r2 bs(sin(\u03b8ks + \u03b8os \u00b1 \u03c8 + \u03c61) \u2212 \u03b8ks cos(\u03b8ks + \u03b8os \u00b1 \u03c8 + \u03c61))) (15) The face-gear position vector of the generated tooth surface can then be obtained by solving the simultaneous equations, equations (10) and (15). Because a standard face gear is generated by a shaper, the pinion corresponding to the shaper is conjugate to and in line contact with the face gear at every instant (Fig. 4). Thus, alignment errors will cause edge contact on the tooth surface and transmission errors of unfavourable shape and impermissible magnitude. Since this paper aims to design a face gear made by a moulding process \u2013 in which the mould surface is directly or indirectly machined by a computer numerical control (CNC) machine tool \u2013 there is great freedom in designing the tooth surface of the face gear. Thus, this paper proposes a methodology that modifies the tooth surface by superimposing a double crowning on the standard face gear both in the contact path and the instant contact line directions, as shown in the following steps"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002400_1.4006324-Figure20-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002400_1.4006324-Figure20-1.png",
        "caption": "Fig. 20 Deformation of the left web gears",
        "texts": [],
        "surrounding_texts": [
            "It has been reported in previous research that it is necessary to calculate the joint stress of the rim and the web when performing strength calculations of thin-rimmed spur gears [14]. This is because it is possible for the joint circle of the rim and the web to experience fatigue failure faster than the tooth root when the rim and the web are very thin because there are significant bending stress concentrations at the joint circle of the rim and the web. It is also well known that the filet radius of the joint circle of the rim and the web has a greater effect on the joint stress of the thin-rimmed gears. Because this effect has been investigated and reported in previous research [14], it is not discussed again in this paper. Only the effects of the web angle and the web thickness on the joint stress are discussed in this section. For a thin-rimmed gear design, it is possible that the filet radius of the joining part of the rim and the web can be varied with the change in the web angle so that the stress concentration of the filet is decreased. However, in this paper, if the filet radius is varied according to the change in the web angle, when investigating the effect of the web angle on the joint stress, it can become difficult to understand the effect of the web angle on the joint stress because the effect of the filet radius on the joint stress is also included in the calculation results. Therefore, a smaller filet radius of R\u00bc 1 mm is fixed for all of the gears with different web angles and web thicknesses as well as different web positions for all of the calculations in this paper. This approach will avoid including the effect of the filet radius in the results and will allow us to accurately determine the effects of the web angle on the joint stress. 6.1 Effect of the Web Angle on the Joint Stress. In this section, the effect of the web angle on the joint stress is investigated for thin-rimmed inclined web gears. The joint equivalent stresses are analyzed for all three types of gears shown in Figs. 2(a)\u20132(c) with different web angles. The filet radius at the joining part of the web and the rim is R\u00bc 1 mm for all of the calculations. Figures 14\u201316 show the calculation results of the left, center, and right web gears, respectively, with different web angles. The Journal of Mechanical Design MAY 2012, Vol. 134 / 051001-7 Downloaded From: http://mechanicaldesign.asmedigitalcollection.asme.org/ on 01/28/2016 Terms of Use: http://www.asme.org/about-asme/terms-of-use 051001-8 / Vol. 134, MAY 2012 Transactions of the ASME Downloaded From: http://mechanicaldesign.asmedigitalcollection.asme.org/ on 01/28/2016 Terms of Use: http://www.asme.org/about-asme/terms-of-use abscissas of Figs. 14\u201316 represent the circumferential angle A as shown in Fig. 3(a), and the ordinates of these figures are the equivalent stresses distributed along the joint circle of the rim and the web. From Figs. 14\u201316, it can be seen that the positions of the maximum joint stresses are changed from approximately 92 deg to approximately 100 deg when the web angle is increased. The maximum stresses are determined, and the relationship between the web angle and the maximum stress is shown in Fig. 17. From Fig. 17, it is apparent that the maximum joint stress is increased with an increasing increment in the web angle. 6.2 Effect of the Web Thickness on the Joint Stress. The effect of the web thickness on the joint stress is also investigated by using thin-rimmed center web gears with 30-deg web angles. In this investigation, the filet radius is also R\u00bc 1 mm. Figure 18 shows the equivalent stresses of the gears distributed along the joint circle of the rim and the web. The maximum stresses are determined, and the relationship between the web thickness and the maximum stress is shown in Fig. 19. Figure 19 shows that the maximum joint stress decreases with an increasing increment in the web thickness."
        ]
    },
    {
        "image_filename": "designv11_12_0002129_s0263574709005426-Figure14-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002129_s0263574709005426-Figure14-1.png",
        "caption": "Fig. 14. Virtual environment to calculate the weld trajectories.",
        "texts": [
            "12 This manipulator is composed of one prismatic joint, which comprises a mobile platform that moves on a flexible rail and six revolute joint (including a spherical wrist). Figure 12 depicts the Roboturb manipulator. The Roboturb manipulator carries out the recovery through an automated welding process, where the material is deposited in parallel chords in interposed layers.13 The trajectories have to follow the recovery welding process specifications14 and an eroded model was constructed to represent the real eroded surface. Figure 13 depicts the real environment and Fig. 14 the simulated environment. To apply the proposed differential kinematics method, the closed chains are firstly defined. To impose a desired trajectory, a 3P3R Assur virtual chain is attached to the end-effector with the base attached to the trajectory reference frame, this closed chain defines circuit 1. To solve the redundancy, a classical proposed solution is to define the prismatic velocity through the projection of the endeffector velocity.12 Adding a second 3P3R Assur virtual chain attached to the mobile platform with the base attached to the reference frame, a velocity can be imposed activating some of the prismatic virtual joints to specify the mobile platform (or prismatic) velocity, and this closed chain defines circuit 2"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003119_j.enzmictec.2012.10.002-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003119_j.enzmictec.2012.10.002-Figure1-1.png",
        "caption": "Fig. 1. Schematic diagram of the analytical activated",
        "texts": [
            " For the experiments, samples were collected from a tower located within the last phase of the landfill in the winter of 2006. These samples had COD and BOD5 concentrations of 12900 and 6300 mg L\u22121 respectively. The measured pH was 7.74, the Eh \u221231.7 mV and the conductivity 33.3 mS/cm. These samples were diluted with distilled water to adjust the leachate strength. 2.2. Microbial fuel cells: standard analytical (25 mL) type and column (1 L) setup The standard size MFCs were assembled and setup as previously described [41] (Fig. 1). Both anode and cathode were made of sheet of carbon fibre veil (PRF Composite Materials Poole, Dorset, UK) of 180 cm2 surface area \u2018folded down\u2019 to 5 cm2. The MFC columns formed a triplicate and were setup as previously described in Greenman et al. [35], but with the exclusion of the oxygen column and the open 3 icrobial Technology 52 (2013) 32\u2013 37 c M ( a e b M s a 2 c E p c m p 2 m p c T P a fi m i o d s 2 2 2 o t C p t r b r t f b 2 i e l m i w p e 2 t a s a t m A 2 C a 1 fl a 4 I. Ieropoulos et al. / Enzyme and M ircuit column (see Fig. 1). Anode consisted of carbon veil electrode (PRF Composite aterials Poole, Dorset, UK) of 360 cm2 surface area, folded down to a suitable size 12-folds of 10 cm \u00d7 3 cm). Cathode was also made of carbon veil with a surface rea of 360 cm2-folded down to a 3D structure of 3 cm \u00d7 5 cm \u00d7 1 cm. During the xperiments the three columns were also fluidically connected in series as described y G\u00e1lvez et al. [36]. The same inoculum, used for the standard analytical (25 mL) FCs, was also used for the column MFCs, and the data that will be presented on ulphate-reduction inhibition (see Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003501_2041302510394742-Figure9-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003501_2041302510394742-Figure9-1.png",
        "caption": "Fig. 9 The forces generated by pairings",
        "texts": [
            " kt1 and kt2 are the rotating coefficients of the two propellers. t is the airflow effect by the front counter-rotating propellers, fuselage, pairings, and mode transition, and the effect is bounded (i.e. t | \u0304t ). The tail counter-rotating propellers generate two reverse moments Mt1 and Mt2, and Mt1 = kc1\u03c9 2 t1, Mt2 = kc2\u03c9 2 t2 (6) where kc1 and kc2 are, respectively, two coefficients of reverse moments of the two tail propellers. 3.3 Forces through the four groups of pairings The forces generated by the four groups of pairings are shown in Fig. 9. In Fig. 9, O1, O2, O3, and O4 are the operating centres of the forces generated by the four groups of pairings, respectively. K is the point of intersection of axes O1O2 and O3O4. f1,1, f1,2, f2,1, f2,2, f3,1, f3,2, f4,1, and f4,2 are the forces projecting on axes xb, yb, and zb, respectively, generated by the tail propeller with the four groups of pairings (see Figs 9 and 10). \u03b41 and \u03b42 are the deflection angles at which the pairings 7 and 8 (Fig. 10) rotate with respect to axis BA, respectively; \u03b43 and \u03b44 are the swerving angle at which the pairings 5 and 6 rotate with respect to axis FE , respectively"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001181_j.wear.2007.04.005-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001181_j.wear.2007.04.005-Figure1-1.png",
        "caption": "Fig. 1. Geometry of contact (a) and geometry of t",
        "texts": [
            " In this case the ploughing friction is dominating. For soft aterials, the difference between the apparent friction and the loughing friction is due to the adhesive friction [11]. Recently elastic recovery was taken into account in the loughing part of the friction for conical indenters [12,13] and or a realistic conical tip with blunted spherical extremity [14]. he true area of contact is the sum of a front area and a rear area ue to the elastic unloading which partially recovers the rear ontact and contributes to decrease the friction. Fig. 1a depicts he geometry of the contact with the elastic recovery which is efined by the rear angle \u03c9. The rear angle is linked to the conact radius a and the rear contact radius ar by ar = a sin \u03c9. The eometry of the real tip is shown in Fig. 1b. Lafaye et al. [13,15] ave shown that for conical tips the cross section of the groove s not a triangular section having a width at the base equal to the ear width of the contact. The cross section is in fact generated y the intersection between a cone and a plane parallel to the xis of the cone which corresponds to a hyperbola. The relationhip of the ploughing friction with the triangular approximation s written as [13]: plough = 2 \u03c0 cot \u03b8 ( \u03c0 cos \u03c9(1 \u2212 sin \u03c9) \u03c0 + 2\u03c9 + sin 2\u03c9 ) (3) The evaluation of the ploughing friction is necessary to undertand the fundamental mechanism of friction which occurs in a cratch test",
            " True solution for a spherical tip For a perfectly spherical tip of radius R, the evaluation of the loughing friction with elastic recovery given by Lafaye et al. 14] does not make any approximation. The ploughing friction s given for a contact radius a. The equation which allows to valuate the ploughing friction coefficient is plough = 2 a2 \u03c12 sin\u22121(a cos \u03c9/\u03c1) \u2212 a cos \u03c9 \u221a \u03c12 \u2212 a2 cos2 \u03c9 \u03c0 + 2\u03c9 + sin 2\u03c9 (5) here = \u221a R2 \u2212 (a cos \u03c9)2 tan2 \u03c9 . True solution for a real tip The real tip is modelled by a conical tip with a blunted sphercal extremity. The tip is split into two parts: the spherical and he conical part (see Fig. 1b). We define the limiting contact adius a0 = R cos \u03b8, which corresponds to the boundary between he conical and spherical parts of the tip. Two cases must be disinguished according to the value of the contact radius and the ear angle. When ar \u2265 a0 (see Fig. 2a), the cross section is given y the solution of a perfectly conical tip, then the ploughing fricion is given by Eq. (4). Conversely when ar \u2264 a0 (see Fig. 2b), he cross section is the sum of the cross section of a spherical tip 552 S. Lafaye / Wear 264 (2008) 550\u2013554 wher o o t T e \u03bc a2 0 co n 2\u03c9 w \u03c1 a f 5 f w g c p g \u03c9 W e t t t m s t m f (4) the ploughing friction is written as: \u03bcplough = 2/\u03c0 cot\u03b8f(\u03c9)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000243_jae-2007-772-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000243_jae-2007-772-Figure1-1.png",
        "caption": "Fig. 1. Equivalent Reluctance Model: (a) For non-conducting elements. (b) For conducting elements.",
        "texts": [
            " All rights reserved In this model, the scalar potential is used in MEC modelling while the analytical solution employs the magnetic vector potential. To analyze the electromagnetic devices with magnetic equivalent circuit method, the related equations are as follows [3]: \u2207 \u00b7 B = 0, \u2207\u00d7 B = \u00b5J , \u2207\u00d7 J = \u2212\u03c3 dB dt = \u2212\u03c3 ( \u2202B \u2202t + \u03c9r \u2202B \u2202\u03b8 ) (1) where, B, J , \u03c9r, \u03c3, and \u00b5 are magnetic flux density, current density, rotational velocity of rotor, electrical conductivity, and magnetic permeability respectively. For non-conducting elements, an equivalent reluctance model shown in Fig. 1a is used. Also, for conducting elements two additional MMF sources in radial orientation are used as shown in Fig. 1b. In cylindrical coordinates, Eq. (1) can be rewritten as in the following: \u2202Jz \u2202r = \u03c3 [ \u2202B\u03b8 \u2202t + \u03c9r \u2202B\u03b8 \u2202\u03b8 ] (2) where, \u03c3 is the conductivity of rotor bars. The first term of the above equation denotes the induced current due to time variation caused by motor input voltage frequency and the second term denotes the induced current due to the relative motion between rotor and stator. The effect of relative motion between rotor and stator can be modeled by multiplying the conductivity by the slip of the rotor"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000427_j.ijsolstr.2007.12.004-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000427_j.ijsolstr.2007.12.004-Figure1-1.png",
        "caption": "Fig. 1. Deformation pattern 1. Elevated middle point.",
        "texts": [
            " Their work was motivated by the mechanics problem observed in the stent deployment procedure. In this paper, we extend the analysis in Chen and Li (2007) by studying the deformation patterns of a planar elastica under a conservative pushing force within a circular channel. At one end of the circular channel the elastica is fully clamped. At the other end the lateral displacement and slope of the elastica are fixed, while the elastica is allowed to slide in and out of the clamp in the longitudinal direction. Fig. 1 shows a circular channel with outer and inner radii ro and ri. w is the span angle of the circular channel. An elastic strip, or the planar elastica, is placed inside the channel. One end of the elastica is fully clamped at the left end B of the channel. On the right end A the elastica is partially clamped. By partially clamped we mean that the strip is allowed to slide freely through the clamp, while the lateral displacement and slope at A are fixed. The strip will be under external pushing force at end A. As a result of this pushing force, the strip originally outside the channel will slide through the partial clamp, and the shape of the strip inside the channel will change accordingly. It is noted that in Fig. 1 we only show the elastica inside the channel. We assume that the elastic strip is inextensible and is stress-free when it is straight. The strip is uniform in all mechanical properties along its length. It is assumed that the thickness of the elastic strip is much smaller than the clearance ro ri of the channel. Before the application of the external pushing force, the elastica is in the form of a circular shape with constant bending moment throughout the entire length, as shown by the circular dashed curve in Fig. 1. The original length of the elastica inside the circular channel is denoted as lo. The solid curve in Fig. 1 represents the first stage of the elastica deformation when an external force FA is applied to push in the elastica at end A. The total length of the strip inside the channel is increased by Dl. In Fig. 1 the elastica is not in contact with the circular walls yet. It is not hard to envision that when the pushing force increases, the elastica will contact the outer radius of the channel first. However, it is not obvious how the elastica will behave thereafter. The purpose of this paper is to study the behavior of the elastica in response to the external pushing force. The easiest way to visualize the elastica deformation is to build an experimental apparatus as described in Section 2 and make observations",
            " It is noted that in the experiment we control the length of the strip being pushed in the channel. Therefore, this is a displacement control procedure instead of a load control procedure. Figs. 3\u20135 show the photographs of the elastica deformations we observed in the laboratory after fixing the left clamp on the rail. Fig. 3 shows 12 deformation patterns in one of the experiments. In Fig. 3(1) the elastica is pushed in a length Dl = 2 cm. The middle point of the elastica is lifted up a small distance. This deformation pattern has been demonstrated in Fig. 1. After the strip is pushed in 4 cm, the elastica makes a point contact with the outer wall, as shown in Fig. 3(2). When Dl increases to 7 cm, the point contact at the middle point evolves to distributed contact in the middle region, as shown in Fig. 3(3). After fixing the end of the strip on the right and poking the strip in the middle by hand, the elastica deformation jumps to Fig. 3(4). In Fig. 3(4) the middle point of the elastica floats in the air while the neighboring segments remain in contact with the outer wall",
            " There are six different types of sub-domains which can be observed when a clamped\u2013clamped elastic strip is placed inside a circular channel with clearance. The difference between these six sub-domains lies in their boundary conditions or the contact conditions, as listed in Table 1. Similar sub-domain analysis was first proposed by Domokos et al. (1997) for a hinged\u2013hinged bar constrained in straight walls. In the following we discuss how these boundary conditions or contact conditions affect the equilibrium equations. Clamped\u2013clamped sub-domain (listed as case (a) in Table 1) occurs only when the whole elastica is a subdomain, as shown in Fig. 1. The strip is in the form of a circular shape before the external longitudinal force FA is applied. When FA increases from zero, the middle point of the strip will be elevated. The deformation is symmetric with respect to the central radius, as observed experimentally in Fig. 3(1). First of all we establish an xy-coordinate system with the origin at point A. The moment equation at any point (x,y) of the deformed strip can be written as Table Six ele Case (a) Cla (b) Cla (c) Cla (d) Co (e) Co (f) Dis EI dh ds \u00bc QA sin w 2 F A cos w 2 y \u00fe QA cos w 2 \u00fe F A sin w 2 x\u00feMA \u00f01\u00de QA and MA are the shear force and bending moment provided by the partial clamp at A"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001196_0094-114x(72)90004-3-Figure13-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001196_0094-114x(72)90004-3-Figure13-1.png",
        "caption": "Figure 13. The plane RP*P*R* mechanism.",
        "texts": [
            " To show the general form of physical realizations of these linkages, Figs. 9-12 show models of linkages of the first four mechanism groups. The fifth group is omitted since it only contains the plane four-bar linkage. In addition to the plane four-revolute mechanism, there are three more plane four-link mechanisms in the table. The RP*R*P* mechanism produced by equation (45) is a plane mechanism as previously noted and shown in Fig. 6. The RP*P*R* mechanism of group 2 is also a plane mechanism with two revolute joints and two prisms or sliders. Figure 13 illustrates this mechanism. And finally, the RRR*P* mechanism of group 4 is the familiar crank-slider linkage. It has been brought to the attention of the author that a similar but not identical linkage summary was simultaneously obtained[10]. Whereas the present study is an extension of the German literature[6-8] the similar study is an extension of the Russian literature[5, 9, 14]. Some significant differences are that this grouping is restricted to linkages with cylindric, revolute and prismatic pairs, that it identifies the types of added constraints in the joints and that it includes closed-form relations for the dimensions of the RR*CC linkage of equation (41) (Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002028_icnsc.2010.5461545-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002028_icnsc.2010.5461545-Figure1-1.png",
        "caption": "Fig. 1. Simplified UH-60A gear transmission train",
        "texts": [
            " The investigation results showed that both vibration and AE sensors were capable of detecting the gear fault in a STG. However, in terms of locating the source of the fault, AE sensors outperformed vibration sensors. I. INTRODUCTION HE The requirement for lower weight in helicopters has lead to the development of the split torque gearbox (STG) to replace the traditional planetary gearbox by the drive drain designer [1]. A simplified planetary gear transmission system described in [2] is shown in Fig.1. This work was supported in part by Goodrich Corporation under contract No. 75095. D. He is with the Department of Mechanical and Industrial Engineering, University of Illinois at Chicago, Chicago, IL, 60607, USA (phone: 312- 996-3410 email: Davidhe@uic.edu) R. Li is with the Department of Mechanical and Industrial Engineering, University of Illinois at Chicago, Chicago, IL, 60607, USA (email: rli8@uic.edu). E. Bechhoefer is with Goodrich Sensors & Integrated Systems, Vergennes, VT, 05419, USA (email: Eric.Bechhoefer@goodrich.com) As shown in Fig. 1, a central sun gear is surrounded by two or more rotating planets in the planetary gear system. The torque is transmitted from the central sun gear through the planets gears to the planet carrier and from the planet carrier to the main rotor shaft. In comparison with traditional planetary gearbox, STG potentially offers the following benefits [3]: (1) high ratio of speed reduction at final stage; (2) reduced number of speed reduction stages; (3) lower energy losses; (4) increased reliability of the separate drive paths; (5) fewer gears and bearings; (6) lower noise"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003817_chicc.2015.7260315-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003817_chicc.2015.7260315-Figure1-1.png",
        "caption": "Fig. 1: Quadrotor and coordinate definition",
        "texts": [
            " Since the settling time of the closed loop system does not depend on the initial state, but only on the controller parameters, the outer and inner loop settling time of the closed-loop system can be assigned offline, which explicitly guarantees the two-time scale separation between the outer and inner loop. In addition, the introduction of sliding mode makes the system robust to the external disturbances and model uncertainty. Finally, simulation results are provided to validate the effectiveness of this algorithm. 2 Problem Description 2.1 Coordinates E(OXY Z) denotes an earth-fixed inertial frame and B(oxyz) a body-fixed frame whose origin o is at the centre of mass of the quadrotor , shown in Fig.1. \u03be = [x, y, z]T represents the absolute position and \u0398 = [\u03c6, \u03b8, \u03c8]T is the attitude in terms of Euler angles. The orthogonal rotation matrix to orient the quadrotor R This work is supported by National Natural Science Foundation (NNSF) of China under Grant 61203022, and Aeronautic Science Foundation of China under Grant 2012CZ51029. is as follows: R = \u239b \u239dc\u03b8c\u03c8 s\u03b8c\u03c8s\u03c6 \u2212 s\u03c8c\u03c6 s\u03b8c\u03c8c\u03c6 + s\u03c8s\u03c6 c\u03b8s\u03c8 s\u03b8s\u03c8s\u03c6 + c\u03c8c\u03c6 s\u03b8s\u03c8c\u03c6 \u2212 c\u03c8s\u03c6 \u2212s\u03b8 c\u03b8s\u03c6 c\u03b8c\u03c6 \u239e \u23a0 (1) 2.2 Mathematical Model In order to simplify the controller designing, two reasonable assumptions are made: Assumption 1: Quadrotor is a rigid body"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000610_1.2647443-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000610_1.2647443-Figure2-1.png",
        "caption": "Fig. 2 View of steady circular whirl in an axial plane",
        "texts": [
            " 5 for M , there results in constraining as a coupled motion, hus 76 / Vol. 129, APRIL 2007 om: http://tribology.asmedigitalcollection.asme.org/ on 01/28/2016 Terms \u00b1 j 1 + 2 = 0 or 1 j 2 = 0 6 This is a description of circular forward- and backward-rotating whirl orbits with an undetermined orbit amplitude and phase reference. One can specify 1 1 to define the standard eigenvectors with the dummy index representing either f or b to connote forward or backward whirling f = 1 \u2212 j and b = 1 j 7 Illustrated in Fig. 2 is a cross-sectional view in an axial plane of the bearing surfaces that contain an eccentric film thickness. The eccentric film profile is invariant to the line of centers that is undergoing a steady whirl motion. The right-hand side of Eq. 1 can now be restated as separate 1-DOF homogeneous equations, respectively, for axial, cylindrical, and conical modes ms2 + Zaxial z = 0 ms2 + Zcylindrical, x cylindrical, = 0 ITs2 \u00b1 jIP s + Zconical, x conical, = 0 8 The perturbed rotor motion is depicted in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001283_6.2009-5794-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001283_6.2009-5794-Figure2-1.png",
        "caption": "Figure 2. System\u2019s control surfaces and axes definition",
        "texts": [
            " Finally, the control design methodology is discussed and is followed by simulation and flight test results. 2 of 9 American Institute of Aeronautics and Astronautics The MiniAV, presented in figure 1, is a commercial-off-the-shelf fixed-wing airplane (Flatana, Great Planes) which has a 0.91 m wing span. The test vehicle has an all-up-weight of 450 g. The propulsion system consists of a RIMFIRE 22M-1000 brushless motor coupled to a 25 A electronic speed controller and a 10x3.8\" propeller. The aerodynamic surfaces are controlled with a standard 72 MHz receiver and micro servos. As shown in figure 2, there are four manipulated variables: ailerons, elevator, rudder and throttle. In hovering mode, these manipulated variables play roles very similar in comparison to level flight. Ailerons are still used for roll rate control. However, their deflection generate strong air stream perturbations on the elevator and the rudder. These perturbations strongly affect the control authority of these actuators. Elevator and rudder are respectively used for pitch and yaw rate control. Finally, the throttle is used for altitude control",
            " It calculates and transmits orientation under cosine matrix or Euler representations at 100 Hz. Its dynamic angle estimation accuracy is \u00b12\u25e6 and the fusion filter used is not documented. Quaternion representation,10,11 which is gaining popularity for expressing a system attitude without singularity, is used for calculating the attitude angles which are converted back to Euler angles for control purpose. To maintain the airplane in hovering mode, angle rates (p, q and r) and attitude angles (\u03c6, \u03b8 and \u03c8) were selected as variables of interest for controlling the MiniAV attitude (figure 2). The selected attitude angles are very similar to vertical Euler angles, however the algorithm proposed by Green6 is used to generate \u03b8 and \u03c8 decoupled from \u03c6 in near hover conditions. Green\u2019s algorithm is normally used at the beginning of a transition between level flight and hovering mode and has the objective of generating a quaternion with no component in X. When converting the resulting quaternion to Euler angles, \u03c6 \u223c= 0 and thus \u03b8 and \u03c8 are nearly decoupled from it. To control \u03c6, the vertical Euler \u03c6 is directly taken"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003501_2041302510394742-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003501_2041302510394742-Figure2-1.png",
        "caption": "Fig. 2 Tail-sitter aircraft design",
        "texts": [
            " The concept demonstrator is a \u2018tail-sitter\u2019 configuration and utilizes co-axial, counter-rotating rotors. A relatively large cruciform tail provides directional control in the airplane mode as well as serving as landing gear in the helicopter mode. Transition between hovering flight and wing-borne flight will be the main focus of the initial flight test series. However, in references [13] to [22] the flying height cannot be kept invariant while the aircraft undergoes a vertical-to-horizontal transition. The aircrafts climbs upward during transition from hover to forward flight (see Fig. 2). Flexible manoeuvrability must Proc. IMechE Vol. 225 Part G: J. Aerospace Engineering at The University of Manchester Library on April 23, 2015pig.sagepub.comDownloaded from Xinhua Wang and Hai Lin be designed for the aircraft, and the weight of the aircraft is strictly limited. More importantly, the sufficiently big thrust force must be provided to finish mode transition. In this article, a novel rotor-fixed wing hybrid aircraft with two free wings is presented. The two free wings are extended from two fixed-wing root portions (which are rigidly and non-rotatably attached to the fuselage for rotation with the fuselage relative to the spanwise axis), respectively, in a manner such that the wings are pivotable about their spanwise axis"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003887_j.procir.2015.06.103-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003887_j.procir.2015.06.103-Figure1-1.png",
        "caption": "Fig. 1. Kinematic diagram of the one-stage nutation drive",
        "texts": [
            " Peer-review under responsibility of the scientific committee of the CIRP 25th Design Conference Innovative Product Creation. Keywords: Screw theory; Spiral bevel gear; Nutation drive; Kinematic modeling 1. Introduction Nutation motion is a kind of transmission proposed on the basis of the motion principle of celestial planet or gyroscopic. The principle of this rotation is used to create a kind of mechanical drive called the nutation drive. The kinematic diagram of one-stage nutation drive system shown in Fig. 1 is composed of an input shaft 3, external spiral bevel gear 1, fixed internal spiral bevel gear 2, equiangular speed ratio mechanism 4 and an output shaft 5. The external gear 1 meshing with the internal gear 2 is free to rotate about its axis, which is inclined to the axis of the input shaft. As a kind of transmission form, the nutation drive has a broad development spaces in the field of robotics wrist and aerospace craft with the advantages of low noise, higher carrying capacity, steady transmission ability, higher transmission ratio and small volume",
            " And 0123 aS 21 21 12 SS SS a (6) The pitch of the sum screw can be obtained by the dot product between the 3S and equation (4) and expressed as cos2 sin)(coscos 21 2 2 2 1 2112212 2 22211 2 11 3 SSSS SSaaSShShSShShh (7) where represents the angle between 1$ and 2$ . The cross product of between the 3S and equation (4) can be given as below cos2 cos)(sin)( 21 2 2 2 1 2 22 2 1121122112 3 SSSS SaSaSSaaSShha (8) Then the magnitude, location, pitch and direction of the sum screw can be determined by equations (3), (7) and (8). Particularly, in the nutation drive, the cone vertex of the bevel gear pair is coincidence, and consequently the point A , B and C is coincidence, i. e. 01a and 012 aa . Compare with Fig. 1 and Fig. 2, the screw system of nutation drive are further illustrated in Fig. 3. In nutation gear drive system, bevel gear pair meshes only for pure rotation, and the pitch of the screw is 021 hh . Thus, the pitch of the screw and the distance between sum screw 3$ and coordinate system ),,( ZYXS is 03h and 03a . The meshing between the external and internal spiral bevel gears in the nutation drive can be considered as the external and internal spiral bevel gear meshing with crown gear (an imaginary gear), which has a pitch cone angle of 90 , and the pitch cone is at right angle to its axis"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001069_1.3070580-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001069_1.3070580-Figure2-1.png",
        "caption": "Fig. 2 Schematic diagram of the adjustable pad with various configurations and angle definitions",
        "texts": [
            "org/ on 01/29/2016 Terms coordinate i.e., circumferential direction , is taken from the position of maximum film thickness. The present analysis requires that the position of maximum film thickness be found beforehand. This is done by assuming an arbitrary value of attitude angle . After each calculation, the attitude angle computed from Eq. 11 is compared with the assumed value of the attitude angle . The value of is modified with a small increment, and Eq. 2 is solved using this modified value until is equal to 0. Figure 2 shows the schematic diagram of various angles and adjustments considered in the present analysis. 8 Validation Externally adjustable fluid film bearing will perform as a conventional partial arc bearing when Radj and Tadj are set to zero. Steady state performance curves in Figs. 3 and 4 compare the results of the present analysis with the results illustrated in Szeri 15 for a centrally loaded 160 deg fixed pad partial bearing with L /D=1.0, operating with a Reynolds number 5000 and 10,000. This comparison validates the implementation of the turbulence model"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003878_1.4031025-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003878_1.4031025-Figure2-1.png",
        "caption": "Fig. 2 Bevel gear flank sketch",
        "texts": [
            " In coordinate system x00y00z00, unit tangent vector to arc OP \u00f0 at point P is n0 \u00bc 0 1 0\u00bd T (superscript T means transpose); therefore, n in x00y00z00 is nx00y00z00 \u00bc sin# cos# 0\u00bd T (17) Hence, unit normal vector n in xyz coordinate system is nxyz \u00bc cos h sin h 0 sin h cos h 0 0 0 1 2 4 3 5 1 0 0 0 cos g sin g 0 sin g cos g 2 4 3 5 sin# cos# 0 2 4 3 5 (18) or, after simplification nxyz \u00bc cos hsin# sin h cos g cos# sin hsin# cos h cos g cos# sin g cos# 2 64 3 75 (19) From Eqs. (16) and (19), coordinates and unit normal vectors of any point P on spherical involute surface are calculated based on two independent surface parameters of r and u. Points of spherical involute surface lie between gear base cone with cone angle db \u00bc 2b and face cone with cone angle df between inner cone ri and outer cone ro, all shown in Fig. 2. For any given point P shown in Fig. 2 located at \u00f0Rp; Lp\u00de, coordinates are calculated through solution to below set of equations. From this, normal to the surface at this point can also be calculated x2 \u00fe y2 \u00bc R2 p; z \u00bc Lp (20) Any point P located at \u00f0Rp;Lp\u00de on the bevel gear flank has a unique set of associated \u00f0r;u\u00de values that relate them by Eqs. (13)\u2013(16); for each point P also there exists a unique unit normal n. Actual surfaces should be measured against its intended theoretical surface. The theoretical surface can be spherical involute, any modifications to it, or other bevel gear surface profiles used Journal of Mechanical Design SEPTEMBER 2015, Vol"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001369_j.jsv.2008.09.045-Figure6-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001369_j.jsv.2008.09.045-Figure6-1.png",
        "caption": "Fig. 6. Link\u2013pulley contact kinematics.",
        "texts": [
            " The slip angle, g, defines the plane where the friction force between the chain link and the pulley acts (i.e. g defines the slip direction). It is the angle which the resultant friction force vector, Ff, makes with a unit tangential direction vector to the pulley. So, in order to get the friction force vector, it is crucial to keep track of the relative velocity vector between the chain link and the pulley. The relative acceleration and the relative velocity between the link and the pulley can be obtained using the contact kinematics depicted in Fig. 6. Almost all models, except a few mentioned in the literature, use classical Coulomb\u2013Amonton friction law to model friction between the contacting surfaces of a CVT. The friction phenomenon described by this law is inherently discontinuous in nature. It is common engineering practice to introduce a smoothening function to represent this set-valued friction law. However, certain friction-related phenomena like chaos, limit-cycles, hysteresis, etc., are neither easy to detect nor easy to explain on the basis of classical Coulomb\u2013Amonton friction theory"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003142_1.3616922-Figure18-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003142_1.3616922-Figure18-1.png",
        "caption": "Fig. 18 Actual shape of tooth contact as determined by etching tests (Concave side of pinion tooth.)",
        "texts": [
            "org/about-asme/terms-of-use where mN \u2022\u2022 Vn ' mo m F m\u201e / = load sharing ratio\u2014ratio of load carried on line of eontact VW to total load V W - 4P mean normal base pitch y / m F 2 + m,p2 modified contact ratio (contact ratio with- in boundaries of ellipse of tooth bearing in zone of action) face contact ratio profile contact ratio distance from center of surface of action, 0, to line of contact, VW, measured in the normal direction = VI3+ \u00a3 V W ~ 4KNVN(K\u201ePN + 2/ ) ] 3 kn = l + E V W - \u00b1KPn(KVN - 2/)]3 4n = l K = Tj2 2 cos z (21 a) a positive integer, which takes on successive values from 1 to a or /3, generating all real terms in the series. Imaginary terms should be ignored. Position of Point of Load Application The true shape of the area of contact on a spiral bevel or hypoid gear tooth, as determined by etching tests [17], ia in the shape of a distorted ellipse, Fig. 18. This differs from the previous assumption that the area was a symmetrical ellipse. Note that the center of pressure is toward the heel (outer end of the tooth). Although there is a distributed load, for simplicity a point is chosen to represent a concentrated load replacing the distributed load. Based on the etching tests and other similar studies by the present author it was determined that point K will lie along the line of contact at a distance j toward the heel of the fc tooth from the center of the line of contact 0'"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001345_robot.2008.4543604-Figure5-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001345_robot.2008.4543604-Figure5-1.png",
        "caption": "Figure 5. The cylinder driving",
        "texts": [
            " The set of all functions thats , can take on at any time is state function space . We will consider that .2L The control forces have the distributed components along the arm, , , that are determined by the lumped torques, tsF , tsFq , Ls ,0 N i tilstsF i 1 , (8) N i qq tilstsF i 1 , (9) where is Kronecker delta, , andllll N21 821 dSppt iii (10) 821 dSppt iii qqq , (11)Ni ,,2,1 In (10), (11), , , , represent the fluid pressure in the two chamber pairs, 1 i p 2 i p 1 iqp 2 iqp , and S, d are section area and diameter of the cylinder, respectively (Figure 5). The pressure control of the chambers is described by the equations [9] q ki k i ki u dt dp a , qki k qi ki u dt dp qb , (12)2,1k where kia , are determined by the fluid parameters and the geometry of the chambers and qbki 00kia (13) 00kib , ; ;2,1k Ni ,,2,1 q, (14) IV. CONTROL PROBLEM The control problem of a grasping function by coiling is constituted from two subproblems: the position control of the arm around the object-load and the force control of grasping. A. Position Control We consider that the initial state of the system is given by Tqs 000 ,,0 (15) Ts 0,0,00 (16) where s,00 , , (17)sqq ,00 Ls ,0 corresponding to the initial position of the arm defined by the curve 0C sqsC 000 ,: , Ls ,0 (18) The desired point in is represent by a desired position of the arm, the curve that coils the load,dC T ddd q, (19) T d 0,0 (20) sqsC ddd ,: , Ls ,0 (21) In a grasping function by coiling, only the last mn 1 elements Nm are used"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003484_mabi.201100060-Figure12-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003484_mabi.201100060-Figure12-1.png",
        "caption": "Figure 12. (a) Fluorescence microscopy image of fibrous AM formed under shear stress and (b) Schematic illustration of the device used to form fibrous AM under application of shear stress.",
        "texts": [
            " Since MTs grow at a faster rate at the plus end than at the minusend, theplusendgrewpreferentially, andthegrowth of the minus end was suppressed in the absence of excess free tubulin. Consequently, MTs were formed with their plus ends oriented towards the cold end. Thus, the correlation among asymmetric polymerization conditions and factors governingpolarity, anisotropy, etc. of suchwelloriented structures is now understandable. An actin and myosin solution self assembled into an oriented actomyosin (AM) gel, as shown in Figure 12a, on application of shear stress followed by subsequent crosslinking, which subsequently shrank in the presence of ATP.[76] This shrinkage dramatically enhanced the regular orientation of actin and myosin filaments in the gel and resembled the in vivo system which also shows a similar effect on applied stress. Hence the application of shear stress also appears as a novel method to control the orientation of protein components within the gel. The system used for applying the shear stress is illustrated schematically in Figure 12b, and Figure 13 shows the schematic shrinking and elongation process of an AM gel. s captured under ectively; (c) schelymerization was Combination of \u2018Top Down\u2019 and \u2018Bottom Up\u2019 Approaches A temporal control on the self-assembly process of a motor protein system has been achieved through a novel in situ method, by combining a \u2018top down\u2019 and \u2018bottomup\u2019approach. For this, a synthetic photoresponsive polycation has been developed, which provided a means for spatiotemporal control over polymer/ actin complex formation"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003571_s0263574711000324-Figure3-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003571_s0263574711000324-Figure3-1.png",
        "caption": "Fig. 3. Planar geometry of the mechanism in the vertical plane.",
        "texts": [
            " (5) The mathematical model of all m motors can be written in a vector form as u = R \u00b7 i + CE \u00b7 \u02d9\u0304\u03b8 CM \u00b7 i = I \u00b7 \u00a8\u0304\u03b8 +Bu \u00b7 \u02d9\u0304\u03b8 \u2212 S \u00b7 (\u2666m \u00b7 \u03b5 + \u03b5m) \u2223\u2223\u2223\u2223\u2223\u2223 \u2211 M=0 about the rotation axis of the each motors . (6) Example. Let us analyze the behavior of the robotic pair consisting of elastic gear and flexible link in the presence of the second mode, as depicted in Fig. 2. The link has two modes (the lower one and the upper one) and each of them is considered as a mode of rectangular cross-section ci,j xbi,j . The presence of the second mode is introduced into the analysis of the robotic pair behavior. The relations between the important angles are defined in Fig. 3. q = \u03b8\u0304 + \u03be + \u03d11,1, \u03b3 = \u03b8\u0304 + \u03be, \u03b4 = \u03d11,2 + \u03c91,1, \u03c91,j = \u03d11,j 2 . (7) The dynamic model (both the model of flexible line and model of the motion of each mode\u2019s tip) is defined according to classical principles but with the previously introduced new DH parameters, using Lagrange\u2019s equations. The following quantities are adopted: q, \u03b4, \u03b3 , and \u03b8\u0304 as generalized coordinates (see Fig. 3). Small bending angles should be taken into consideration and it should be adopted that lsi,j = li,j and if tan \u03d1i,j = ri,j li,j , then \u03d1i,j \u2248 tan \u03d1i,j and ri,j = li,j \u00b7 \u03d1i,j , i = 1, j = 1, 2 \u03d1i,j = ri,j li,j . (8) The magnitude ri,j is the maximal deflection, i.e., the deflection at the each mode\u2019s tip. The bending angles are expressed in Fig. 3 via generalized coordinates, Eqs. (7) and (8) r1,1 = l1,1 \u00b7 (q \u2212 \u03b3 ), r1,2 = l1,2 \u00b7 ( \u03b4 \u2212 q \u2212 \u03b3 2 ) , (9) m\u0302el1,1 = 33 140 \u00b7 w\u03041,1 \u00b7 (x1,1 \u2212 x\u03021,1), m\u0302el1,2 = 33 140 \u00b7 w\u03041,2 \u00b7 (x1,2 \u2212 x\u03021,2), (10) J\u0302elzz1,1 = m\u0302el1,1 \u00b7 ( x1,1 \u2212 x\u03021,1 2 )2 , J\u0302elzz1,2 = m\u0302el1,2 \u00b7 ( x1,2 \u2212 x\u03021,2 2 )2 . (11) Equation (10) sources from ref. [22]. Kinetic and potential energies of the mechanism presented in Fig. 2 are denoted as \u02c6\u0302Ekm and \u02c6\u0302Ep. All the angles in the expression for kinetic and potential energies characterizing flexibility of the links should also be expressed via generalized coordinates",
            " The magnitude M\u03021,1 includes the external force Fuk that across the Jacobi matrix J\u0302 maps on the direction of the first generalized coordinate. This is just the procedure for obtaining Euler\u2013Bernoulli equation by which the motion of any point on the flexible line of the first mode is performed [H\u03021,1 H\u03021,2 H\u03021,3 0] \u00b7 \u03c6\u0308 + h\u03021 + J\u03021,1 \u00b7 Fuk x + J\u03022,1 \u00b7 Fuk y \u2212 1 2 \u00b7 Cs1.2 \u00b7 l1,2 \u00b7 r1,2 \u2212 1 2 \u00b7 Bs1,2 \u00b7 l1,2 \u00b7 r\u03071,2 + \u03b21,1 \u00b7 \u22022(y\u03021,1 + \u03b71,1 \u00b7 \u02d9\u0302y1,1) \u2202 x\u03022 1,1 = 0. (19) Fuk (N) is the dynamic force of the contact (in this case). The component of the entire external force in the radial direction (see Fig. 3) is Fc = (me \u00b7 \u0308 + be \u00b7 \u0307 + \u00b7Fo c + ka1 \u00b7 ), whereas the friction force is Ff = \u2212\u03bc p\u0307s |p\u0307s | \u00b7 Fc, as in ref [23]. The friction coefficient is \u03bc. The velocity of the robot tip is p\u0307s . is the distance from the point \u201c0\u201d to the trajectory, marked with \u03bb on Fig. 2, and = l \u2212 , l = l1,1 + l1,2. me (kg) is the equivalent mass, be (N/(m/s)) is the equivalent damping, ka1 (N/m) is the equivalent rigidity. \u03c2 = C\u03be \u00b7 \u03be + B\u03be \u00b7 \u03be\u0307 is the elasticity moment of gear and \u03b5i,j = (Csi.j \u00b7 ri,j + Bs i,j \u00b7 r\u0307i,j ) \u00b7 li,j is the bending moment of each mode\u2019s tip motion \u03c6 = [q \u03b4 \u03b3 \u03b8\u0304]T, H\u03021,1 = m\u0302el1,1 \u00b7 (x1,1 \u2212 x\u03021,1)2 + (m + mel1,2) \u00b7 l2 1,1 + (m + mel1,2) \u00b7 l2 1,2 + 2 \u00b7 (m + mel1,2) \u00b7 l1,1 \u00b7 l1,2 \u00b7 cos \u03b4 + 9 4 \u00b7 J\u0302elzz1,1 + 9 16 \u00b7 (Jzz + Jelzz1,2), H\u03021,2 = ",
            "\u201d A geometric link between these characteristics (internal coordinates) and the space of Cartesian coordinates (external coordinates) has been defined by using the transformation matrix, or so-called \u201cdirect kinematics\u201d in the robotics. The rotation matrix that describes the change of position (Cartesian coordinates) and orientation (Euler angles) of the tip of every mode of segment has the form Te i\u22121 i = \u23a1 \u23a2\u23a2\u23a2\u23a3 cos \u03c1i,j \u2212 sin \u03c1i,j cos \u03b1i,j sin \u03c1i,j sin \u03b1i,j li,j \u00b7 cos \u03c1i,j sin \u03c1i,j cos \u03c1i,j cos \u03b1i,j \u2212 cos \u03c1i,j sin \u03b1i,j li,j \u00b7 sin \u03c1i,j 0 sin \u03b1i,j cos \u03b1i,j di,j 0 0 0 1 \u23a4 \u23a5\u23a5\u23a5\u23a6. (35) \u03c1i,j , li,j , \u03b1i,j , and di,j are the new DH parameters that also encompass the rigidity characteristics, see Fig. 3 and Eq. (7) (\u03b1i,j = 0o and di,j = 0 (m)): Te 0 1 = \u23a1 \u23a2\u23a2\u23a2\u23a2\u23a3 cos q \u2212 sin q 0 l1,1 \u00b7 cos q sin q cos q 0 l1,1 \u00b7 sin q 0 0 1 0 0 0 0 1 \u23a4 \u23a5\u23a5\u23a5\u23a5\u23a6 . (36) The matrix rotation (36) describes position change and orientation of the top of the first mode\u2019s tip of the link Te 1 2 = \u23a1 \u23a2\u23a2\u23a2\u23a2\u23a3 cos \u03b4 \u2212 sin \u03b4 0 l1,2 \u00b7 cos \u03b4 sin \u03b4 cos \u03b4 0 l1,2 \u00b7 sin \u03b4 0 0 1 0 0 0 0 1 \u23a4 \u23a5\u23a5\u23a5\u23a5\u23a6 . (37) The matrix rotation (37) describes position and orientation change of the top of the second mode\u2019s tip of the link. For thirty matrix rotations adopted l1,3 = 0: Te 2 3 = \u23a1 \u23a2\u23a2\u23a2\u23a2\u23a3 cos \u03c91,2 \u2212 sin \u03c91,2 0 0 sin \u03c91,2 cos \u03c91,2 0 0 0 0 1 0 0 0 0 1 \u23a4 \u23a5\u23a5\u23a5\u23a5\u23a6 ",
            " The overall transformation matrix describes the change of the position and orientation of segment tip in a coordinate frame x, y, z : Te 0 3 = Te 0 1 \u00b7 Te 1 2 \u00b7 Te 2 3. (39) The Jacobi matrix for a manipulator with elastic joints and links maps the velocity vector of external coordinates p\u0307s into the velocity vector of internal coordinates\u03c6\u0307 : \u03c6\u0307 = J\u22121(\u03c6) \u00b7 p\u0307s, (40) where p\u0307s = [x\u0307 y\u0307 z\u0307 \u03c8\u0307 \u2118\u0307 \u03d5\u0307]T defines the velocity of a given point of the robotic system in Cartesian coordinates, whereas \u03c6\u0307 = [\u03c1\u03071,1 \u03c1\u03071,2 \u03c1\u03071,3 \u03c1\u03071,4 . . . \u03c1\u03071,n]T defines the velocity vector of internal coordinates. In this example, see Fig. 3 and Eq. (40) have the form [ q\u0307 \u03b4\u0307 ] = [ l1,2 \u00b7 sin(q + \u03b4) + l1,1 \u00b7 sin \u03b4 l1,2 \u00b7 sin(q + \u03b4) l1,2 \u00b7 cos(q + \u03b4) l1,2 \u00b7 cos(q + \u03b4) + l1,1 \u00b7 cos q ]\u22121 \u00b7 [ x\u0307 y\u0307 ] . (41) Elements of the Jacobian are only functions of the elements of the homogenous transformation matrix Te 0 3. It is clear that each branched chain in the complex mechanism has its finite transformation matrix, as well as its Jacobi matrix. The term \u201ckinematics\u201d is commonly used in the terms of rigid robot systems because, in this case, mechanisms geometry defines the position and orientation of the robot tip"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003822_s10483-013-1706-9-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003822_s10483-013-1706-9-Figure1-1.png",
        "caption": "Fig. 1 Model of multi-disk rotor-bearing system with twenty-two-degree-of-freedom",
        "texts": [
            " The result provides a new way to analyze the high-dimensional rotor-bearing systems by the bifurcation theories. The analysis of this paper is useful in fault diagnosis and the optimization of system parameters and provides theoretical guidance for the nonlinear dynamical design of rotor-bearing systems. 2 System description and modeling The high pressure rotor-bearing system of an aero engine with a pair of liquid-film lubricated bearings is modeled as a twenty-two-degree-of-freedom system, as shown in Fig. 1. The axial and torsional vibrations of the rotor and the gyro-moment are assumed to be negligible. oi (i = 1, 2, \u00b7 \u00b7 \u00b7 , 11) are the geometric centres of the discs, m is the equivalent lumped mass, ci (i = 1, 2, \u00b7 \u00b7 \u00b7 , 11) are the equivalent damping coefficients at the position of the lumped mass, and ki (i = 1, 2, \u00b7 \u00b7 \u00b7 , 10) are the equivalent stiffnesses of the corresponding discs. The equation of motion of the system can be obtained by the Lagrange method, and the dimensionless form can be obtained, see (A1) and (A2) in Appendix A for details, where the nonlinear oil-film forces[15] are described as follows:( fx fy ) = \u2212 ((x \u2212 2y\u0307)2 + (y + 2x\u0307)2)1/2 1 \u2212 x2 \u2212 y2 \u00b7 ( 3xV (x, y, \u03b1) \u2212 sin \u03b1 G(x, y, \u03b1) \u2212 2 cos\u03b1 S(x, y, \u03b1) 3yV (x, y, \u03b1) + cos\u03b1 G(x, y, \u03b1) \u2212 2 sin\u03b1 S(x, y, \u03b1) ) , (1) where \u03b1 = arctan y + 2x\u0307 x \u2212 2y\u0307 \u2212 \u03c0 2 sign y + 2x\u0307 x \u2212 2y\u0307 \u2212 \u03c0 2 sign(y + 2x\u0307), G(x, y, \u03b1) = 2 (1 \u2212 x2 \u2212 y2)1/2 (\u03c0 2 + arctan y cos\u03b1 \u2212 x sin \u03b1 (1 \u2212 x2 \u2212 y2)1/2 ) , V (x, y, \u03b1) = 2 + (y cos\u03b1 \u2212 x sin \u03b1)G(x, y, \u03b1) 1 \u2212 x2 \u2212 y2 , S(x, y, \u03b1) = x cos \u03b1 + y sin \u03b1 1 \u2212 (x cos \u03b1 + y sin \u03b1)2 , in which the function \u03b1 can be written as \u03b1 = arctan y + 2x\u0307 x \u2212 2y\u0307 \u2212 \u03c0 2 (y + 2x\u0307)(x \u2212 2y\u0307) |y + 2x\u0307||x \u2212 2y\u0307| \u2212 \u03c0 2 y + 2x\u0307 |y + 2x\u0307| "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001835_s12239-010-0044-y-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001835_s12239-010-0044-y-Figure1-1.png",
        "caption": "Figure 1. Photograph of different types of constant velocity joints.",
        "texts": [
            " This is mainly due to both technological enhancements, which reduce the vibration and noise of the powertrain, and the mature driving skills of ordinary people. In this study, the axial forces were measured. The assembly modules of different types of tripod CV joints were used for the measurements, and a review of the results was very helpful for understanding the relationships between idle motions and the spider locations in the CV joint. Car manufacturers have devoted considerable effort to improving the noise and vibrations that are related to CV joints. Based on the TJ-type joint, the internal components, as shown in Figure 1, have been modified to solve the shudder problem. The FTJ (free-ring tripod joint), SFJ*Corresponding author. e-mail: jirehk@kookmin.ac.kr (shudder-less free-ring joint), PTJ (pillow-type tripod joint), and UTJ (u-type tripod joint) are examples of modified CV joints that are applied in various vehicles with specific purposes. However, these joints are designed to first address the shudder of the vehicle; thus, the emerging issue of idle vibration characteristics was not carefully studied in previous designs"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002716_s11340-011-9514-z-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002716_s11340-011-9514-z-Figure1-1.png",
        "caption": "Fig. 1 Loading configurations: (a) parallel plate compression, (b) central load compression via a rigid sphere",
        "texts": [
            " Commercially available contact lenses are characterized using the two aforementioned loading configurations. The materials parameters will be extracted using the elastic model, and will be compared with standard tensile test of a straight strip and indentation. A linear, elastic and isotropic convex shell in the form of a spherical cap with elastic modulus, E, Poisson\u2019s ratio, v, thickness, t, radius of curvature, R, base radius c, sits in a circular recess with the periphery bounded by a rigid shoulder (Fig. 1). A number of restraints are applied to be consistent with the sample lens dimension and measurement, namely, (t/R)~0.011, (t/c)~0.013, and (w0/R)<1/8. It is noted that though the shell experiences large geometrical deformation, strain remains small. Sample lenses are fully immersed in an aqueous solution at all times to ensure the absence of meniscus and the associated negative Laplace pressure at the shell-substrate interface. Adhesion is safely ignored in the theoretical model. Parallel Plate Compression Figures 1(a) and 2 show the cross section and the loading configuration of the parallel plate compression",
            " Equations (1\u2013 6) for the outer region are solved numerically, and the computational details are given in the Appendix. Values of {w u \u03b2 V HMf} T are determined at every shell segment. Once H and u in the traction free annular shell and the constant A are found, {w u \u03b2 V HMf} T in the contact circle can be determined. The measurable quantities of central displacement, applied force, and the contact radius, respectively, are found to be w0 \u00bc w\u00fe R\u00f01 cos f \u00bb 0\u00de \u00f010\u00de F \u00bc 2pRV sin f \u00bb 0 \u00f011\u00de a \u00bc a\u00f0f\u00bb0\u00de \u00fe u \u00f012\u00de Central Load Compression External load can also be applied via a small solid rigid sphere (Fig. 1b). Upon loading, the shell conforms to the spherical geometry within the central contact region, and remains so without slippage throughout subsequent loading. A circular dimple forms around the center. Increase of external load raises both the contact radius and dimple radius. Equations (1-6) remain valid though the boundary conditions change accordingly. It is mathematically involved to capture the ever expanding contact circle during loading. We note that a ring load causes a \u201csnap-through\u201d buckling of a hemispherical shell at a critical load, but, if the same load is delivered via a concentrated area, the shell apex deforms continuously without buckling"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000587_2007-01-2232-Figure10-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000587_2007-01-2232-Figure10-1.png",
        "caption": "Fig. 10 Torsional Damper Cross-Sectional View",
        "texts": [
            "5 Time (s) Gasoline engine TDC signal MG2 feedback control Crank angle calculation Engine torque fluctuation estimation Conventional controller Filter MG1 feed-forward control MG1 MG2 Angle Angle Wheel speed estimationGain MG2 Speed 1200 800 400 0 Vibration control: OFF Motoring start Ignition start F lo or a cc el er at io n (m /s 2 ) E ng in e sp ee d (r pm ) Vibration control: ON Torsional angular velocity The torsional vibration level of the drive train can also be reduced by increasing the damping of the torsional damper. Increasing the hysteresis torque is an effective way to increase damping, but it raises the concern of worse booming noise of engine explosion first-order component because the torque fluctuation to the drive train increases. In consideration of the different torsional angles of these two phenomena, a damper with twostage hysteresis characteristics was adopted (Fig.10). During engine start, the hysteresis torque is high in the high-amplitude region, but the hysteresis torque is low in the very low-amplitude region where booming noise tends to be a problem (Figs. 11 and 12). In a vehicle powered by a V6 engine, in addition to the third-order component of engine revolution, the booming noise of the second-order component of engine revolution sometimes becomes a problem. The secondorder excitation force of revolution is mainly a secondorder couple of the reciprocating inertia of the piston"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002400_1.4006324-Figure24-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002400_1.4006324-Figure24-1.png",
        "caption": "Fig. 24 Deformation of the loaded tooth of the right web gear (web angle 5 0 deg)",
        "texts": [
            " To understand the deformation characteristics of the loaded tooth of the thin-rimmed gear well, an image of the deformed right web gear with a web angle of 0 deg is shown in Fig. 23(a), and an enlarged view of the loaded tooth is also shown in Fig. 23(b). Figure 23 aids in explaining the deformation characteristics of the loaded tooth of the thin-rimmed right web gear with a web angle of 0 deg in the following discussion. Based on Fig. 23, the tooth positions before deformation and after deformation are sketched in Fig. 24(a). From Fig. 24(a), the deformation of the loaded tooth can be roughly divided into two types of deformation: one type is an upward and downward deformation of the loaded tooth as shown in Fig. 24(b), and the other type is a rotation deformation of the loaded tooth as shown in Fig. 24(c). In Fig. 24(b), end A of the loaded tooth has a downward deformation, and end B has an upward deformation. This is because end A is further away from the web than end B, so end A is more flexible than end B. Thus, end A can be deformed more easily than end B. Therefore, when the tooth is very rigid, an inclined deformation (end A is down and end B is up) of the loaded tooth, as shown in Fig. 24(b), occurs. In Fig. 24(c), when the tooth is loaded, end A moves toward the right and end B moves toward the left. It appears that the loaded tooth rotates around the web (the web is the axis). This deformation is called a rotation deformation of the loaded tooth in this paper. The asymmetrical web position (the web center is offset from Journal of Mechanical Design MAY 2012, Vol. 134 / 051001-9 Downloaded From: http://mechanicaldesign.asmedigitalcollection.asme.org/ on 01/28/2016 Terms of Use: http://www.asme.org/about-asme/terms-of-use the tooth center) can be considered to be the main reason for the tooth rotation deformation. As stated above, because end A is further away from the web than end B, the rigidity of end B is greater than that of end A. Therefore, end B shares most of the tooth\u2019s load as shown in Fig. 8(a). When the tooth is very rigid, the greater tooth load on end B will cause the tooth to rotate around the web, as shown in Fig. 24(c). When the rotation deformation of the loaded tooth occurs, end A of the loaded tooth will approach the neighboring tooth, and end B will move far away from the neighboring tooth on the right side, as shown in Fig. 24(a). This allows the tooth root of the loaded tooth to experience a counterintuitive compressive stress on end A and a tensile stress on end B on the side of the loaded tooth surface. When the web is inclined, the supporting rigidity of the web to the teeth will become smaller. In this case, the deformation of the loaded tooth will be affected by the web\u2019s supporting rigidity, and the deformation of the loaded tooth will become more 051001-10 / Vol. 134, MAY 2012 Transactions of the ASME Downloaded From: http://mechanicaldesign"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003998_s0026261713040139-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003998_s0026261713040139-Figure1-1.png",
        "caption": "Fig. 1. Diagram of a microbial fuel cell (MFC).",
        "texts": [
            " Comparable results were obtained for strain FRB1. Thus, the rates of oxidation reduction pro cesses were higher in the mutants than in the original strain. These results suggested the application of these mutants in microbial fuel cells. Electricity generation in MFCs by Shewanella strains. Numerous modifications of MFCs exist, dif fering in the chamber volume, anode and cathode materials, distance between the electrodes, etc. An MFC consists of two chambers (anode and cathode ones) separated by an ion selective membrane (Fig. 1). Organic matter and bacteria are located in the anode chamber under anaerobic conditions. The cathode is under aerobic conditions. The membrane permits proton transport from the anaerobic anode chamber into the aerobic cathode one and prevents oxygen diffusion into the anode chamber. The anode is connected with the cathode via an electric circuit with a certain resistive load. The electrons arrive to the ter minal acceptor (a proton in the cathode chamber) via the anode and the electric circuit, thus creating the current [8]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002569_1350650112451218-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002569_1350650112451218-Figure1-1.png",
        "caption": "Figure 1. Plain cylindrical journal bearing.",
        "texts": [
            " An EHD analysis is briefly described to enable a better understanding of which phenomena occur with bearing under consideration. The applied radial load varies from 10 to 150 kN and the relative shaft displacements induced by the deformations of the bushing due to pressure fields are taken into account. An experimental study is presented in this article on the variation of the radial load, which aims to make a comparison with numerical analysis. In the case of bearing operating in steady state,2 the generalized Reynolds equation under classic assumptions in the (o, ! , z!) coordinates system (Figure 1) is expressed in the following form 1 R2 @ @ g \u00f0h\u00de3 @p @ \u00fe @ @z g \u00f0h\u00de3 @p @z \u00bc 6! @h @ \u00f01\u00de In this expression, is elasticity modulus that varies with the lubricant compressibility in the fullfilm region. is expressed by \u00bc @p @ The pressure p is given by p \u00bc g ln \u00fe pC \u00bc C \u00f02\u00de pC is the cavitation pressure. In the full-film zone, is the ratio of the local fluid density to the oil cavitated density C. In the rupture zone, represents the local fraction of the bearing gap occupied by lubricant. The cavitation index g is equal to 0 in the cavitation zone ( < 1) and equal to 1 in the full-film zone ( 5 1)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003501_2041302510394742-Figure4-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003501_2041302510394742-Figure4-1.png",
        "caption": "Fig. 4 Rotor-fixed wing hybrid aircraft: (a) the whole structure of the hybrid aircraft, (b) in VTOL flight or in hover, (c) rotating free wings in hover, (d) mode transition from hover to forward flight or vice versa, and (e) forward flight",
        "texts": [
            " The fixed-wing root portions in horizontal flight mode perform as the wings by generating lift in association with the free wings. Therefore, the normalized aerodynamic lift force and drag force can be obtained even during flying modes transition. Moreover, a backstepping controller is designed to control the hybrid aircraft to keep the flying height invariant during modes transition. The flying modes transition is shown in Fig. 3. 2 DESIGN OF ROTOR-FIXED WING HYBRID AIRCRAFT A designed rotor-fixed hybrid aircraft is shown in Figs 4(a) to (e). When the aircraft is in VTOL flight or in hover (see Fig. 4(b)), the thrust generated by the co-axial counterrotating propellers 1 provide lift force. The fixed-wing root portion 3 advantageously remains in the slip stream and dynamic pressure acting thereon tends to provide some degree of directional stability. From reference [23], rotating the free wings 2, the portions of propeller wash will not be blocked by the free wings 2 and propeller wash only operate over the fixed-wing root portion 3 and fuselage 4. Controlling the differential setting of the pairings 5\u20138, respectively, yaw dynamics are regulated. Free wings 2 controlled by gearing arrangements 19 can keep a given small angle of attack during mode transition. The free wing with gearing arrangements 19 will be introduced. To transition from hover to horizontal flight, the gearing arrangements are controlled. Assuming that the aircraft is hovering (see Fig. 4(c)), the aircraft is initially lifted by the co-axial counter-rotating propellers 1. The free wing 2 is rotated to a given angle of attack. As the pairings 7 and 8 swerve forward, a backward pitch moment is generated by tail propeller 9 with the deflection pairings 7 and 8 (see Fig. 4), and the fuselage 4 pitches towards the horizontal which in turn causes the horizontal speed of the aircraft to increase (see Fig. 4(d)). The rotatable free wings 2 are controlled by gearing arrangements and obtain Proc. IMechE Vol. 225 Part G: J. Aerospace Engineering 832 at The University of Manchester Library on April 23, 2015pig.sagepub.comDownloaded from a given angle of attack in accordance with the relative wind. The effects of relative wind acting on the free wings 2 quickly overcome the braking effects of the air flow over the fixed-wing root portion 3 from the propulsion system 1. The gravity of the aircraft is counteracted mainly by the vertical force of the thrust generated by the co-axial counter-rotating propellers 1, and the height of the aircraft is kept invariant. With increasing horizontal speed, wings 2 and 3 develop lift. The aircraft soon transitions into horizontal flight in a free wing straight and level flight mode (see Fig. 4(e)). For transition from horizontal flight to hover, the reverse procedure is used. As the pairings 7 and 8 deflect upward, the fuselage 4 pitches towards the vertical which in turn causes the horizontal speed of the aircraft to decrease and the vertical thrust vector gradually increases to keep the flying height invariant. The aircraft thus slows and converges to hover. The advantage of the rotating wing is that transition to stall can be delayed, and the co-axial counter-rotating propellers can provide lift force to keep the height invariant although the wing will still stall as forward velocity goes to zero"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000645_s00170-007-1183-9-Figure3-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000645_s00170-007-1183-9-Figure3-1.png",
        "caption": "Fig. 3 Free-body diagram of the moving platform",
        "texts": [
            " Pai is the spherical joint position in local coordinate frame {P}. Superscripts of P and W represent the local and reference coordinate frames, respectively, and whenever the coordinate system is not specified, frame {W} would be inferred. The vibration behavior of the moving platform as the end effector carrying the workpiece during machining operations is the main concern in analyzing the vibration of the hexapod table. This is the subject of subsequent sections. The free-body diagram of the moving platform is illustrated in Fig. 3. In this figure, u P and q P are the linear and angular acceleration of the platform expressed in reference coordinate frame {W}, respectively; PFmac and PMmac, the harmonic machining force and moment vectors in local coordinate frame {P} being arbitrarily exerted to the moving platform, respectively. The gravity and coriolis forces are ineffective in vibration analysis and have been ignored in this work. FKi and FCi are the total stiffness and damping forces, respectively, exerted to the platform and can be obtained as follows: FKi \u00bc KTi \u0394lTi ; FCi \u00bc CTi \u0394l Ti \u00f01\u00de where KTi and CTi are the equivalent stiffness and damping coefficient of the support carrying the moving platform at the junction of this platform and the ith spherical joint, respectively; \u0394lTi and \u0394l Ti are the absolute displacement and velocity of this junction along the ith pod\u2019s axis and can be written as follows: \u0394lTi \u00bc \u0394lui \u00fe\u0394lsi \u00fe\u0394lai \u00fe\u0394ldi \u00fe\u0394lli \u00fe\u0394lbi \u00f02\u00de \u0394l Ti \u00bc \u0394l ui \u00fe\u0394l si \u00fe\u0394l ai \u00fe\u0394l di \u00fe\u0394l li \u00fe\u0394l bi \u00f03\u00de where \u0394lsi;\u0394lui;\u0394lai;\u0394ldi;\u0394lli and \u0394lbi are the displacements of the ith spherical joint, pod\u2019s upper part, sliding joint, pod\u2019s lower part, universal joint and the junction of the lower platform with the ith universal joint, respectively; \u0394l si;\u0394l ui;\u0394l ai;\u0394l di;\u0394l li and \u0394l bi are the corresponding velocity increments",
            " (1\u20135), the equivalent stiffness of the ith supporting chain beneath the upper platform, KTi, and similarly the equivalent damping coefficient, CTi, can be found as follows: 1 KTi \u00bc 1 Kui \u00fe 1 Ksi \u00fe 1 Kai \u00fe 1 Kdi \u00fe 1 Kli \u00fe 6 Kb \u00f06\u00de 1 CTi \u00bc 1 Cui \u00fe 1 Csi \u00fe 1 Cai \u00fe 1 Cdi \u00fe 1 Cli \u00fe 6 Cb \u00f07\u00de In the above relations, it is implicitly assumed that the lower platform can be modeled as six concentrated vibratory elements; each situated under one pod, and combined in parallel. The stiffness coefficient of each element can be found by modal experiment. However, it is possible to model the deformation of each element by a homogeneous beam with uniform cross section. The stiffness of that element can then be found as follows: Ke \u00bc EeAe le \u00f08\u00de where Ee is Young modulus; Ae, the area of the cross section and le, the length of the element. Considering the free-body diagram of the moving platform illustrated in Fig. 3, the force equilibrium (Newton) equation can be written as follows: MPu P \u00fe X niFCi \u00fe X niFKi \u00bc WRP PFmac \u00f09\u00de where ni is the unit vector along the ith pod axis in coordinate frame {W}. The moment equilibrium (Euler) equation ( P M \u00bcIPq P; M is the moment) about the geometry center of the moving platform can be expressed as follows: wRP PMmac \u00fe GC PFmac X wRP: Pai niFCi X wRP: Pai niFKi \u00bc IPq p \u00f010\u00de where GC is the position vector of the external force in the local coordinate system. Using the following definition, qai \u00bc wRP: Pai \u00f011\u00de Equation (10) can be stated as follows: IPq P \u00fe X qai niFCi \u00fe X qai niFKi \u00bc wRP PMmac \u00fe GC PFmac \u00f012\u00de Substituting Eq"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001724_tmag.2010.2072910-Figure3-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001724_tmag.2010.2072910-Figure3-1.png",
        "caption": "Fig. 3. Magnetic flux density. (a) Slip . (b) Slip . (c) Slip .",
        "texts": [
            " The electrical steel sheet was assumed to be 50A470, and the rolling direction was parallel to 0 (horizontal direction of this figure). As shown in Fig. 2, three-phase stator windings were distributed. The model core was divided into the linear triangular elements. The number of nodes was 22 936, and the number of elements was 44 263. In this mesh, the air gap was divided into eight layers. In the analysis, the periodic boundary condition of 180 was used, and the fixed boundary condition was employed at the outer air region. Fig. 3(a)\u2013(c) shows the distributions of the maximum magnetic flux density at slip 0, 0.2, and 0.5, respectively. At the no-load condition slip as shown in Fig. 3(a), because influence of eddy current was very small, the magnetic flux density increased at the back yoke part of the slots. At the small slip condition slip as shown in Fig. 3(b), the magnetic flux density did not increase so much in comparison to that of Fig. 3(a). This means that the eddy current generated in the rotor bars because of the phase lag of the secondly side behind primary side flux. As mentioned, at slip , as shown in Fig. 3(c), the magnetic flux density became smaller at the back core part in comparison to ones of Fig. 3(a) and (b). However, in such larger slip condition, the magnetic flux concentrated at the teeth edge part due to the eddy current. Also, it is evident that the magnetic flux density is easily increased in the rolling direction (RD: horizontal direction of the figures), because of higher magnetic permeability. The magnetic permeability in the rolling direction is about 1.5 times larger than that in the transverse direction. Fig. 4(a)\u2013(c) shows the distributions of the maximum magnetic field intensity of each condition (slip 0, 0.2, and 0.5). As shown in Fig. 4(a), as well as Fig. 3(a), the magnetic field intensity was the largest because the eddy current was very small. With increasing the load and the slip, the magnetic field intensity became smaller as well. In addition, in the transverse direction (TD: vertical direction of the figures), the magnetic field intensity increased because of magnetic anisotropy. Fig. 5 shows the iron-loss distributions calculated from the vector relation between and . Usually, the iron loss is assumed to be in proportion with the square of the maximum magnetic flux density, "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001512_s11668-010-9398-8-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001512_s11668-010-9398-8-Figure1-1.png",
        "caption": "Fig. 1 (a) Heat source and temperature distribution [4] and (b) heat accumulation area Fig. 2 Gear heat transfer model [5, 6]",
        "texts": [
            " Using this method, the service life is extended, and the thermal damage that occurs on the gear tooth surface is reduced. The gear service life depends on many parameters, such as the gear temperature, gear stiffness, gear macro/micro-geometries, and operating conditions. For the operational tests in this study, the thermal damage at the surface was decreased because of the presence of the cooling holes compared to the damage seen in similar gears without the holes. Many studies have been performed to measure tooth surface temperatures. One such study was performed by Terashima et al. and Fig. 1 from that study shows the temperature profile of a tooth gear operating at room temperature. The highest temperature of the tooth surface is located near the pitch point region. In addition, the inner temperature was about 10 C lower than that of the tooth surface [4]. The heat generated in the plastic gears is a result of not only the frictional heat between the tooth surfaces but also of the heat caused by the hysteresis loss in the viscoelastic materials; it is, therefore, difficult to calculate the amount of the heat generated or the temperature distribution within the gear. In standard gears, the resulting heat can be partially removed because of convection from the surface. However, the heat accumulates on the elliptical region, as shown in Fig. 1. Owing to the accumulated heat on the outer surface of the gears, the plastic material degrades and exhibits bulging at the tooth flank [21]. Owing to the low thermal conductivity of plastic, nearly all of the generated heat must be removed by convection. The air temperature around the gears remains close to ambient, and there is only a small difference between the temperatures of the contacting and non-contacting flanks of the gears, which suggests the possibility of a simplified model [5, 16, 20]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001038_j.finel.2008.01.006-Figure7-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001038_j.finel.2008.01.006-Figure7-1.png",
        "caption": "Fig. 7. 1st and 2nd modes of the induction motor.",
        "texts": [
            " Besides, the mass eccentricity of the silicon steel core was lumped into nodal masses named as m23 through m28 corresponding to each of the element nodes 23 through 28. The radius of eccentricity, e, is defined as the distance of the located lumped mass eccentricity that is relative to the geometric center on each of the silicon steel laminations. Its value was initially assumed to be the same as the air gap with Cr = 0.35 mm for a series of analysis processes. In other words, the eccentricity ratio e/Cr = 1 is applied at the outset of the transient analysis. Fig. 7 shows the first two mode shapes of the induction motor as resulted from the calculation of the developed solution modules. As observed from the figure, the mode shapes are distinguished into the cylindrical and conical modes, respectively. Also, those nodes on the rotor with significant vibration amplitudes are located at the balance planes as seen in the cylindrical mode and are at the bearing locations as indicated in the conical mode correspondingly. A ratio of the nodal vibration amplitude to the air gap is defined as the amplitude factor = q/Cr"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001289_robot.2008.4543428-Figure3-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001289_robot.2008.4543428-Figure3-1.png",
        "caption": "Fig. 3. Stick diagrams and ankle trajectories of swing leg",
        "texts": [
            " In this section, the relationship between the flexion actions and the parameters of circular arc foot is demonstrated. Figure 8 shows the coefficient q in Eq. (17) when the parameters of circular arc foot are varied. The horizontal axis denotes the center angle of circular arc of foot \u03b4. The vertical axis denotes the curvature radius of foot \u03c1/l (dimensionless). The contour line denotes the coefficient q. If \u03b4 and \u03c1/l are increased, the flexion action of foot is enhanced. In addition, \u03c1/l has more influence than \u03b4. As shown in Fig. 3 (a), the swing leg swings through the stance leg while the knee of swing leg is inflecting. During this, passive walker has a high risk for striking its toe on the slope. In order to reduce the collision risk, the flexion action of knee of swing leg must be enhanced. Therefore, it is desired to take advantage of the flexion action of circular arc foot. At the same time, we must consider the disadvantage of the circular arc foot becoming an obstacle of swinging through the stance leg. In this paper, we derived the equation of angular acceleration of knee joint from the simplified and linearized model"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000980_ac801289t-Figure3-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000980_ac801289t-Figure3-1.png",
        "caption": "Figure 3. Schematic of the spectroelectrochemical cell: (a) sample inlet, (b) sample outlet, (c) reference electrode (Ag/AgCl), (d) SF6 coupling prisms, (e) electrical contact with counter electrode, (f) counter electrode (stainless steel foil), (g) gasket, (h) working electrode (ITO glass), and (i) cell base.",
        "texts": [
            " The angle of the incident light into the prism was adjusted to maximize the ATR throughput, as determined by measuring the intensity of the outcoupled light. Light propagated through the ITO glass slide was outcoupled with another SF6 coupling prism and focused on a photodiode detector (Photonic Detectors, Digi-Key PDB-V107). The ATR spectroelectrochemical cell made with black Delrin was equipped with a reference electrode, electrical contact with a counter electrode, and sample inlet-outlet holes as shown in Figure 3. The counter electrode (stainless steel foil, 0.05 mm thick) was sandwiched between 0.25 mm thick silicone gaskets and located between the Delrin spectroelectrochemical cell and ITO glass. The light path was 4.5 cm long. Electrochemical operation was done with a conventional three op-amp potentiostat built in our laboratory. All potential values were applied against an Ag/ AgCl reference electrode (EE 008 miniature reference electrode, Cypress Systems (Lawrence, KS)). Thin layer spectroelectrochemical measurements used a CV27 potentiostat (BAS, West Lafayette, IN) and a Hewlett-Packard 8453 diode array spectrophotometer",
            " 1998, 10, 2481\u20132489. (40) Limoges, B.; Degrand, C.; Brossier, P.; Blankespoor, R. L. Anal. Chem. 1993, 65, 1054\u20131060. (41) Chen, H. Y.; Ju, H. X.; Xun, Y. G. Anal. Chem. 1994, 66, 4538\u20134542. (42) Shi, Y.; Seliskar, C. J. Chem. Mater. 1997, 9, 821\u2013829. 9645Analytical Chemistry, Vol. 80, No. 24, December 15, 2008 silica, PAA-silica, PSSA-silica, and heparin-silica film coated electrodes that were first immersed in 10-5 M NB for 15 min. The measurements were done in the spectroelectrochemical cell shown in Figure 3. The repetitive potential scans between 0.5 and -0.8 V modulate the optical signal. The magnitude of change in the optical signal shows the amount of NB accumulated within the film. The highest modulation amplitude of NB is obtained with the Nafion-silica and PSSA-silica films. However, the width of the optical modulation peaks obtained at Nafion-silica is slightly greater than for the PSSA-silica. Optical modulation is not observed in the PAA-silica film because NB does not penetrate into the film"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001281_waina.2008.106-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001281_waina.2008.106-Figure1-1.png",
        "caption": "Figure 1: Multi-sink wireless sensor network. Base stations (BSs) can directly communicate with each other. In addition, the tree structure with the BSs as roots is constructed by the sensor nodes.",
        "texts": [
            " Besides our algorithm finds the shortest path from a sensor node to the nearest BS, it calculates the hop counts from a sensor node to all BSs deployed in the WSN, respectively. They could contribute to the location estimation of each sensor node. The remainder of this paper is organized as follows. The WSN model is explained in Section 2 and the algorithm is introduced in Section 3. In Section 4, some experimental results are shown; finally, the conclusions and future work are described in Section 5. 2. Sensor Network Model In this paper, we consider the WSN model shown in Fig. 1. The WSN consists of multiple base stations (BSs) and many sensor nodes. An observer can obtain sensed data and transmit sensing requirements from/to the WSN through the BSs. Each sensor node has wireless communication, data processing functions, and sensors. Although the functioning of each sensor node is relatively simple, the observer can obtain environmental information from many networked sensor nodes. Here, we assume that the WSN is organized autonomously. The organizing steps are as follows",
            " In general, TDMA control is realized in a centralized manner, e.g., the BS allocates a time slot to each node [3]. However, it is difficult to implement approach because our target is a decentralized WSN. We apply our coloring algorithm to the time slot allocation in TDMA. It enables an autonomous and dynamic time slot allocation. The behavior of the proposed system can be summarized as follows. A multiple tree structure whose nodes comprise the deployed sensor nodes is constructed and maintained by using the distributed algorithm. Figure 1 shows an example of a multitree network topology. The root of each tree is a BS. Data aggregation is performed using the trees constructed with the TDMA. To realize the system, the following process should be performed in each sensor node. \u30fb Maintain and update a list of its neighboring nodes. \u30fb Determine the \u201clevel\u201d and \u201cparent\u201d to construct the multi-sink shortest paths, i.e., the tree topology with multiple BSs. \u30fb Send the sensed data to the parent node. \u30fb Perform time slot allocation by using the distributed coloring algorithm to avoid packet collisions",
            " We assume that the number of time slots is more than the number of BSs deployed. The pseudocodes of the time slot allocation algorithm are shown in Fig. 9. This algorithm features the packet collision avoidance mechanism as mentioned above. 3.3 Calculation of Hop Counts from a Node to All BSs In WSN, location of each sensor node is important for many protocols such as a routing and data aggregation. Hence, a location estimation method for sensor nodes is often required. In our WSN model shown in Fig. 1, if each sensor node knows the minimum hop count from each of BSs, the relative location of the node can be estimated from the obtained hop counts. Though the location estimation is not main task of our proposed algorithm, we include a mechanism that calculates the minimum hop-counts from all BSs surrounding the sensor node. Figure 10 shows the procedure. In the procedure, the minimum hop-count or the shortest path from each BS to the sensor node is calculated using the received packet information stored in list P"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001546_j.jmatprotec.2010.12.002-Figure9-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001546_j.jmatprotec.2010.12.002-Figure9-1.png",
        "caption": "Fig. 9. Temperature distribution on the",
        "texts": [
            " / Journal of Materials Processing Technology 211 (2011) 675\u2013687 681 e t v c i a p o t o t i o i T w b t fi h xperimental results; therefore, experiments were repeated five imes and the error was estimated in the order of 4%. Temporal ariation of temperature predicted and obtained from the thermoouple data is shown in Fig. 8. It is evident that both results are n good agreement. The differences in both results may be associted with experimental error and the assumption of the uniform roperties in the simulations. Fig. 9 shows two-dimensional view of temperature distribution n the outer surface of the tube at four different times. The effect of he moving heat source on the distribution of temperature on the uter surface of the tube with time and the rapid heat transfer in he welding zone is evident from this figure. It also appears that the sotherm line presents ellipse. The isotherm line is dense in front f the laser source and the temperature level is high there while n the back of the moving laser source, the complexion is contrary",
            " Temporal variation of von Mises stress at three different locations. outer surface at different times. the attainment of high thermal strain along the weld path. The von Mises residual stress fields are identical in trend and magnitude on both outer and inner surfaces of the tube. Thermal stress developed during high temperature heating contributes to the stress field developed during the cooling period. Consequently, von Mises stress at any time does not follow exactly the temperature distribution as shown in Fig. 9. Fig. 15 shows temporal variation of von Mises stress at the same locations shown in Fig. 12. von Mises stress increases as the heat source approaches these locations with increasing temperature. During phase change, stress value reduces with minimum at the peak temperature when the substrate material is in molten state. However, von Mises stress attains high values for the time when temperature decay is large. This follows by gradual increase in von Mises stress with the progressing time. Moreover, two local maxima (sharp peaks) are observed in von Mises stress at each location shown in the figure"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002335_detc2010-28926-Figure3-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002335_detc2010-28926-Figure3-1.png",
        "caption": "Figure 3. Example of a graph that is not an Assur Graph",
        "texts": [
            " The main property of the graph is that removal of any vertex with its incident edges makes the graph non-rigid. The graph, appearing in Figure 2(a) is an Assur Graph since the number of the edges is twice the number of the inner vertices, it is rigid and all its sub-graphs are not rigid. For example, the graph in Figure 2(b) is obtained from the graph Figure 2(a) by deleting vertex C and all its incident edges, resulting in a linkage. The system in Figure 2(c) is obtained by deleting vertex D and is also a linkage. In contrast, the structure in Figure 3(a) is not an Assur Graph since deleting vertex C results in an Assur Graph, known as the Triad, shown in Figure 3(b). A B A B A B a) Assur Graph. b,c) The graphs after deleting vertices C and D, respectively. In each Assur Graph there are two types of vertices: ground vertices, called also pinned vertices, and inner vertices. For example, in a triad type Assur Graph (Figure 3b) there are three inner and ground vertices while in the dyad type Assur graph there are two ground vertices and one inner vertex. The composition rule for constructing a determinate truss from its components (Assur Graphs) is done as follows: Let G1 and G2 be two Assur Graphs. G1 is defined to be preceding G2 if at least one ground vertex of G1 is connected to an inner vertex of G2. The decomposition process can be presented by a directed graph in which an edge e=<u,v> indicates that the Assur Graph corresponding to vertex u is preceding another Assur Graph, presented by vertex v"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000757_1.1663415-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000757_1.1663415-Figure1-1.png",
        "caption": "FIG. 1. Coordinates chosen to describe the half-wall a termi nating along the junction Oz.",
        "texts": [
            " Kleman: Stresses due to magnetic wall junctions Finally, the total displacements, including those coming from the e~j' are given by9,10 U = 0 3(1-211) X (1 -1)- X2 6W Ca )\",Ca) Xl 411\"(1 _II) 1 np 211\" a 3 'I' , 03(1 - 211) ) Z!. \"\" Ca) Ca) U\"2 = 4(1-11) x2 (Inp - 1 + 211\" L;! W3 cf> , U =0. X3 (28) In this expression, w~a) is the component of the rotation 0 3 due to wall a and cf> Ca) is the azimuthal angle of the point (Xll X 2) from the axis Ox Ca ), such that the half wall a is situated on the negative part of this axis, -11\" <cf>a <+11\" (see Fig. 1). Let us notice that when the w~) add up to zero the junction does not produce elastic strains. We shall call such a junction a Nye junction; its effects are pure ro tations. Clearly, any N~elline on a plane Bloch wall is a Nye junction. Hence, the elastic strains produced by a N~elline (or a Bloch line on an infinite Neel wall) can not be analyzed by the present formalism, which does not take into account the distribution of M inside the wall. However, we get that result that strains due to Bloch or Neellines are very short range, quite certain ly screened at a distance of the order of the wall thickness",
            " Such complex defects are still to be studied. Let us be re minded that a similar approach has already been under taken by Li,13 who computed the elastic distortions of polygonal dislocations by using as a unit the elastic dis tortions due to each segment of the dislocation. By starting from Eq. (26), the dislocation and dis clination densities for a given half-wall a are given, in the frame attached to the wall, by (29) all other components vanish. We assume S to be along the negative part of the xia ) axis, as it is in Fig. 1. We also introduce, in conformity with Eq. (24), the quantities Et) and wr): E:~) = ll.e~l , E~) = - [11/(1 -1I)](ll.e~l + ll.e~3) , J. Appl. Phys., Vol. 45, No.3, March 1974 E Ca)' A 0 33 =~e33' E;:')=ll.e~3' E:~)=O, E~)=O, w1v,) = -ll.e~2' w2v,) = 0, w~a) = ll.e~2 , where the superscripts (a) have been omitted in the second members fo r the sake of clarity of writing. We get the following results: (a ) cf> v,) v, ) 1 . v, ) Ca ) Ca ) w~a ) ell =21T Ell - 411\"11 smcf> coscf> E22 + 41T(1 _II) x[(1-211)lnp+sin2 cf>Ca) +1-411], cf> Ca) 1 w(a) e Ca ) = --ECa) + - sincpV,) cos'\" V,)EV,) + 3 22 21T 22 41T1I 'I' 22 41T(1 _ II) X[(1-211)lnp + cos2cf>(a) + 1-4v], cf> (a) e(a)---E(a) 33 - 211\" 33, 1 E(a) w(a) 1380 (30) (31) e(a)=-(E(a) _ECa\u00bb(l +lnp) _ :::..22.... cos2cf> Ca) _ 3 12 411\" 11 22 811\"v 41T(1 _ II) Xsincf> (a) cos2 cf>(a) v,) E;:') e 23 = 41T (1 +lnp) , Ca) cf> Ca ) (a) wl Ca ) e31 = 21TE 31 - 411\" (1 +lnp) , KV,)= _l_(coscf> (a)E Ca ; - sin\",(a)\",v,\u00bb 11 21Tp 31 'I' 1 , K2~)= -2 1 (-sincf>Ca)E:~) +coscf>(a)w~\u00bb. 1I\"p B. Elastic fields of a junction (32) Finally, we add up the results of Eqs. (31) and (32) over all the walls. We express the sum in a frame X, Y. We define cf>a = (OX, Ox:a\u00bb (see Fig. 1). cf>a is an angle between -1T and +1T. Call a the polar angle in the frame X, Y. We have the following relation: [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: 137.207.120.173 On: Tue, 25 Nov 2014 14:40:26 1381 M. Kleman: Stresses due to magnetic wall junctions ua(e) = 0 for other values of e, where 0 <cf>a <1T Ua(e) = -1 for 1T - cf>a <e <1T, where -7r <cf>a <0 Ua(e) = 0 for other values of e, where -7r <cf>a <0"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000616_002029408001300701-Figure4-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000616_002029408001300701-Figure4-1.png",
        "caption": "Fig 4 Eden/s horizontal projector at NPL. c 1916 (Courtesy of NPL, Crown Copyright)",
        "texts": [],
        "surrounding_texts": [
            "Fig 3 Eden's first screw diameter measuring machine at NPL, 1915. (shown dismantled) (Courtesy of I\\JPL, Crown Copyright)\nThe last machine that Eden produced, before leaving the NPL to join the Research Department of GEC in 1918, was a 'millionth' comparator. Unlike Whitworth's so-called millionth machine, reiying on a micrometer screw, Eden's machine was a true comparator with a very small range of measurement. Such an instrument had become necessary with the increasing use of Johansson gauges and because a new method of manufacturing such gauges was being developed at NPL by A. J. C. Brookes. The comparator was simple in its basic design. The fixed measuring face carried three small balls to provide three-point contact with one face of the gauge and the moving measuring face had a spherical tip. Thus perfect kinematic location was achieved. All the moving components of the instrument were mounted on flexible strips and variation in gauge size was transmitted to a light aluminium pointer, movement of which was projected through an optical system onto a scale in front of the observer. An overall mechanical and optical magnification of 20000 was achieved, so that variations of a micro-inch could be readily estimated on the scale. Quite a number of these comparators were made commercially for metrology laboratories all over the world. It was the first instrument to measure accurately and reliably to the millionth of an inch (0.025 pm). It has only been superseded in recent years by high magnification electronic comparators but one large manufacturing company in England still uses the basic unit of such com parators, substituting for the optical system a sensitive electronic indicator. These are used for the routine checking of hundreds of sets of gauges in use in the factory.\nOptics and electronics\nThe period from the end of World War I to the present time has seen tremendous developments in metrology in industry. Increases in production volume and the introduc tion of automation have only been possible with the close control of dimensional accuracy in all components. The idea of the individual fitting of parts in an assembly had virtually disappeared by 1939, except in a very few cases of specialist manufacture. It would be impossible, in the space of this article, to do more than make brief reference to a few basic principles on which much of this development has been based. The two most significant fields are obviously optics and electronics. The younger reader might even query the importance of optics, in the light of the design of modern equipment when almost any required function can be achieved with a few transistors, micro-circuits and,\nMEASUREIVIENT A 1\\1 0 CONTROL, Vcl 13, July 1980\nsometimes, quartz crystals. Nevertheless, in its time, the use of optics revolutionised the control of accuracy to a similar, if not such a widespread, extent.\nWe have seen that, during the first war, optical methods were introduced for the direct magnification of gauge profiles, such as screw threads, and for the amplification of mechancial movements, as in Eden's instruments. These applications were, for the most part, for laboratory use although projectors had become widely used in industry by the end of the war. In the 1920 Leipzig Fair, the firm of Carl Zeiss of Jena, Germany, exhibited several instruments including a screw-thread measuring microscope and an optical comparator which they called the 'Optimeter'. These differed from the early NPL instruments in having robust enclosed optical systems, quite suitable for use in inspection departments and toolrooms, as well as in metrology laboratories.\nCarl Zeiss, born in 1816, worked as a mechanic in the University of Jena and, in 1846, set up on his own as an optical instrument maker, working mainly for the University. Ernst Abbe, born in 1840, became professor of physics at an early age and was to have a great influence on Zeiss and his company, first as consultant, then as director of research, becoming Zeiss's partner in 1875. Under the direction of the scientist and the practical mechanic, the company attained a high reputation for optical equipment, from binoculars to astronomical telescopes and planetariums. It was, of course, many years before they turned their attention to industrial instruments but, as has happened in many companies, advanced equipment made for their own use quickly became a saleable product. During the first war, they made micrometers and, in 1918, received orders for some 10000 such instruments. The 'Department for Technical Measuring Tools' was established on the day of the armistice in 1918. Being expert in optics, it would have seemed a natural step from purely mechanical instruments to those incorporating optical systems.\nAnother highly successful development was the incorporation of surveying instrument techniques in industrial equipment. In this way, the optical circular dividing head and dividing tables were introduced. Up to that time, the only dividing head in the toolroom was the worm and wormwheel device in which accuracy relied on the worm, the wormwheel and the dividing plate by which fractions of turns of the worm were set or measured. By incorporating a finely divided glass circle, totally enclosed in the body of the dividing head and observed through a microscope, accuracy of dividing was markedly\n239",
            "improved and, at the same time, the measuring process was separated from the mechanical operation of the machine. Even up to the present time, engineers have become so used to the transporting mechanism, such as a leadscrew, being also the sole measuring device that the conflict between these two functions is often overlooked.\nMany other instruments, mainly optical, were made by Zeiss and these were the mainstay of industrial metrology all over the world during the 1930s. Societe Genevoise of Switzerland and Adam Hilger of London made a few optical instruments, Hilger projectors being quite widely used. When World War 11 came, the supply of Zeiss instru ments dried up instantly and several optical firms in Britain started to make replacements of the same or similar design. These firms were Cooke, Troughton and Simms, Ltd; Optical Measuring Tools Ltd; Ca newly formed company) and E. R. Watts Ltd. CTS and Watts had been making surveying instruments, microscopes, etc, for many years. After the war, the combined company Hilger and Watts Ltd was formed and, in more recent years, this has been absorbed with Taylor, Taylor and Hobson Ltd into Rank Precision Industries. The Zeiss company split into two, one in West Germany and the other remaining at Jena in the Eastern sector.\nVarious small companies climbed onto the band wagon of electronics and made comparators with huge magnifica tions, simply by stringing together rows of valve amplifying stages. Most of these were unreliable electronically, but worse, their designers had little idea of the mechanical stability and rigidity, or the fine mechanical adjustments, also necessary with such equipment. Two companies, Sigma Instruments Ltd and Taylor, Taylor and Hobson Ltd, made very good electrical comparators and various electronic instruments were developed later.\nDuring the late 1930s, the problems of measuring the finish or texture of machined surfaces began to arise. The performance of engineering products, particularly aero engines, depended more and more critically on precise surface finishes, and not always 'the finer the better'. Richard E. Reason, of Taylor-Hobson, set to work on an instrument which would measure surface roughness in numerical terms and he developed the range of Talysurf instruments which, with later developments, have had world-wide success ever since. He was not the first; a good deal of work had been going on in America, but his machine was the first to trace over a surface with a fine diamond point and, through electronic amplification, produce a\n240\nMEASUREMENT AND CONTROL, Vol 13, July 1980",
            "profile graph of the surface. He had realised that the character of the surface was just as important as a single figure representing a roughness value and that a profile graph would allow this to be judged. Once a sample machined surface from a particular process was found acceptable, subsequent samples could be quickly judged and controlled from a meter reading alone.\nAn almost natural development from the surface texture instrument was one, also working on the same electronic magnification principle, to check roundness of machined components. Such a machine, the Talyrond, needed a highly accurate bearing to rotate the measuring stylus on a true axis and this was achieved with an accuracy of a few millionths of an inch. This was, of course, before air or fluid bearings had been invented. This machine also has been developed over the years and its more recent successor, the Talycenta, is capable of measuring a portion of a cylindrical surface and calculating its deviation from a mean radius. Such is the power of the computer.\nConclusion\nIt would be impossible here to attempt a survey of more recent and modern instruments, many of which perform similar functions in various ways. Fortunately, it is not difficult to obtain technical details from recent literature and from the manufacturing companies. It is however, more important to have traced some of the early history of the development of engineering metrology, much of which is hidden away in the dusty (literally) archives of scientific and engineering institutions.\nReferences\nPart 1\nBarrell, H. 1958. 'The bases of measurement', IProdE Journal, 37 (l). Barrell, H. 1962, 'The metre', Contemporary Physics, 3 (6). Barrell, H. 1968. 'A short history of measurement standards\nat the NPL', Contemporary Physics, 9 (3). Cochrane, R. C. 1966. Measures for progress, us Dept of\nCommerce.\nPart II\nAlthin, T. K. W. C. E. Johansson 1864-1953, AB C. E. Johansson, Stockholm. Jena Review. 1967. No 5, Carl Zeiss, Jena, Germany. Musson, A. E. 1966. Joseph Whitworth (IMechE. Exhibition\nBooklet),IMechE. National Physical Laboratory Annual Reports, HMSO 1900\nonwards. Rolt, F. H. 1936. E. M. Eden Memorial Lecture, Private\nPublication. The Science Museum. Production Engineering Gallery,\nS Kensington, London.\nCommon to both parts\nHume, K. J. 1980. A history of engineering metrology. MEP. (To be published shortly). Rolt, F. H. 1929. Gauges and Fine Measuremen ts, Macmillan. Rott, F.H. 1952. 'Developments in engineering metrology',\nThe Production Engineer, 31 (l).\nWEST DC50 Industrial Digital Process Controller\nAM50Auto/Manual Unit Featuring:\n\u2022 Wide choice of input and output options with PROM linearisation \u2022 Fully adjustable PlO control action with overshoot inhibit \u2022 Set point calibration accuracy 0.25% \u2022 Bumpless transfer in changing between manual and auto modes \u2022 Drives to safe level in fault condition \u2022 Simple operation through illuminated push buttons with clear legends \u2022 Drift-free output in the manual mode\ngul\"\"\" Gulton Europe Limited Brighton Sussex BN2 4JU t.L.a I Telephone (0273) 606271 Telex 87172\n...\nc\n\u2022 \\eVE SI DC SO .-;;;\nMEASUREMENT AND CONTROL, Vol 13, July 1980\n2*\n241"
        ]
    },
    {
        "image_filename": "designv11_12_0000160_icarcv.2006.345352-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000160_icarcv.2006.345352-Figure1-1.png",
        "caption": "Fig. 1. The Experimental platform - Eagle",
        "texts": [
            " In this paper, we discuss the capabilities of the research platform and sensor integration, and also elaborate on our linear systems identification using time and frequency domain for the largest platform. Finally, we discuss the future direction of our research. Experiments for this research were conducted on two different helicopter platforms. The largest of these was a 70kg Yamaha R-Max helicopter, already used for autonomous helicopter research at many other institutions [10],[11]. A second platform was constructed based on a 60 size Hirobo \u201dEagle\u201d helicopter kit shown in Figure 1. The latter model was converted to electric power using a brushless DC motor to remove the complications of an internal combustion engine. Both platforms were fitted with identical NovAtel OEM4-G2LRT2 GPS units. Differential corrections from a DGPS base station were provided using 19600 baud rate RF modems. The NovAtel units are configured for a 20Hz update rate. In real-time kinematic (RTK) mode the GPS units are quoted as having a circular error probable (CEP) of 2cm. We plan to perform higher risk testing with the Eagle helicopter owing to its lower replacement cost"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003895_ls.171-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003895_ls.171-Figure2-1.png",
        "caption": "Figure 2. Foil bearing cross-section schematic and coordinate systems.",
        "texts": [
            " \u2020E-mail: lahmar.mustapha@univ-guelma.dz Copyright \u00a9 2012 John Wiley & Sons, Ltd. characteristics for the safety operation, especially against the external excitations such as the unbalance mass forces. Therefore, it is needed to investigate theoretically and experimentally the steady-state and dynamic characteristics of such bearings for better designing of high-quality eco-friendly or free-oil rotating machinery. Figure 1 shows a picture of a typical elastically supported foil bearing. As illustrated in Figure 2, it schematically consists of a cylindrical shell (sleeve) lined with corrugated bumps (bump foil) topped with a thin flat foil (top foil). The bump foil serves as a support for the top metal foil, and its compliant feature allows the top foil to deform under the action of aerodynamic pressure. This latter is generated in the air film when the shaft (journal) rotates over a certain angular speed. Air foil bearings were constantly modified several times to improve their performances. In fact, there are three generations of air foil bearings"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001322_13506501jet290-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001322_13506501jet290-Figure1-1.png",
        "caption": "Fig. 1 Stator and rotor vane seals",
        "texts": [
            "eywords: seal, rotary, vane, elastomer, polytetrafluoroethylene, leakage, friction In a typical linear hydraulic actuator, a piston translates in to a sealed cylinder with application of high hydraulic pressure. This type of actuator is widely used in industrial, automotive, and aviation applications as it is a simple, robust, and efficient design. There are some applications, however, where space is precious and another type of actuator is better suited, namely, a rotary vane actuator (RVA for short; see the top two drawings in Fig. 1). An RVA contains a stator and a rotor with two or more vanes. The vanes create chambers between the rotor and the stator, \u2217Corresponding author: Tribology Group, Department of Mechanical Engineering, Imperial College London, 3 Princes Mews, Hounslow, Middlesex TW3 3RF, UK. email: g.nikas@imperial.ac.uk which are filled with hydraulic fluid. By controlling the pressure to at least two of the said chambers, the rotor is forced to rotate and exert force to an external mechanism. In this way, a rather bulky linear hydraulic actuator can be replaced by a more compact RVA and save valuable weight and space",
            " Other applications of RVAs include antiroll suspension systems on high performance cars as well as in industrial machine tools, for example, multi-directional tube bending equipment [1]. Efficient operation of an RVA is based upon the effective sealing of its fluid chambers. Principal sealing elements used in RVAs are either \u2018Goalpost\u2019 or \u2018Window\u2019 type arrangement, referring to their configuration. JET290 \u00a9 IMechE 2007 Proc. IMechE Vol. 221 Part J: J. Engineering Tribology at Purdue University on June 5, 2015pij.sagepub.comDownloaded from In the studied case (Fig. 1), these are of rectangular cross-section and, having a single horizontal and two vertical sections, resemble a goalpost. They are mainly constructed of an elastomeric material. As the shape to be sealed is rectangular, all corners of the vane seals must be clean and sharp; otherwise, fluid leakage will take place between RVA chambers when the hydraulic pressure is applied, resulting in power loss (loss of transmitted torque) and reduction of efficiency. In applications such as the control of passenger-aircraft wing surfaces, such power loss from ineffective sealing could have obvious and serious consequences both in terms of human and economic risk",
            " In the most recent research project [44] and in an application involving the control of aircraft wing surfaces via an innovative RVA for a major passenger-aircraft manufacturer, vane seals were modelled and two papers published [47, 48], which were based on the model developed in reference [44] for the horizontal section of the vane seals (Fig. 2). In the present study, a brief presentation of the RVA seal modelling and optimization work is attempted, involving the whole (three-dimensional) goalpost vane seals, in an effort to show the performance of such seals and how this can be analysed and optimized. A composite vane seal as shown in Fig. 2 comprises two vertical sections such as ABCDEF and a horizontal section EFGH. There are two types of vane seals in an RVA as in Fig. 1: (a) a stator seal with its housing being part of a stationary body (stator) and the seal acting on a moving surface (rotor) (b) a rotor seal with its housing being part of a rotating body (rotor) and the seal acting on a stationary surface (stator). The two bottom drawings in Fig. 1 depict the horizontal sections of the two vane seals. The seals are (generally) shown as having a corner radius r at free state in the direction of motion (x), which is owed to either the normal production process or to wear. The dashed lines in Fig. 1 represent the seals in their housings with the counterfaces absent. Hydraulic fluid under sealed pressure pcyl may exist either side of a seal housing (left or right). A seal section (either horizontal or vertical) has, generally, initial dimensions bx , by , and bz , and is installed in its housing with initial interferences \u03b4x , \u03b4y , and \u03b4z in the three principal axes. Finally, the seals modelled here are composite with one elastomeric and two PTFE parts as in Fig. 2, the relative sizes of which can be changed in the model to resemble any intermediate state, from a purely elastomeric seal (by nullifying the width of the PTFE parts) to a purely PTFE seal (by nullifying the width of the elastomeric part)",
            " Normal strains on a seal from the initial interferences are \u03b5x = \u2212 \u03b4x bx , \u03b5y = \u2212 \u03b4y by , \u03b5z = \u2212 zsr bz \u2212 zr (1) where zsr = { zsta = \u03b4s \u2212 Rr \u2212 zr + \u221a R2 r \u2212 x2, (stator seal) zrot = \u03b4r + Rs \u2212 zr \u2212 \u221a R2 s \u2212 x2, (rotor seal) } (2) zr = \u23a7\u23aa\u23a8 \u23aa\u23a9 r \u2212\u221a r2 \u2212 [x \u2212 sgn(x)xr]2 for x \u2208 [\u2212xr \u2212 r, \u2212 xr] \u222a [xr, xr + r] 0 for |x| xr \u23ab\u23aa\u23ac \u23aa\u23ad (3) JET290 \u00a9 IMechE 2007 Proc. IMechE Vol. 221 Part J: J. Engineering Tribology at Purdue University on June 5, 2015pij.sagepub.comDownloaded from where xr \u2261 (bx \u2212 \u03b4x)/2 \u2212 r and sgn(x) = 1 for x > 0 or sgn(x) = \u22121 for x < 0. Equations (2) and (3) have been derived geometrically from Fig. 1. The maximum normal strain is usually lower than 10 per cent, which, according to a study of the present authors in references [41] and [45], justifies the use of the theory of linear elasticity. Thus, the contact stresses of the seal are \u03c3x \u223c= ( \u03b5x + \u03b5y + \u03b5z ) \u03bb + 2G\u03b5x \u2212 \u03b1 E \u03b8 1 \u2212 2\u03bd \u03c3y \u223c= ( \u03b5x + \u03b5y + \u03b5z ) \u03bb + 2G\u03b5y \u2212 \u03b1 E \u03b8 1 \u2212 2\u03bd \u03c3z \u223c= ( \u03b5x + \u03b5y + \u03b5z ) \u03bb + 2G\u03b5z \u2212 \u03b1 E \u03b8 1 \u2212 2\u03bd \u23ab\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23ac \u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23ad (4) where \u03bb \u2261 \u03bdE(\u03b80)/[(1 + \u03bd)(1 \u2212 2\u03bd)], \u03bd, E , and \u03b1 are the Poisson\u2019s ratio, elastic modulus, and thermal expansion coefficient of the seal, respectively; \u03b8 = \u03b8 \u2212 \u03b80 where \u03b80 is the temperature of seal installation in its housing and \u03b8 is the operating temperature; E \u2261 [E(\u03b80) + E(\u03b8)]/2 is the average elastic modulus of the seal between the temperatures of installation and operation; G = E/(2 + 2\u03bd) is the shear modulus of the seal, and \u03b5\u0304z = 1 W \u222bW /2 \u2212W /2 \u03b5z dx (5) is the average normal z-strain with W being the contact width (normally, W \u223c= bx \u2212 \u03b4x)",
            " These are again derived from the condition of contact pressure continuity at the elastomer\u2013PTFE boundaries x = x0, that is [p(rubber) y = p(PTFE) y ]x=x0 , which gives \u03b4(PTFE) x = C + \u03bbrubber ( \u03b4x/b(rubber) x ) ( 2\u03bbrubber/b(rubber) x ) + ( \u03bbPTFE/b(PTFE) x ) , \u03b4(rubber) x = \u03b4x \u2212 2\u03b4(PTFE) x (13) where C \u2261 (\u03bbPTFE \u2212 \u03bbrubber) ( \u03b5y + \u03b5z ) + 2 [GPTFE (\u03b80) \u2212 Grubber (\u03b80)] \u03b5y + ( \u03b1rubber Erubber 1 \u2212 2\u03bdrubber \u2212 \u03b1PTFE EPTFE 1 \u2212 2\u03bdPTFE ) \u03b8 (14) EHL analysis of the horizontal sections for steady-state conditions has been presented in reference [47] for elastomeric materials and in reference [48] for composite (PTFE\u2013elastomer\u2013PTFE) materials. The analysis is based on the one-dimensional Reynolds equation because by bx and corner seals (Fig. 1) are used to prevent side leakage. The lubrication (Reynolds) equation to be solved is d dx ( \u03c1h3 \u03b7 dp dx \u2212 6V \u03c1h ) = 0 (15) where h is the local film thickness, V is the tangential velocity of the moving surface in the sealing contact, \u03c1 and \u03b7 are the local fluid mass density and dynamic viscosity, respectively, calculated from the following semi-empirical equations \u03c1 = ( 1 + c\u03b1p 1 + cbp ) \u03c10, \u03b7 = \u03b70e\u03b1Bp (16) where \u03c10 = \u03c1(p = 0), \u03b70 = \u03b7(p = 0), c\u03b1 and cb are constants of the sealed fluid, and \u03b1B = 34",
            "5 GHz personal computer for the complete evaluation of the pair of stator and rotor seals in steady-state conditions, including the solution of the EHL problem and the computation of leakage, friction, and extrusion, based on the equations of sections 2 and 3. Some representative examples for steady-state and transient conditions are presented in this section, together with results from a parametric study and an extensive optimization study to establish critical parameters, conditions, and optimum values for maximum sealing performance. Lack of space necessitates presentation of only a small number of interesting results from the vast amount of those obtained in reference [44]. Data used here are based on a real application of an RVA (Fig. 1) for the control of aircraft wing surfaces [44] but dimensions and conditions have been altered to avoid confidentiality concerns. The vane seals are composite as in Fig. 2. The data are as follows: (bx , by , bz) = (4, 80, 5) mm for the horizontal sections and (4, 5, 25) mm for the vertical sections; b(rubber) x = 1.4 mm, b(PTFE) x = 1.3 mm (Fig. 2); r = 0.1 mm (Fig. 1) (this is mainly used to account for some wear after a number of operation cycles); (\u03b4x , \u03b4y , \u03b4z) = (500, 0, 100) \u03bcm (except in Fig. 6), where \u03b4z = \u03b4r = \u03b4s; \u03b80 = 22 \u25e6C; \u03b8 = \u221254 \u25e6C (this is the operating temperature at normal aircraft altitude); Rr = 40 mm, Rs = 65 mm (Fig. 1); zc = 0.1 mm (except in Fig. 7); \u03c9 = 0.7 rad/s (except in Figs 5 and 7); pcyl = 30 MPa (except in Fig. 7). Sealed fluid: Skydrol 5 aviation hydraulic fluid [51]; \u03c10 = 1031 kg/m3 at \u221254 \u25e6C; \u03b70 = 2.6288 Pa s at \u221254 \u25e6C; c\u03b1 = 0.6 GPa\u22121, cb = 1.7 GPa\u22121 (density-versuspressure data were not available for Skydrol 5 and the values of the coefficients used here are for typical mineral oils); \u03b1B = 39.0 GPa\u22121 at \u221254 \u25e6C (using Wooster\u2019s formula, quoted in section 2.3). As already explained, the composite vane seals comprise two materials, an elastomer and a form Proc",
            " As explained in section 4.1, the application presented here concerns an RVA for the control of aircraft wing surfaces and the operating temperature at normal altitude is about \u221254 \u25e6C. Figure 4 shows the contact pressure and film thickness on the horizontal sections of both RVA seals, followed by the film thickness and leakage along the vertical sections of the rotor seal. The contact pressure distribution on the horizontal sections is easily visualized by looking at the seals in the lower drawings of Fig. 1 and accounting for the different elastic moduli and Poisson\u2019s ratios of the composite seal parts: the central, elastomeric part is almost incompressible (Poisson\u2019s ratio equal to 0.499) and the two outer, PTFE parts are compressible. Following this, the corresponding film thickness distribution is easy to understand, taking into account the rotation of the rotor, which, for Fig. 4, is clockwise. The average film thickness for the horizontal section of the stator seal is lower than that of the rotor seal owing to the lower peripheral velocity of the former (Fig. 1). Both average film thicknesses are in the order of 0.6 \u03bcm, which is realistic for the given conditions and viscous (at operating temperature) hydraulic fluid. Average film thickness along a circular arc and corresponding leakage for the vertical sections of the rotor seal are shown in the lowest diagram in Fig. 4 as functions of the distance from the rotor axis. Naturally, these vary linearly with the distance from the rotor axis owing to the variation of the local peripheral velocity of the related contact",
            " In view of these observations, the optimum is \u03b4y = 0. Interference \u03b4z has insignificant effect at higher temperatures but it has a substantial effect on friction and extrusion at low sub-zero temperatures (Fig. 6). Leakage is not significantly decreased by increasing this interference. Therefore, the focus is on friction and extrusion, both of which rise markedly with \u03b4z , especially for \u03b4z > 200 \u03bcm. In view of these observations, the optimum is 50 < \u03b4z < 100 \u03bcm. Composite vane seals used in, for example, rotary vane actuators (Fig. 1) are machine elements of complex mechanical behaviour. They have to perform under transient operating conditions between temperature and pressure extrema. The present study dealt with the modelling of their mechanics and elastohydrodynamics, with emphasis on computational speed but without sacrificing precision. The mathematical analysis was kept as simple as possible. The sealing performance of vane seals was evaluated on the grounds of mass leakage rate, hydrodynamic friction force, and extrusion size during operation",
            " 2) c\u03b1, cb coefficients in the density\u2013pressure function C auxiliary variable defined in equation (14) E elastic modulus E [E(\u03b80) + E(\u03b8)]/2 EPTFE [EPTFE(\u03b80) + EPTFE(\u03b8)]/2 Erubber [Erubber(\u03b80) + EPTFE(\u03b8)]/2 F hydrodynamic friction force (equations (27) and (32)) G E/(2 + 2\u03bd); shear modulus G\u0304 weighted average shear modulus of the composite seals (equation (30)) GPTFE, Grubber shear modulus of the PTFE and the elastomer h local film thickness H \u03c1h hm value of h at an extremum point of q (equation (20)) Hm value of H at an extremum point of q (equation (18)) H\u03b1 H (x = x\u03b1) (equation (25)) K1, K2, K3 auxiliary coefficients \u2013 equation (9) p local contact pressure pcyl sealed pressure pin elastohydrodynamic inlet pressure py contact pressure on seal vertical sections (equation 12) q auxiliary variable, defined implicitly in equation (34) Q mass leakage rate (equations (26) and (31)) r seal corner radius in the x-direction R radial distance from the rotor axis Rr inner radius of the stator (Fig. 1) Rs outer radius of the rotor (Fig. 1) S length of the seal extruded part (Fig. 3 and equations (28) and (33)) t time U auxiliary variable (U \u2261 V for the hor- izontal seal sections and U \u2261 R\u03c9 for the vertical sections) V contact sliding velocity W contact width; W \u223c= bx \u2212 \u03b4x xin local coordinate in the EHL inlet zone x(\u03b1) in distance between the inflexion point of q and the nearest edge of the \u2018dry\u2019 contact zone xr [(bx \u2212 \u03b4x)/2] \u2212 r x\u03b1 distance of the inflexion point of q from point x = W /2 x0 position of an elastomeric\u2013PTFE boundary zc stator\u2013rotor clearance (Fig. 3) zr length (Fig. 1 and equation (3)) zrot rotor\u2013seal local interference (Fig. 1 and equation (2)) zsr auxiliary variable defined in equation (2) Proc. IMechE Vol. 221 Part J: J. Engineering Tribology JET290 \u00a9 IMechE 2007 at Purdue University on June 5, 2015pij.sagepub.comDownloaded from zsta stator\u2013seal local interference (Fig. 1 and equation (2)) \u03b1 thermal expansion coefficient \u03b1B pressure\u2013viscosity coefficient \u03b1PTFE, \u03b1rubber thermal expansion coefficients of the PTFE and elastomer \u03b4r, \u03b4s rotor and stator seal z-interferences (Fig. 1) \u03b4x , \u03b4y , \u03b4z principal seal interferences at installation \u03b4(rubber) x , \u03b4(PTFE) x compressions of the elastomeric and PTFE parts of the composite seal (Fig. 2) \u03b8 \u03b8 \u2212 \u03b80 \u03b5x , \u03b5y , \u03b5z seal normal strains \u03b5\u0304z average normal z-strain (equation (5)) \u03b7; \u03b70 dynamic viscosity (equation (16)); \u03b70 = \u03b7(p = 0) \u03b8 , \u03b80 operating temperature and temperature of seal installation \u03bb \u03bdE(\u03b80)/[(1 + \u03bd)(1 \u2212 2\u03bd)] \u03bbPTFE, \u03bbrubber \u03bb for the PTFE and elastomer \u03bd Poisson\u2019s ratio \u03bdrubber, \u03bdPTFE Poisson\u2019s ratios of the elastomer and PTFE \u03c1; \u03c10 mass density (equation (16)); \u03c10 = \u03c1(p = 0) \u03c1\u03b1 \u03c1(x = x\u03b1) \u03c3\u0304x , \u03c3\u0304y , \u03c3z seal normal stresses (equation (4)) \u03c4\u0304zx average viscous shear stress in a sealing contact (Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002856_ac201072u-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002856_ac201072u-Figure1-1.png",
        "caption": "Figure 1. Diagram showing the construction of the nanofiber junction reference electrode (NFJRE).",
        "texts": [
            " The impedance spectra were generated using a three-electrode systemwith a Ag/AgCl (3MKCl) half cell as reference, platinum disk counter electrode, and an Autolab PGSTAT 30 potentiostat/galvanostat equipped with a frequency response analyzer module (Eco Chemie, Netherlands). Resistance was measured by scanning current between(1 pA and recording the resulting potential using a Keithley model 6430 electrometer with the NFJRE as a working electrode and a platinum wire as the reference/counter. Electrode Construction. A schematic of a NFJRE is shown in Figure 1. Electrode bodies were assembled by heating a 1-mmdiameter Ag wire (AlphaAesar, 99.999% purity) to approximately 900 C and dipping into molten AgCl for several seconds until a 2 mm pellet formed at the tip of the wire to act as the internal reference element. This Ag/AgCl pellet was then epoxied into an in-house fabricated PVC electrode body. The nanocomposite membrane was prepared by dissolving 9 parts PMMA and 1 part CNF in m-xylene. The mixture was sonicated for 10 min before applying to the electrode body in 30 \u03bcL aliquots"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002984_00368791311303474-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002984_00368791311303474-Figure2-1.png",
        "caption": "Figure 2 Worn bearing geometry",
        "texts": [
            " The fluid film journal bearing operating in turbulent regime can be modeled by the following equation as (Awasthi et al., 2006): \u203a \u203aa h3 Ga \u203a p \u203aa \u00fe \u203a \u203ab h3 Gb \u203a p \u203ab \u00bc V 2 \u203a h \u203aa \u00fe \u203a h \u203at \u00f01\u00de Ga and Gb are the turbulence coefficients which are the functions of local Reynolds number and the expression is given as (Constantinescu, 1959; Constantinescu and Galetuse, 1965; Anjani et al., 1996): Ga \u00bc 12 \u00fe 0:026\u00f0Re\u00de0:8265 \u00f02\u00de Gb \u00bc 12 \u00fe 0:0198\u00f0Re\u00de0:741 \u00f03\u00de For the laminar flow, Ga \u00bc Gb \u00bc 12 and local Reynolds number Re \u00bc 0. The geometry of the worn zone is shown schematically in Figure 2 given as (Dufrane et al., 1983) and the expression for non-dimensional nominal fluid-film thickness h in a worn misaligned hole-entry hybrid journal bearing system may be expressed as (Jain et al., 1997): h \u00bc 1 2 Xjcosa2 Zjsina\u00fe d:b:cosa2 s:b:sina\u00fe \u203a h \u00f04\u00de where, Xj and Zj are the equilibrium co-ordinates of the journal centre, s and d are the misalignment parameters and: \u203a h \u00bc dw 2 1 2 sina for ab # a # ae \u00f05\u00de \u203a h \u00bc 0 for a , ab or a . ae \u00f06\u00de The angles ab and ae are calculated by equating equation (5) with equation (6) as: sina \u00bc dw 2 1 \u00f07\u00de The flow of lubricant through an orifice and capillary restrictors are expressed in non-dimensional form as (Awasthi et al"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001242_s12239-009-0050-0-Figure14-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001242_s12239-009-0050-0-Figure14-1.png",
        "caption": "Figure 14. Rear edge swing-out.",
        "texts": [
            " Therefore, the transition function was derived based on the response of the first-order system to a step-function input shown in equation (13). Figure 13 shows the relation between the pilot angle and the steered angle for the limit values of the steering angles, 5o and 2o. If (12) If (13) (14) where \u03b5 is the permissible value, \u03b8 is the steering angle, \u03b80 is the limit value of the steering angle, and Pmax is the maximum value of the virtual rigid axle. 3.4. Suppression of Rear Swing-out The rear of the vehicle generally overhangs the rear axle. As a result, the rear of a vehicle swings to the outside of the rear axle, as shown in Figure 14. In the case of a vehicle with AWS, the swing-out increases because of the rear steering at a reverse phase angle. To prevent rear swingout, the angles of the rear wheel were suppressed for a constant distance and then the values of the virtual rigid axle were moved to maximum setting values, as shown in equation (15). (15) (16) where \u03c9 is the angular velocity of the steering axle and \u03b4max is the maximum steering angle. 3.5 Steering Angle according to Vehicle Speed When a vehicle travels around a curve, speed is restricted for safety"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003713_978-1-4614-2419-2-Figure27.6-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003713_978-1-4614-2419-2-Figure27.6-1.png",
        "caption": "Fig. 27.6 Full field displacement with iso-lines of CX-100 due to a 350 lb load applied at 6.75 m from root",
        "texts": [
            " These separate fields of view were combined, with the use of reference points, to measure full field blade deformations. Figure 27.5 shows the measured surface, in each state, overlaid on top of each other. The stitching method was applied within the DIC software package S-View in order to calculate direct displacement and strain for the blade surface under loading. The missing data on the loaded blade surface is due to the DIC pattern being covered by the loading apparatus. The full field displacement from unloaded to loaded states is shown in Fig. 27.6. The curvature of the centerline of the blade with respect to span location measured from the root along the X-axis is shown in Fig. 27.7. A displacement of 146 mm at the blade tip was measured due to the 350 lb load applied at the 7 m span location. The curvature from one measured field of view to the next is continuous showing no visible degradation of accuracy due to stitching. 280 B. LeBlanc et al. The span wise strain overlaid on the CX-100 surface for the same loading condition is shown in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003509_j.proeng.2013.08.203-Figure5-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003509_j.proeng.2013.08.203-Figure5-1.png",
        "caption": "Fig. 5. Evaluation of the Cartesian stiffness for the ABB IRB6660 (a) and for PKM Tricept (b)",
        "texts": [
            " The Cartesian stiffness Kx can be defined by the relation: (1) The results are really different from one architecture to the other because, to sweep the surface with the robot with a parallelogram closed loop, all the actuator moves but only the joints defined by the actuated motions q2 to q5 are charged. Concerning the value obtained, the stiffness is more homogeneous in its workspace. If the wrist is crooked, the force is distributed on link 4 and 5 and the IRB6660\u00ae is stiffer in this configuration (Fig. 5a). Concerning the parallel robot, the extension of the legs leads to a loss of stiffness. Moreover, it can be observed the same pattern on the edges of the swept area (Fig. 5b). Concerning the manipulator with parallel architecture, the deformation induced by the charge is relatively proportional to the median leg length. The following criterion is introduced (Robin et al. (2011)): (2) dX T K x 3 321 qqq rt Regarding serial manipulators, many studies have considered that static deformations are mainly located in the actuated joints (Dumas et al. (2011)). Nevertheless, as far as our structure is concerned, deformations are also located in the links and in the passive joints (Pashkevitch et al"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001847_1.4000269-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001847_1.4000269-Figure2-1.png",
        "caption": "Fig. 2 Schematic view of test rig",
        "texts": [
            " The upper part of a test bearing is provided with an oil supply groove, which allows supplying lubricating oil from the groove into a bearing clearance. The bearing clearance is set relatively large because the research subject of our experiment is concentrated to small-bore high-speed bearing used in small size lightly loaded, relatively low cost machinery such as small-size compressor or turbine. The relatively large clearance has some advantages in manufacturing assembling and cost etc. In addition, the large clearance makes it possible to get more a clear vision of the cavitations. A schematic view of the test rig is shown in Fig. 2. The test rig consists of a rotor installed in its central part and a revolving shaft supported by two bearings on its left-hand and righthand sides, respectively. A cylindrical journal bearing, which is an object of measurements, is installed on the right-hand side. The JANUARY 2010, Vol. 132 / 011703-110 by ASME of Use: http://www.asme.org/about-asme/terms-of-use s o p v j l w g s o c 0 Downloaded Fr haft is driven by a dc motor with a possibility to vary the number f revolutions continuously up to 10,000 rpm"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002522_mesa.2012.6275544-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002522_mesa.2012.6275544-Figure2-1.png",
        "caption": "Figure 2. The model parameters for an -legged robot.",
        "texts": [
            " This reduces the calculation costs of the MFFSM to only one axis. Also, definitions of and \u210e are different. The vector for a spatial robot is the vector created by connecting the CG perpendicularly to the tipover axis and is referred to as the tipover axis normal vector. For a spatial robot, \u210e is defined as the height of the CG with respect to the tipover axis and is aligned with the gravity vector. Based on the above definitions, and \u210e are calculated mathematically. Consider an -legged robot as shown in Fig. 2. In Fig. 2, & is th tipover axis and &' = & /\u2016& \u2016 is its unit vector. The support polygon is not restricted to be planar because the robot is assumed to be on uneven terrain, which is another advantage of the presented criterion. The following equations show how and \u210e are calculated and substituted into (2) for stability analysis. & = * \u2212 (3) where = 1, \u2026 , \u2212 1. When = , then * will turn to . From Fig. 2, is that portion of the vector + * \u2212 -which is perpendicular to the tipover axis. This indicates that can be obtained by subtracting that portion of the vector * \u2212 that is along the tipover axis as follows. = + * \u2212 - \u2212 .+ * \u2212 - \u2219 &' 5&' (4) Finally, = \u2016 \u2016. The height of the CG with respect to the tipover axis is the other term which is needed to be calculated and substituted in (2). The height is the vertical portion of the tipover axis normal vector which is given by \u210e = \u2219 +\u221267- (5) where 67 is the unit vector of 9-axis of the global coordinate system"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003196_j.proeng.2011.04.140-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003196_j.proeng.2011.04.140-Figure1-1.png",
        "caption": "Fig. 1. The photograph of battery bracket in vehicle trunk Fig. 2. The location of accelerometer",
        "texts": [
            " However, currently there is no information related to the acceleration level and testing time. Although the up-to-now study results and specifications relevant to the vibration-related lifecycle all show their own distinct features, they are still confined to a limited scope. As such, this study introduces the single-DOF acceleration test method that is suitable for battery fixing brackets. The battery bracket module(40kg) is installed in the vehicle trunk and accelerometers are attached symmetrically on the module fixed on the trunk bottom. (Fig. 1 and Fig. 2) Vibrations were measured while driving on expressways, unpaved roads, national highways, local roads and city streets. The vibration on the battery fixing bracket was measured for each type of road using the accelerometer. Table 1 shows the standard distance which is the result of \u201cTesting method standardization and evaluation technique for automobile parts\u201d[8]. Table 1 also shows actual driving distance(mileage), driving time and expansion factor. Mileage includes time for breaks and stops"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002167_cca.2009.5280933-Figure8-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002167_cca.2009.5280933-Figure8-1.png",
        "caption": "Fig. 8. Drive acceleration method implementation scheme.",
        "texts": [
            "0 m/s2 has a little effect on gap closing time and it has to be chosen by the electric motors limits. First zone of drive regulation is assumed (below nominal speed). Backlash gap variation in the range 0.01-0.03 rad for the different spindles wear and angular position also has a little effect on closing time for more than 0.6 m/s2 acceleration level. Lower acceleration limit is restricted by the drive control system (about 0.1 m/s2 referred to the work rolls). An implementation scheme of drive acceleration method is represented in Fig. 8 for nonreversing rolling stand with one edger. Slabs are being feeding to the rolling stand by the individually driven rollers. A vertical edger begins rolling side edges and then horizontal rolling stand reduces strip thickness. Hydraulic scale removing is provided at the exit of stand. Time interval between successive slabs usually is about 30-40 s. A rolling speed is in the range 1-4 m/s for different stands. Drive dynamic torque should be of enough level to close backlash gap and overcome reactions in the drive train: ),max( minijijRijdyn TTT >\u22c5 (2) where Tdyn ij = Jij a - dynamic torque; Jij - summary inertia of drive chain from the rolls to Cij coupling; a - rotation acceleration referred to Cij coupling; TR ij - reaction torque in coupling which is required to control; Tmin ij - minimal dynamic torque after first period of oscillation"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002930_978-3-642-31988-4_26-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002930_978-3-642-31988-4_26-Figure1-1.png",
        "caption": "Fig. 1 Conceptual model of the FAST",
        "texts": [
            "hina is building a Five-hundred meter Aperture Spherical radio Telescope (FAST) [1]. Figure 1 shows the conceptualization of the FAST system, where the feed support system moves over the active reflector. The feed support system of FAST includes two parts: first-level adjustable feed support system, which is a six-cable driven parallel manipulator with large span that provides the coarse positioning, and a secondlevel adjustable feed support system (A\u2013B rotator and a Stewart platform) that can compensate the positioning error and achieve the required accuracy. At present, a R. Yao (B) \u00b7 H"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000880_tmag.2007.891399-Figure5-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000880_tmag.2007.891399-Figure5-1.png",
        "caption": "Fig. 5. Distributions of eddy current density vectors: (a) t = 0:125 ms; (b) t = 4:25 ms.",
        "texts": [
            " In this calculation, it is supposed that the ring magnet is not inserted into the gap between the conductors at the initial condition in this impact drive mechanism. Therfore, the initial potentials of all edge elements are set to be zero in order to simulate under the conditions of that the ring magnet is rotated at constant speed and suddenly inserted into the gap between the conductors, and the ring magnet is rotated from the rotation angle of 45 . Figs. 5 and 6 show the distributions of eddy current and Lorentz force density vectors, respectively, when the ring magnet is rotated at the rotation speed of 2000 rpm. From Fig. 5(a) and (b), the distributions of eddy current density vectors are amazingly different because the flux density vectors are suddenly occurred at the initial position. For that reason, the distributions of Lorentz force density vectors are also different as shown in Fig. 6(a) and (b). Fig. 7 shows the calculated time variations of transient torque when the ring magnet is rotated at the rotation speed of 2000 rpm. The peak value of torque is about 0.8 at the position where the boundary of ring magnet passes through the center of conductors"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000154_iros.2007.4399278-Figure3-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000154_iros.2007.4399278-Figure3-1.png",
        "caption": "Fig. 3. Schematic of Variable Pitch Propeller (VPP) System",
        "texts": [
            " The second term in the right side of equation (14) represents an aerodynamic moment produced by the change of the roll rate, normally opposing to the roll moment, that is why, the derivative, C\u2113\u03c6\u0307 = 0.36, is known as roll damping derivative. Then equation (13) can be rewritten as follows: \u03c6\u0308 = (F \u00b7 d\u2212 C\u2113\u03c6\u0307 \u03c6\u0307)/Jx (15) The lift force in each rotor can be considered as the thrust and can be calculated by the following expression: T = Ct\u03c1n 2D4 (16) where Ct is the thrust coefficient, \u03c1 is the density of the air, n is the number of revolutions per second of the motor and D is the diameter of the propellers. The thrust coefficient is a function of the pitch angle propeller \u03d5, which is shown in Figure 3. The thrust coefficient in a linear region can be calculated by: Ct = Ct\u03d5\u03d5 (17) where Ct\u03d5 is a derivative which represents the thrust slope with respect to the VPP angle. This derivative has been estimated using a shareware program called JavaProp [7]. This program uses the number of blades, the velocity of rotation, the diameter of the propellers, the velocity and the power of the motor to give the value of Ct for an operational range 5\u25e6 \u2264 \u03d5 \u2264 15\u25e6 as is shown in Figure 4. Then, using MatlabTM a first order polynomial (dashed line) can be constructed using the values of the thrust coefficient for each \u03d5 angle. The dashed line slope is the derivative of this polynomial which in fact represents the derivative Ct\u03d5 , and its value is estimated to be 0.0025. Then using the inertia values given in Table I and applying the Laplace transform, the following transfer function for the roll angle with respect to the VPP angle displacement is obtained. \u03c6(s) \u03d5(s) = 5 s2 + 25s (18) Now, the VPP dynamics will be determined. In the Figure 3, it can be seen that the aerodynamic pitch moment of the blades must be equal to the moment generated by the servo mechanism. Considering that the blade profile corresponds to the NACA0014, then the following approximation can be used to obtain the blade pitch moment: mb = \u03c1V 2 tb Sbcb 2Jyb [ Cm\u03d5 \u03d5+ Cm\u03d5\u0307 \u03d5\u0307 ] = ksfs\u03b4s (19) where the subscript b denotes the blade. The term Vtb denotes the total velocity of the propeller at the tip, it is given by: Vtb = \u221a v2 axial + v2 rotational (20) where vrotational = \u03c0nD and vaxial is the aircraft velocity"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001343_aim.2009.5229984-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001343_aim.2009.5229984-Figure2-1.png",
        "caption": "Figure 2. DH coordinates and the position of grasp points with respect to C.G. of object.",
        "texts": [
            " The MAG Index could be used to evaluate grasp configuration in object manipulation tasks that are consisting of static, kinematics, and kinetics aspects. Also with the MAG Index, the cooperative object manipulation tasks could be evaluated. First consider the grasp matrix for n-contact points that has a structure as follows, [19], 1 3 3 3 3 3 3 3 3 6 6n 1 0 ... 1 0 = (1) ... j T T T T obj p obj obj p obj G S r S S r S \u00d7 \u00d7 \u00d7 \u00d7 \u00d7 \u00d7 \u00d7 where rpj is the position vector of the j-th EE, with respect to the coordinate system placed in the object center of mass and parallel to principal axes of object, Fig.2, and rx pj is the cross operator of vector rpj. The matrix objS is obtained by: 1 0 0 (2) 0 obj S S C C S S C C \u03b3 \u03d5 \u03b3 \u03d5 \u03d5 \u03b3 \u03d5 \u2212 = \u2212 where \u03b3 and \u03d5 are pitch and roll angels. Suppose that, NC , denotes the inverse condition number of the grasp matrix G : ( ) ( )min max/ (3)NC G G\u03c3 \u03c3= where min\u03c3 and max\u03c3 are the smallest and the largest singular values of G respectively. If NC is close to zero, the grasp is singular. This means that the End-Effector of at least one manipulator in system is incapable of moving the object in an arbitrary direction",
            " The shape the object and position of its center of mass is shown in fig 8. We choose some candidate points on the grasped object where we calculate of the MAG index for them. In practice, these points are located using a vision system. Coordinates of candidate grasping points on the object are presented in TableII. The object path is the straight line along X-axis. Joints trajectory are quintic functions as follows, ( ) ( ) 2 3 4 5 0 1 2 3 4 5 ( ) 1 (8) 10 X t a a t a t a t a t a t Y t t\u03c8 = + + + + + = = \u00b0 with coefficients that presented in [18]. Fig. 2 shows trajectory of ( )x t and its derivatives. For calculating the terms of (7), we consider the velocity of the grasped object from two ways. First, with the transformation of the actuators velocity, and then with the task predefined trajectory. Equality of these relations gives us configuration of the manipulator to reach defined task.The velocity of the contact points computes from the velocity of the actuators as follows, [ ] [ ]6 1 6 3 3 1 (9)cV J \u03b8 \u00d7 \u00d7 \u00d7 = where cV denotes the velocity array of the contact points on the grasped object as follows, 6 1 (10) T c c c c x y zV x y z \u00d7 = \u03a9 \u03a9 \u03a9 cx , cy and cz are the linear components and x\u03a9 , y\u03a9 and z\u03a9 are the angular components of the contact point velocity array"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003903_tdei.2015.005053-Figure10-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003903_tdei.2015.005053-Figure10-1.png",
        "caption": "Figure 10. Preparation of the sample insulator.",
        "texts": [
            "0 kV and the conductivity and spray velocity are 2-3 mS/cm and 2-3 L/h respectively. During the test process, NaCl solution is sprayed to the surface of insulator to form electrolyte. The experimental set-up is shown in Figure 9. Before the test, the copper electrode should be pasted and fixed onto the lower surface of the insulator. The distance between the electrode and the porcelain is about 1 cm. The metal wire connected to the other end of the copper electrode shall be fixed on the locking device, as shown in Figure10. 5.2 ANALYSIS OF TEST RESULTS XZP2-300 type porcelain insulators installed different organic sleeve are carried out simulation test by means of spray water method. The test electric charge quantity is set as 45000 C, according to the maximum average annual amount of corrosion charge (1500 C/year) obtained from Chusui transmission line and the 30-year service life of porcelain insulator. After tests, 1000 kg of cement block has been hung on these insulators for six months, as shown in Figure 11. Six months later, the tested insulators are carried out tensile load test in accordance with Chinese National Standard GB/T19443 [23]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000193_icma.2005.1626722-Figure10-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000193_icma.2005.1626722-Figure10-1.png",
        "caption": "Figure 10 Transformation lines from one camera, and the subsequent shadowing effect due to the clad\u2019s cross-section.",
        "texts": [
            " As mentioned at the beginning of this sub-section, the projection of melt pool boundaries into the vertical work plane provides a very good estimate of the clad\u2019s cross-section. This information can therefore be utilized to estimate the rate of solidification of the clad. Since the melt pool represents melted metal, the interface of the melt pool and the previously deposited cald provides a solid-liquid interface, indicative of the solidification rate. With two images from independent camera angles, a projection into the horizontal work plane (essentially the substrate\u2019s plane) can reveal the width of the clad being deposited. Figure 10 illustrates the transformation lines for one camera\u2019s image plane to the work plane (substrate). It can be seen that there is a shadowing effect, due to the clad\u2019s shape. Thus, this camera can only see one true edge of the clad. With a second camera on the opposite side, the opposing true edge of the clad can be viewed. The combination of the two cameras produces the clad\u2019s width. Figure 11 shows a typical images projecting on horizontal work plane To evaluate the performance of the aforementioned algorithms, a test setup shown in Figure 12 was created"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003903_tdei.2015.005053-Figure6-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003903_tdei.2015.005053-Figure6-1.png",
        "caption": "Figure 6. Calculation results of strain.",
        "texts": [
            "3 Hz, and peak as 20 kN to simulate the dynamic load under iced conductor galloping condition. 4 20sin(2 0.3 )kNT t   (18) Take the first insulator counted from cross arm as the target, its axial tensile force is shown in Equation (19). 1 2 3 4 72.58 20sin(0.6 )kNT T T T T t      (19) In order to ensure the calculation accuracy, 3D models are used. The Young's Elasticity modulus and Poisson's ratio of each element are listed in Table 2.The calculation results at the time of maximum mechanical load are shown in Figure 6. Figure 6 indicates that, the pin withstands large tensile load during the process of iced conductor galloping. Especially, the bottom half part exposed to the air and the annular section of cement-zinc sleeve interface. These two areas have high strain and are easier to be damaged, as seen in Figure 6b. If the cross-section of pin reduces due to corrosion, the exposed part is easier to be tension fractured. If the bonding strength between cement and zinc sleeve has been reduced due to expansion of corrosion by-products, it can result in the pulling out of pin form cement. The failure forms of pin corroded insulators retrieved from \u00b1800 kV Chusui line in March, 2014 are all pins pulled off from cement, as shown in Figures 5c and 5d, the cross section of pins have not been reduced since for the protective zinc sleeve, but the bonding strength between cement and zinc sleeve has been reduced due to the expansion of corrosion by-products. From this point of view, the calculation result illustrated in Figure 6b is consistent with the failure forms of mechanical load test that are shown in Figures 5c and 5d. According to the previous research results, the mechanical strength of the operating insulators decreases even their pins are not corroded and become thin. The expansion caused by corrosion by-products reduces the bonding strength between the cement and zinc sleeve. Thickening the zinc sleeve is no use for this condition. Therefore, the recommendation that installing organic material sleeve onto the zinc sleeve is proposed, as shown in Figure 7"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003954_213548-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003954_213548-Figure2-1.png",
        "caption": "Figure 2: Basic structure of HJB (the dimension of the clearance is magnified).",
        "texts": [
            " In this work, an ANN is used and is trained with the solutions of the Reynolds equation (RE). The RE is valid for finite-length HJB (0.25 \u2264 \ud835\udf06 \u2264 4) and is valid for a range of eccentricities between 0.1 and 0.9. The general form of RE is [5] \ud835\udf15 \ud835\udf15\ud835\udf03 (\u210e 3 \ud835\udf15\ud835\udc5d \ud835\udf15\ud835\udf03 ) + ( \ud835\udc37 2 ) 2 \ud835\udf15 \ud835\udf15\ud835\udc66  (\u210e 3 \ud835\udf15\ud835\udc5d \ud835\udf15\ud835\udc66  ) = 6\ud835\udf07\ud835\udc62 \ud835\udc37 2 \ud835\udf15\u210e \ud835\udf15\ud835\udf03 + 6\ud835\udf07V( \ud835\udc37 2 ) 2 \ud835\udf15\u210e \ud835\udf15\ud835\udc66  . (3) In this form, \ud835\udc5d is the pressure,\ud835\udc37 the diameter of the journal, \u210e the film thickness, \ud835\udc62 the radial velocity, V the axial velocity (which is zero for this work), \ud835\udf03 the angular coordinate, and \ud835\udc66 the axial coordinate visible in Figure 2. For this solution, the finite difference techniquewas employed. If themisalignment of the rotational axis on the plane \ud835\udc67 -\ud835\udc66 is considerable (>0.005\u2218), the film thickness formula is \u210e = \ud835\udc36 + \ud835\udf00\ud835\udc36 cos (\ud835\udf03 \u2212 \ud835\udf03 0 ) + tan (\ud835\udefe) (\ud835\udc66 \u2212 \ud835\udc3f 2 ) cos (\ud835\udf03 \u2212 \ud835\udf03 0 ) , (4) where \ud835\udc36 is the clearance, \ud835\udc3f is the length of the bearing, \ud835\udf03 0 is the attitude angle, and \ud835\udefe is themisalignment angle that can be evaluated as follows (angle between the \ud835\udc66  axis and plane \ud835\udc65 -\ud835\udc66): \ud835\udefe = \ud835\udc39\ud835\udc3f 2 \ud835\udc53 16\ud835\udc38\ud835\udc3d . (5) Above, \ud835\udc39 is the force acting on the bearing, \ud835\udc3f \ud835\udc53 is the distance between the application point of the force, and the plane by guest on June 23, 2016ade",
            " The last reason why 10\u2218C is chosen instead of 20\u2218C is that a large amount of lubricants, without additives, suffer from rapid degradation due to temperature rise as explained in [2] specially if higher than 10\u2218C. (iii) Maximum pressure: the choice of 28MPa has been made to impose a great constrain on pressure, because by guest on June 23, 2016ade.sagepub.comDownloaded from The minimum supply diameter is 4mm; the maximum is 9mm. The inlet temperature is 25\u2218C. The hole\u2019s position can be any suitable angular direction in the upper housing comprised between \u221265\u2218 and 65\u2218 (according to the scheme in Figure 2). The optimization parameters for the two techniques adopted are listed in Table 2. The results of the optimization process are shown in Table 3. Here, the results obtained with different design techniques are listed. The design generated with Ghorbanian technique saves 39.8% power loss and over 50% of the mass flow. On the other hand, the combination of deciding variables generates a bearing with a minimum film thickness of 1.8 \ud835\udf07m, a maximum pressure of 44MPa, and a maximum temperature of 13.1\u2218C"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001731_s10008-009-0899-x-Figure8-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001731_s10008-009-0899-x-Figure8-1.png",
        "caption": "Fig. 8 a, b Scheme for the description of the solid-state electrochemical reduction of an organic dye via concerted proton- and electron-transfer processes in two successive steps",
        "texts": [
            " Taking into account the foregoing set of considerations, a tentative scheme for electrochemical reduction of alizarin, in principle able to be extended to similar dyes, could involve (a) coupled insertion of protons and electrons with formation of a layer of reduced molecules with no changes in the crystal structure, (b) advance of the above layer with formation of a new crystalline phase, and (c) depending on the proton permeability of the new phase, blocking of the reduction process or subsequent advance of the new phase and the preceding proton diffusion layer along the crystal. A tentative pictorial representation is depicted in Fig. 8. ATR-FTIR experiments act in support of the idea that a layer of new compounds is formed during electrochemical oxidation or reduction of alizarin microparticulate deposits. Thus, Fig. 9 compares the spectra for alizarin-modified FTO electrodes (a) before and (b, c) after application of a constant potential step of (b) +1.0 and (c) \u22121.0 V during 30 min in contact with potassium phosphate buffer. Following recent band assignments [65, 66], the spectrum of alizarin exhibits strong bands at 1,258, 1,296 (\u03bd(CO)/ \u03bd(CC)/\u03b4(CCC)), 1,452, and 1,478 cm\u22121 (\u03bd(CC)/\u03b4(COH)/ \u03b4(CCC)) accompanying the carbonyl stretch band at 1,633 cm\u22121"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000586_978-1-4020-8889-6_6-Figure5-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000586_978-1-4020-8889-6_6-Figure5-1.png",
        "caption": "Fig. 5 Load Transportation Device",
        "texts": [],
        "surrounding_texts": [
            "In Fig. 6 two different slung load transportation tasks are shown. The slung load transportation using one helicopter was performed in Utrera (Spain), April 2008 (the experiment was conducted for the first time (by the authors) in Berlin in November 2007). The helicopter transported one liter of water in a jerry can. The multi UAV load transportation was performed in Berlin in December 2007: the three UAVs together transported a load of 4 kg. Figure 7 shows the load transportation using a single helicopter. All coordinates are given in a local Newtonian system N1,2,3. The origin of the frame is defined at the UAV take-off position. The coordinate X3 represents the height of the UAV. Until the load was deployed the coordinates X1,2 show the position of the load and after the deployment the X1,2 coordinates represent the position of the UAV itself. The deployment is indicated by a vertical dotted line. The dotted step is the desired position as is presented to the controller. The colored area shows an internal state of the controller. This internal state represents the input of the controller, which has been filtered to match the flight dynamics of the UAV. The UAV performed an autonomous take-off using a conventional controller; after the load was lifted (at approximate 5 m) the controller designed for the slung load transportation was used. The working height of 15 m was reached in two steps. The load was transported 41 m along n1 and after the desired position had been reached, the height was decreased until the load was deployed from approximately 2 m height. A few seconds before the deployment of the load the conventional controller was reactivated. The transition between the two controllers was not smooth and therefore a movement of approximately 0.5 m can be observed just before the deployment. The conventional controller was used to return the UAV to it\u2019s take-off position and to perform autonomous landing. The performance of the controller was quite good, despite the stormy weather conditions on that day. Even with steady wind of 30 km/h and wind gusts of 40 km/h the controller was able to stabilize the helicopter and damp upcoming oscillation of the load. The load transportation using three small size helicopters is shown in Fig. 8. A load of 4 kg was transported and ropes with a length of 13 m were used. The helicopters were arranged as an equilateral triangle on the ground with a side length of 8 m. Figure 8 shows the coordinates of all three helicopters during the flight. During the experiment the coordinates of the helicopters were recorded in different Newtonian frames N1,2,3, which have the same orientation, but different origins. That\u2019s why there is no static offset between the trajectories of the different helicopters. The steps given to the controller are shown as dotted lines and the colored areas are internal controller states, showing the input steps after a prefilter has been applied, to match the input steps to the translation dynamic of the helicopters. The horizontal dotted line shows the lift-off height of the load. The helicopters performed an autonomous take-off and increased their height to 10 m (with the load still on the ground). Then the height of all three helicopters was increased to 15 m. The load was lifted at approximately 12.4 m. The additional weight of the load was not included in the controller, which leads to a small disturbance in the trajectories along X3 at the moment the load was lifted from the ground. For the same reason disturbances in X3 can be observed during strong acceleration in X1,2-direction. Steps of \u00b110 m along n2 and of +10 m along n1 were performed. The achieved position error during hovering was about \u00b10.3 m. The helicopters only changed their relative position about \u00b10.3 m during flight and the triangular formation of the helicopters was preserved. The load showed almost no oscillation during the whole flight."
        ]
    },
    {
        "image_filename": "designv11_12_0000726_s11044-007-9072-4-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000726_s11044-007-9072-4-Figure1-1.png",
        "caption": "Fig. 1 A hub-beam system with spatial motion",
        "texts": [
            " Based on precise strain-displacement relation, the geometric stiffening effect is taken into account, and the rigid-flexible coupling dynamic equations are derived using velocity variational principle, and finite element method is employed for discretization. Simulation of a hub-beams system is used to investigate the rigid-flexible coupling dynamics of the system. In this section, a dynamic model of an elastic beam is established based on the following assumptions: The beam is homogeneous and isotropic, and the effects due to eccentricity are not considered. The beam has a slender shape so that the shear effect is neglected. As shown in Fig. 1, k is an arbitrary point of a flexible beam undergoing three-dimensional large overall motion, and k0 is the corresponding point on the neutral axis of the beam. Three coordinate systems are introduced to describe the motion of the spatial beam: The inertial frame O0\u2013X0Y0Z0, the body-fixed frame O\u2013XbYbZb of the beam, and the coordinate system k0\u2013X\u2032 bY \u2032 bZ \u2032 b , which is fixed on the neutral axis of the beam after deformation. Let \u03c1 be the relative displacement of the point k0 with respect to O\u2013XbYbZb, \u03b8 and \u03be be the torsion angle of the cross-section and the relative displacement of k with respect to k0\u2013X\u2032 bY \u2032 bZ \u2032 b . The absolute displacement vector r of the point k can be written as r = r0 + \u03c1 + \u03be, (1) where as shown in Fig. 1, r0 is the position vector of the reference point O with respect to O0\u2013X0Y0Z0. Let \u03c9 be the angular velocity of O\u2013XbYbZb with respect to O0\u2013X0Y0Z0, the position vectors r, r0, \u03c1, \u03be and the angular velocity \u03c9 can be written as r = eT 0 r, r0 = eT 0 r0, \u03c1 = eT b \u03c1, \u03be = e\u2032T b \u03be \u2032, \u03c9 = eT b \u03c9, (2) where e0 = [ i0 j0 k0]T , eb = [ ib jb kb]T and e\u2032 b = [ i \u2032 b j \u2032 b k\u2032 b]T are columns of the unit basic vectors of O0\u2013X0Y0Z0, O\u2013XbYbZb and k0\u2013X\u2032 bY \u2032 bZ \u2032 b , respectively. r, r0 are the coordinates of r and r0 defined in O0\u2013X0Y0Z0, respectively, and \u03c1 = [\u03c11 \u03c12 \u03c13]T ,\u03c9 = [\u03c91 \u03c92 \u03c93]T are the coordinates of \u03c1, \u03c9 defined in O\u2013XbYbZb , respectively, and \u03be \u2032 = [0 y z]T is the coordinate of \u03be defined in k0\u2013X\u2032 bY \u2032 bZ \u2032 b . Then one obtains r = r0 + A\u03c1 + A\u2032\u03be \u2032, (3) where A and A\u2032 are the transformation matrices of O\u2013XbYbZb and k0\u2013X\u2032 bY \u2032 bZ \u2032 b with respect to O0\u2013X0Y0Z0. The relation between A\u2032 and A is given by [11] A\u2032 = AT , (4) where T = \u23a1 \u23a3 c\u03bb \u2212s\u03bbs\u03b7 \u2212s\u03bbc\u03b7 s\u03b7s\u03bb c\u03b72 + s\u03b72c\u03bb s\u03b7c\u03b7(c\u03bb \u2212 1) c\u03b7s\u03bb c\u03b7s\u03b7(c\u03bb \u2212 1) s\u03b72 + c\u03b72c\u03bb \u23a4 \u23a6 [1 0 0 0 c\u03b8 \u2212s\u03b8 0 s\u03b8 c\u03b8 ] , (5) and \u03bb is the angle between the tangent of the deformed neutral axis X\u2032 b and Xb, as shown in Fig. 1, which satisfies c\u03bb = cos\u03bb = \u2202\u03c11/\u2202x [(\u2202\u03c11/\u2202x)2 + (\u2202\u03c12/\u2202x)2 + (\u2202\u03c13/\u2202x)2]1/2 , (6) s\u03bb = sin\u03bb = [(\u2202\u03c12/\u2202x)2 + (\u2202\u03c13/\u2202x)2]1/2 [(\u2202\u03c11/\u2202x)2 + (\u2202\u03c12/\u2202x)2 + (\u2202\u03c13/\u2202x)2]1/2 , (7) and \u03b7 satisfies c\u03b7 = cos\u03b7 = \u2202\u03c13/\u2202x [(\u2202\u03c12/\u2202x)2 + (\u2202\u03c13/\u2202x)2]1/2 , (8) s\u03b7 = sin\u03b7 = \u2202\u03c12/\u2202x [(\u2202\u03c12/\u2202x)2 + (\u2202\u03c13/\u2202x)2]1/2 , (9) and c\u03b8 = cos \u03b8, s\u03b8 = sin \u03b8 . Substitute (4) into (3), the position vector is given by r = r0 + A(\u03c1 + \u03be), (10) where \u03be is the coordinate vector of \u03be defined in the body-fixed frame, which reads \u03be = [ \u2212ys\u03bbs(\u03b7 + \u03b8) \u2212 zs\u03bbc(\u03b7 + \u03b8) y[s\u03b7c\u03bbs(\u03b7 + \u03b8) + c\u03b7c(\u03b7 + \u03b8)] + z[s\u03b7c\u03bbc(\u03b7 + \u03b8) \u2212 c\u03b7s(\u03b7 + \u03b8)] y[c\u03b7c\u03bbs(\u03b7 + \u03b8) \u2212 s\u03b7c(\u03b7 + \u03b8)]y + z[c\u03b7c\u03bbc(\u03b7 + \u03b8) + s\u03b7s(\u03b7 + \u03b8)] ] "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001114_17543371jset42-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001114_17543371jset42-Figure1-1.png",
        "caption": "Fig. 1 Main parts of the 3D skiff model, reference frame, and adjustable set-up. The rowlock detailed view shows the different rotation axes. The adjustable parameters were: ho the rowlock height, hfs the foot stretcher height, lfs the foot stretcher longitudinal position, lin the inner arm lever length, lout the outer arm lever length, lia the inter axes distance, a the foot stretcher angle, and b the slide angle",
        "texts": [
            " The boat-and-oars mechanical model was developed by means of the mechanical modeller ADAMS (MD 2008 R3), which works with a Lagrangian formalism. The advantage of using ADAMS was to avoid elaborate mechanical calculations. The boat components were modelled as rigid solids. Each real mechanical element was specified in terms of its mass and inertia matrix. To include flexible elements is possible (oars, for example), but this feature was not used. The 3D mechanical system boat\u00feoars is presented in Fig. 1. The origin of the boat reference frame (OB) was located at the front upper point of the seat when this point was lying in the vertical transverse plane containing the rowlock pins. As shown in Fig. 1, the oar kinematics took into account the relative position of the different rotational axes of the rowlocks. The model could be adjusted for different rower morphologies. The setting parameters were those used on competition boats. To adjust these parameters, some reference points were localized in the boat reference frame: (a) the rowlock pins reference points; (b) the foot stretcher referencepoint as shown in Fig. 1. The feet and the hands of the rower were attached at a nominal position on the foot stretchers and on the oar handle, respectively. The principle for installing the rower in the boat was based on virtual and massless setting devices. These devices were built with suitable joints which permitted translations and rotations depending on their type. This principle was applied for setting the position and the angle of the foot stretcher, the angle of the seat slides, the rowlock, and the oars. It was possible to adjust lin, the effective inner lever and lout, the effective outer lever with two sliding joints and the centre of mass and the inertia of the oars were automatically modified. The lever arms lin and lout are defined in Fig. 1. The virtual setting devices could eventually be driven during the simulation like the other commands without interrupting the simulation. Table 1 gives the adjustable parameter values introduced into the personalized model. These values were measured on the boat of the rower which was modelled for the presented simulations. To simulate the movement of the boat on the water a six-degree-of-freedom virtual joint was adopted whose degrees of freedom could be blocked according to the aim of the simulation"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000645_s00170-007-1183-9-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000645_s00170-007-1183-9-Figure2-1.png",
        "caption": "Fig. 2 Vibration model of the hexapod table",
        "texts": [
            " The pods are connected to the upper moving platform through spherical joints and to the lower stationary platform through universal joints. The lower part of the pod is fixed and actuates the upper part. The latter has a linear motion along the pod\u2019s longitudinal axis. In order to be more specific, the hexapod table under study has been developed for a three-axis CNC milling machine. The physical specifications of the hexapod are presented in Appendix (b). The vibration model proposed by the authors for the hexapod table is illustrated in Fig. 2. To describe the motion of the upper platform as the end effecter, three coordinate systems have been used. The world coordinate frame {W} is fixed to the geometry center of the lower platform. The local coordinate frame {P} is attached to the moving platform at the latter\u2019s geometry center. The position of frame {P} is specified with vector T (x ; y ; z ) defined in frame {W}. The orientation of frame {P} rotating with the moving platform, is described with reference to another local coordinate system {A} located with its center at the geometry center of the upper platform, moving together with frame {P}. Frame {A} does not, however, rotate, and its axes are constantly oriented along their counterparts in frame {W}. The orientation of frame {P} and thus the upper platform is described by \u03c5x, \u03c5y and \u03c5z. This rotation can be expressed in frame {W} with the aid of a rotation matrix, wRP (Appendix a). The upper platform carrying the workpiece for machining has thus six degrees of freedom of motion consisting of three linear movements and three rotations. In the model shown in Fig. 2, the upper and lower parts of the pods, the joints, and the stationary platform have been considered as flexible elements, each possessing its individual mass or inertia, stiffness, and damping. Only one pod is shown in Fig. 2. The moving platform should be strengthened with stiffeners in order to withstand deflection under the payload. This makes this platform rigid in comparison to other elements of the hexapod, with negligible internal friction. This platform has thus been assumed as a lumped mass with infinite stiffness and negligible structural and viscous damping. The parameters shown in Fig. 2 are as follows: MP and IP are the total mass and mass inertia matrices of the moving platform together with the workpiece and fixtures fastened to it, respectively; mdi, mui, mb are the mass of lower and upper part of the pods and the lower platform, respectively (i=1\u20136 for six pods); lpi, lui, ldi and lbi are the displacements of the upper platform, upper and lower parts of the pods and lower platform along the pod\u2019s longitudinal axis, respectively. These displacements are expressed each relative to its lower mass, for instance lPi is measured from the position of mui; Ksi, Kui, Kai, Kdi, Kli and Kb are the stiffness coefficients of the spherical joints, the upper part of pods, the sliding joints, the lower part of pods, the universal joints and the lower platform, respectively",
            " FKi and FCi are the total stiffness and damping forces, respectively, exerted to the platform and can be obtained as follows: FKi \u00bc KTi \u0394lTi ; FCi \u00bc CTi \u0394l Ti \u00f01\u00de where KTi and CTi are the equivalent stiffness and damping coefficient of the support carrying the moving platform at the junction of this platform and the ith spherical joint, respectively; \u0394lTi and \u0394l Ti are the absolute displacement and velocity of this junction along the ith pod\u2019s axis and can be written as follows: \u0394lTi \u00bc \u0394lui \u00fe\u0394lsi \u00fe\u0394lai \u00fe\u0394ldi \u00fe\u0394lli \u00fe\u0394lbi \u00f02\u00de \u0394l Ti \u00bc \u0394l ui \u00fe\u0394l si \u00fe\u0394l ai \u00fe\u0394l di \u00fe\u0394l li \u00fe\u0394l bi \u00f03\u00de where \u0394lsi;\u0394lui;\u0394lai;\u0394ldi;\u0394lli and \u0394lbi are the displacements of the ith spherical joint, pod\u2019s upper part, sliding joint, pod\u2019s lower part, universal joint and the junction of the lower platform with the ith universal joint, respectively; \u0394l si;\u0394l ui;\u0394l ai;\u0394l di;\u0394l li and \u0394l bi are the corresponding velocity increments. It is obvious from Fig. 2 that various supporting elements situated under the upper platform down to the lower one are connected to each other in series along the ith pod\u2019s axis. The following relations hold for the spring forces and damping forces when the elements are in series: FKi \u00bc FKui \u00bc FKai \u00bc FKdi \u00bc FKbi \u00f04\u00de FCi \u00bc FCui \u00bc FCai \u00bc FCdi \u00bc FCbi \u00f05\u00de Using Eqs. (1\u20135), the equivalent stiffness of the ith supporting chain beneath the upper platform, KTi, and similarly the equivalent damping coefficient, CTi, can be found as follows: 1 KTi \u00bc 1 Kui \u00fe 1 Ksi \u00fe 1 Kai \u00fe 1 Kdi \u00fe 1 Kli \u00fe 6 Kb \u00f06\u00de 1 CTi \u00bc 1 Cui \u00fe 1 Csi \u00fe 1 Cai \u00fe 1 Cdi \u00fe 1 Cli \u00fe 6 Cb \u00f07\u00de In the above relations, it is implicitly assumed that the lower platform can be modeled as six concentrated vibratory elements; each situated under one pod, and combined in parallel"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003467_gt2013-95424-Figure3-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003467_gt2013-95424-Figure3-1.png",
        "caption": "Figure 3: Concept of segmented seal fins",
        "texts": [
            " The use of segments as seal fins instead of rolled and caulked in fins will eliminate tangential tensions and forces. Reducing the radial and circumferential stiffness will lead to a lower contact pressure and, therefore, lower wear rates in case of a collision with the rotor. However, with the curved design the pressure difference over the seal will cause an additional down force. To lower the impact on the contact pressure the seal will be manufactured with a slightly larger angle , so that the pressure drop will push the seal from position (1) to its nominal position 2 (cf. Figure 3). Under design conditions the seal will be in contact with the support structure, which holds the seal in place even when the pressure drop exceeds the design limits. In addition, the support seals the circumferential slits caused by the segmentation. Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 01/31/2016 Terms of Use: http://www.asme.org/about-asme/terms-of-use 3 Copyright \u00a9 2013 by ASME Three labyrinth seals configurations were investigated in this study. The geometries are outlined in Figure 4"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000016_elan.200603590-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000016_elan.200603590-Figure1-1.png",
        "caption": "Fig. 1. Schematic diagram of the assembly for exploitation the cylindrical silver electrode. a) electrode during the activation process, b) electrode ready for use, c) construction of the working electrode. 1) corpus with the activation cell, 2A) and B) fittings for inert gas, 3) auxiliary electrode (Pt foil), 4) small electrode Ag/AgCl/3 M KCl, 5) measuring cell, 6) standard reference electrode Ag/AgCl, 7) auxiliary electrode (Pt wire, \u02d8\u00bc 0.7 mm), 8) stirrer bar, 9) and 10) O-ring, 11) slider from the stainless steel, 12) Ag electrode, 13), 14) resin, 15) thermo shrinkable tube.",
        "texts": [
            " The applied electrode was polycrystalline solid silver, which was renovated close before the measurement and was stored inside the device between recordings. The electrode was reactivated before the series ofmeasurements or before each one. Long time stability of metrological parameters of the electrode and reliability of the device components were observed. The roughness factor (RF) in HClO4 solution was determined. The potential of zero charge (Epzc) in NaF solution was calculated and oxygen reduction voltammograms in NaClO4 were recorded. Construction of the device and the procedure of electrode refreshing is explained in Figure 1. The measuring system consists of two electrolytic cells, which are separated. The upper cell, assigned toworking electrode activation is localized inside the small device. The device with the electrode is fastened to the cover of the measuring cell in such manner that its lower part is plunged in the tested solution (Fig. 1). The cylindrical working electrode made from any conducting material mechanically processed moves between two cells. In this work the results for polycrystalline solid silver electrode are presented. The electrode is being stored and electrochemically refreshed between measurements in the activation cell (1) with the volume of about 3 \u2013 5 mL. Therefore this cell is filled by the pure supporting electrolyte. Solution from the activation cell may be continuously deaerated using otherminiature fittings (2A,B)",
            "7 mm) electrode (7) equipped with stirrer bar (8). Both cells are separated by the special O-ring (9), made of chemically inert, elastic material. The O-rings (9) (\u02d8i.d. \u00bc 1.5 mm, \u02d8o.d.\u00bc 10 mm) and (10) (\u02d8i.d. \u00bc 1.5 mm, \u02d8o.d.\u00bc 12 mm) were made from a silicon rubber 4 mm thick. The O-ring prevents from mixing of the solutions and transporting the mechanical contaminants between compartments. The device requires use of the special construction working electrode. Detailed description of the working electrode is given on Figure 1c. Stainless steel wire (11) is inserted in the middle of the silver tube (12) 2 mm in diameter and 1.5 mm long. The silver tube is mounted in such a way that the wire protrudes 6 \u2013 10 mm. Part of the wire surface is covered by isolation resins or composite (13), the excess of which was than mechanically removed (14). The applied resins (or composite) must be mechanically polished and chemically inert to electrolyte and tested sample solution. In the described device both surfaces, i.e., resins or composite and theworking surface of the electrode were hand polished to mirror surface, with polishing grits, from 600 \u2013 2000, followed by alumina 0",
            " In the proposed solution the tested silver electrodes demonstrate many properties which are specific only for the dropping mercury electrode. The described device simplifies the process of electrode surface preparation before or during measurements, however dedicated technology and materials are necessary. The working electrode, presented in this paper on the example of solid silver electrode, may be manufactured from any metallic, glassy carbon, solid amalgam or conducting ceramic material. The project of the device assumes (Fig. 1) that the electrode should bemade from a small tube or bar with the suitable diameter. Operation of the device may be automatically controlled, and can be also used in the flow-through systems. The electrode stays in permanent contact with the supporting electrolyte or tested solution and being exposed for external storage potential maintains working parameters for months. Additionally the electrode is protected against the contamination and mechanical or Electroanalysis 18, 2006, No. 17, 1710 \u2013 1717 www"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002431_iet-nbt.2011.0042-Figure9-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002431_iet-nbt.2011.0042-Figure9-1.png",
        "caption": "Fig. 9 CVs of cobaltferritin on DTSP-modified Au electrode in 0.1 M pH 7.5 containing 0.01 M EDTA",
        "texts": [
            " No aggregated cobaltferritin was observed on the surface. The cobaltferritin molecules were densely packed like a full monolayer. These results indicate that cobaltferritins were successfully immobilised on DTSPmodified gold electrode with approximately less than 40 nm in diameter. IET Nanobiotechnol., 2012, Vol. 6, Iss. 3, pp. 102\u2013109 doi: 10.1049/iet-nbt.2011.0042 Ethylenediaminetetraacetic acid (EDTA) is a powerful complexing agent of Co2+. Therefore it can be used to investigate the exit and entry of ferritin cobalt by cyclic voltammetry. Fig. 9 shows the current\u2013potential curve of cobaltferritin when the electrode immersed in PB pH 7.5, containing 10 mM of EDTA. In the first scan, in the absent of EDTA both the cathodic and anodic peaks presented the same results. In the second scan, both the cathodic and anodic peaks are missing (Fig. 9b) which never appeared in the following scans. It indicated that in the first scan the reduced Co2+ was released through the hydrophilic channel in the ferritin protein shell and was chelated by EDTA. The cobalt ion did not enter the core of the cobaltferritin which could be proved by the absence of the cathodic and anodic peaks in the following scans. This result indicates that cobalt can be induced to exit the ferritin core following the reduction of cobaltferritin. a and b Bare gold electrode c and d Immobilised cobaltferritin on DTSP-modified gold electrode 107 & The Institution of Engineering and Technology 2012 The ferritin protein has hydrophobic and hydrophilic molecular channels through the protein shell, which makes it possible to remove and insert new inorganic phases"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003495_j.mechatronics.2011.11.006-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003495_j.mechatronics.2011.11.006-Figure1-1.png",
        "caption": "Fig. 1. Schematics of the pendulum section. Effective length can be changed by replacing counterweights.",
        "texts": [
            " In Section 2, a brief description of the mechanical part is presented. For making laboratory experiments and on-line control, electrical design, data exchange interface and software tools were created. Their description is given in Section 3. Possible ways for improving the data exchange in multiagent mechatronic complexes are outlined in Section 4, where both hardware-software and algorithmic solutions are presented. Some results of experiments with MMS IPME are described in Section 5. The schematic of a single pendulum section is presented in Fig. 1 (see also photo in Fig. 2). The foundation of the section is a hollow rectangular body. An electrical magnet and electronic controller board are mounted inside the body. The support containing the platform for placing the sensors in its middle part is mounted on the foundation. The pendulum itself has a permanent magnet tip in the bottom part. The working ends of the permanent magnet and the electrical magnet are posed exactly opposite each other and separated with a non-magnetic plate in the window of the body"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003713_978-1-4614-2419-2-Figure13.2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003713_978-1-4614-2419-2-Figure13.2-1.png",
        "caption": "Fig. 13.2 FEA model of the follicle used to support the hair fiber",
        "texts": [
            " In order to accurately model the way in which the hydrogel supports the hair, three models were developed with decreasing fixity at the hair\u2019s base. Modeling was performed using the finite element package Abaqus. The first model included only the hair follicle, with one degree of freedom (translation in one direction), and modeled the hair as a cantilevered beam element. The second model also consisted of only the hair follicle, supporting the base using a simple torsional spring-damper system. The last model fixed the hair into the hydrogel using an embedded region constraint (Fig. 13.2). 13 Characterization of Bio-Inspired Synthetic Hair Cell Sensors 139 The hair was modeled using beam elements, with a length of 19.60 mm, a circular diameter of 69.85 mm and using the properties supplied by researchers at Virginia Tech (Young\u2019s modulus of 8.5 GPa and density of 1.3 g/cm3). The hydrogel was modeled as a cylinder of length 1.50 mm and radius of 0.50 mmwith a hemispherical base having the same radius. It was assumed to have a Young\u2019s modulus of 0.72 kPa and a Poisson\u2019s ratio of 0"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001749_iros.2009.5354769-Figure4-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001749_iros.2009.5354769-Figure4-1.png",
        "caption": "Fig. 4. Desired region seen by robot i",
        "texts": [
            " We consider a bidirectional interactive force between each pair of neighbors. That is, if robot i keeps a distance from robot j then robot j also keeps a distance from robot i. Next, we define a vector x\u0307ri as x\u0307ri = x\u0307o \u2212 \u03b1i\u2206\u03b6i \u2212 \u03b3\u2206\u03c1ij (15) where \u2206\u03b6i is defined in (9), \u2206\u03c1ij is defined in (14), \u03b1i and \u03b3 are positive constants. Let \u2206\u01ebi = \u03b1i\u2206\u03b6i + \u03b3\u2206\u03c1ij , (16) we have x\u0307ri = x\u0307o \u2212 \u2206\u01ebi (17) When the robot i keeps a minimum distance from all its neighboring robots inside the desired region (as illustrated in figure 4), then \u2206\u01ebi = 0. Differentiating (15) with respect to time we get x\u0308ri = x\u0308o \u2212 \u2206\u01eb\u0307i (18) A sliding vector for robot i is then defined as: si = x\u0307i \u2212 x\u0307ri = \u2206x\u0307i + \u2206\u01ebi (19) where \u2206x\u0307i = x\u0307i \u2212 x\u0307o. Differentiating (19) with respect to time yields s\u0307i = x\u0308i \u2212 x\u0308ri = \u2206x\u0308i + \u2206\u01eb\u0307i (20) where \u2206x\u0308i = x\u0308i \u2212 x\u0308o. Substituting equations (19) and (20) into (1), and using property 4 we have Mi(xi)s\u0307i + Ci(xi, x\u0307i)si + Di(xi)si +Yi(xi, x\u0307i, x\u0307ri, x\u0308ri)\u03b8i = ui (21) where Yi(xi, x\u0307i, x\u0307ri, x\u0308ri)\u03b8i = Mi(xi)x\u0308ri + Ci(xi, x\u0307i)x\u0307ri + Di(xi)x\u0307ri + gi(xi)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003818_9781118354162.ch1-Figure1.1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003818_9781118354162.ch1-Figure1.1-1.png",
        "caption": "Figure 1.1 An allegory of a chemical sensor. A sensor is an assembly of a receptor and a signaling (transduction) unit. Adapted with permission from [3]. Copyright 1995 The Royal Society of Chemistry.",
        "texts": [
            " Besides chemical species, micro-organisms and viruses can be traced by means of specific biocompounds such their nucleic acid or membrane components. Physical sensors are devices used to measure physical quantities such as force, pressure, temperature, speed, and many others. The first (and also best known) chemical sensor is the glass electrode for pH determination, which indicates the activity of the hydrogen ion in a solution. When operated, a chemical sensor performs two functions, recognition and transduction, which are exemplified by the allegory in Figure 1.1. First, the analyte interacts in a more or less selective way with the recognition (or sensing) element, which shows affinity for the analyte. The sensing element may be composed of distinct molecular units called recognition receptors. Alternatively, the recognition element can be a material that includes in its composition certain recognition sites. Beyond this, the recognition element can be formed of a material with no distinct recognition sites, but capable of interacting with the analyte. In a chemical sensor, the recognition and transduction function are integrated in the same device"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002331_cdc.2011.6160819-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002331_cdc.2011.6160819-Figure1-1.png",
        "caption": "Fig. 1: Planar UAV model representation.",
        "texts": [
            " The flap\u2019s deflection \u03b4f \u2208 [\u03b4fmin , \u03b4fmax ], the elevator\u2019s deflection \u03b4e \u2208 [\u03b4emin , \u03b4emax ] and the thrust per propeller T \u2208 [Tmin, Tmax] are assumed to be readily available, i.e. the propellers dynamics are much faster than the aircraft dynamics therefore they can be disregarded for control design purposes. Furthermore, the thrust is collinear with the zero-lift line. The aircraft dynamic model construction requires the definition of an Inertial Reference Frame {I} and a Body Reference Frame {B} depicted in Figure 1. The reference frame {I} is fixed at some point in the Earth\u2019s surface which is considered to be flat and still for the current application. It is identified by the set of unitary vectors {iI ,kI} where iI is directed to geographic North and is parallel to the ground and kI is perpendicular to iI and is directed towards the Nadir. The reference frame {B} is identified by the set of unitary vectors {iB ,kB} which match the principal axis of inertia and has its origin at the vehicle\u2019s center of gravity (CG), as represented in Figure 1. For the sake of simplicity, iB is coincident to the zero-lift line. The configuration of the body frame {B} with respect to {I} can be viewed as an element of the Special Euclidean group, (R,p) \u2208 SE(2) where p = [x z]\u1d40, R = [ cos \u03b8 sin \u03b8 \u2212 sin \u03b8 cos \u03b8 ] , and the rotation matrix is parametrized by the pitch angle \u03b8. The kinematic equations of motion for the aircraft are described by \u03b8\u0307 = q, p\u0307 = Rv. (1) where v = [u w]\u1d40 \u2208 R2 is the linear velocity and q \u2208 R the angular velocity, expressed both in {B} (we borrow the standard aircraft nomenclature described in e.g. [18]). The forces and moments acting on the aircraft body which affect its behavior are the gravity fg \u2208 R2, the aerodynamic lift L \u2208 R, the aerodynamic drag D \u2208 R, the thrust T \u2208 R and the aerodynamic torque Ma \u2208 R which are depicted in Figure 1. The gravity force acting on the aircraft is directed downwards along kI and it is given by fg = [0 mg]\u1d40, where m is the vehicle\u2019s mass and g is the acceleration of gravity. The aerodynamic force takes into account both lift and drag components which are described by (see e.g. [19]) Li = 1 2 \u03c1u2 \u221eAiCLi(\u03b1), Di = 1 2 \u03c1u2 \u221eAiCDi(\u03b1), respectively, where i is the lifting surface identifier1, u\u221e denotes the airspeed, \u03c1 denotes the atmospheric pressure, A is the planform area, CL is the Coefficient of Lift and CD is the Coefficient of Drag",
            " where the subscripts p, hs and p, w have been introduced to distinguish the effects of the free-stream flow and that of the propeller slipstream. Since the main contribution to the aircraft drag is that of the wing, the horizontal and vertical stabilizers drag are not taken into account (this approach yields the positive experimental results reported in [6]). Both the wing lift and the horizontal stabilizer lift are normal to the aircraft velocity while the drag is collinear to it, as depicted in Figure 1. According to this geometry, one may construct the aerodynamic forces along the x-axis (Xa) and the z-axis (Za) in terms of both lift and drag as follows[ Xa Za ] = \u2212 [ cos\u03b1 \u2212 sin\u03b1 sin\u03b1 cos\u03b1 ] [ Dw Lhs + Lw ] . (2) The lifting surfaces also create torques which affect the aircraft dynamics because their aerodynamic centers are displaced from the CG along the x-axis. The corresponding torques are given by Mi = \u2212xaciZai (3) where i is the surface identifier and xac is the aerodynamic center location along the x-axis of the {B} frame"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000597_978-3-540-73129-0-Figure6.79-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000597_978-3-540-73129-0-Figure6.79-1.png",
        "caption": "Fig. 6.79. Simple MR fluid direct-shear rotary brake with disc geometry",
        "texts": [
            " Unlike passive viscous dampers, with the MR damper it is easy to achieve large force at very low speed. In a similar fashion, the force developed by a direct-shear device can be divided into F\u03b7 the force due to the viscous drag of the fluid and FMR the force due to magnetic field induced shear stress: F\u03b7 = \u03b7pvSLw h (6.36) FMR = \u03c4MR(H)Lw , (6.37) where vS is the relative velocity. The total force developed by the direct-shear device is the sum of F\u03b7 and FMR. An example of a simple, direct-shear device is shown in Fig. 6.79. In this brake MR fluid is located between the faces of the disc-shaped rotor and the stationary housing. Rotation of the shaft causes the MR fluid to be directly sheared as the rotor moves relative to the housing. A coil fixed in the housing produces a toroidal shaped magnetic field that interacts with the MR fluid in the fluid gaps on each side of the rotor. Torque versus current for the small MRB-2107 brake by Lord Corporation is shown in Fig. 6.80. Measured on-state yield strength \u03c4MR and flux densityB versus magnetic field intensity H for several standard MR fluids from Lord Corporation are given in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001954_j.jsv.2009.01.057-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001954_j.jsv.2009.01.057-Figure1-1.png",
        "caption": "Fig. 1. Multi-folding pantographic systems: (a) multi-layer model, (b) one half of multi-layer model, (c) basic three layer model.",
        "texts": [
            " [3] show that an atomic matrix cellular model can suffer from an explosion of unstable post-buckling states associated with an n-fold compound critical point. The atomic models are analysed in a block diagonal context making use of the local and global symmetries of the structures. The behaviour identified by Ikeda et al. [3] is for bifurcation of rigid bar models where nodes can displace. This paper presents a theoretical approach to examine the elastic stability of a folding multi-layered truss model as shown in Fig. 1(a). The theoretical folding patterns obtained are compared to the actual patterns ee front matter r 2009 Elsevier Ltd. All rights reserved. v.2009.01.057 ing author. esses: mario@hiroshima-u.ac.jp (I. Ario), a.watson@lboro.ac.uk (A. Watson). ARTICLE IN PRESS I. Ario, A. Watson / Journal of Sound and Vibration 324 (2009) 263\u2013282264 seen in the laboratory. The simplest truss has one layer and two elements and is better known as the Von Mises truss. The (shallow) truss has zero stiffness on the point of snap-through behaviour or unstable bifurcation phenomena allowing for geometrical nonlinearity [4\u20137]",
            " the members which were subject to the impact loading initiating the ARTICLE IN PRESS I. Ario, A. Watson / Journal of Sound and Vibration 324 (2009) 263\u2013282 265 folding process. This paper builds upon that work by presenting a generalised analysis of the multiple folding of a multiple layer truss and looking in detail at the folding behaviour of a three layer truss. I.e. folding behaviour initiated by any members in the multiple layer truss. Hence we present in this paper, the folding patterns for the system and derive a general formula for a truss with multiple layers shown in Fig. 1(a). The folding patterns of the three layer truss initiated by the top members presented in the previous paper [11] (see Figs. 6 and 7) are included. The multi-layered truss has right\u2013left symmetry, hence in this paper the theoretical bifurcation analysis is limited to considering symmetric fold patterns. As such, by allowing for symmetric models only, we can therefore consider the half model shown in Fig. 1(b) for the theoretical analysis. The bifurcation paths from the multiple singular point are found using bifurcation analysis. In this paper we establish multiple paths emanating from the initial compound bifurcation and limit point. In general, there is a difference between experimental behaviour and behaviour computed using 0-eigenvalue analysis for the singular stiffness matrix (e.g. see Ref. [12]). For static analysis it is shown that normal modes (approximate eigenvectors) depend on the geometrical nodal condition without eigenvectors from the nonlinear stiffness matrix. The static and dynamic numerical simulations, of the basic truss shown in Fig. 1(c), each identify several folding patterns and these show good agreement with the experimental multi-folding behaviour. The numerical models allow for geometric nonlinearity and contact between nodes. To develop the numerical model the authors estimated the energy that initiates the multi-folding of the structure under impact. The experimental behaviour identified by Holnicki-Szulc et al. [9,10] was controlled by limiting the stress in an individual member and hence active control of the folding characteristics of the truss was achieved, resulting in the several different folding patterns in the experiment. For each of the different folding patterns the folding is initiated by a vertical impact load at the top node (i.e. node 1 in Fig. 1(c)). The model allows for symmetrical folding patterns only and therefore each node is restricted to have a vertical degree of freedom. (The authors do briefly discuss asymmetric folding in the paper.) The folding patterns of a multi-layered system are more complex and depend on the number of layers in the system. For the multi-layered model there are multiple snap-throughs identified after the initial bifurcation point (BP). The folding patterns can be controlled by introducing an imperfection to the position of the central node that displaces in the local snap through",
            " In his paper [13] on the snap through and snap back response of concrete structures Crisfield states that snap through involves a dynamic jump to a new displacement at a fixed load level. Snap back involves a dynamic jump to a new load level at a fixed displacement. In this paper the authors commit the offence of using the term snap through in both contexts. In this section, we consider the folding mechanisms for the three layer truss structure subject to a vertical impact load at the top node shown in Fig. 1(c). The system is a pin-jointed elastic truss and all nodes of the system displace vertically only. No allowance is made for friction or gravity for this geometrically nonlinear problem. We assume a periodic height for each layer of hi \u00bc giL where the width L of the truss is fixed. Therefore, an initial length for each bar in the geometry of the figure is expressed as \u2018i \u00bc ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi L2 \u00fe h2 i q \u00bc L ffiffiffiffiffiffiffiffiffiffiffiffi 1\u00fe g2i q for i \u00bc 1; . . ",
            " The stability of the system is given by a nonzero value for the determinant of the tangent stiffness matrix, the Jacobian for J 2 Rn n. J is defined as follows: J \u00f0Jij\u00de \u00bc @2V @vi@vj \u00bc @2V @v\u0304i@v\u0304j @v\u0304i @v\u0304j @vi @vj \u00bc @Fi @v\u0304j @v\u0304j @vj for i; j \u00bc 1; . . . ; n (17) and instability is defined as det J\u00f0vi\u00de \u00bc 0. (18) It is then possible to determine the buckling load and the post-buckling mode shape of the truss at the singular points from the nonlinear equations during instability. We now determine the equilibrium paths for the basic model shown in Fig. 1(c). The height of each layer is identical, i.e. hi \u00bc h, hence gi \u00bc g. In order to solve for the variable v\u0304i, we use the implicit function theorem and substitute n \u00bc 3 into Eqs. (13) and (14) which gives the solutions as follows: v\u03043 \u00bcF3\u00f0v\u03042\u00de \u00bc v\u03042=2 for primary path; \u00bc 1 2 \u00f0v\u03042 ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 3v\u030422 \u00fe 12gv\u03042 8g2 q \u00de for bif : paths; 8< : (19) v\u03042 \u00bcF2\u00f0v\u03041\u00de \u00bc 2v\u03041=3 for primary path; \u00bc g\u00fe v\u03041 ffiffi 3 p 3 ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi \u00f0v\u03041 g\u00de\u00f0v\u03041 5g\u00de p for bif : paths: ( (20) The use of the implicit function theorem (16) and/or (11) for v\u03041 leads to the following equation: F1\u00f0v\u03041;F2\u00f0v\u03041\u00de\u00de \u00bc f bF1\u00f0v\u03041\u00de \u00bc 0 (21) hence it is seen that the relationship between the load parameter and the displacement v\u03041 is nonlinear f \u00bc bF1\u00f0v\u03041\u00de",
            " The dynamic analysis equation for the folding truss combines mass, damping and nonlinear stiffness F\u00f0v\u00de in the following equation: M \u20ac\u0304v\u00f0t\u00de \u00fe C _\u0304v\u00f0t\u00de \u00fe F\u00f0v\u0304\u00f0t\u00de\u00de \u00bc 0, (25) where M 2 RN N is the mass matrix; C 2 RN N is the damping matrix; F\u00f0 \u00de is the nonlinear stiffness vector; f \u20ac\u0304vi\u00f0t\u00deg T \u00bc \u20ac\u0304v\u00f0t\u00de 2 RN is normalised acceleration; f _\u0304vi\u00f0t\u00deg T \u00bc _\u0304v\u00f0t\u00de 2 RN is the velocity; fv\u0304i\u00f0t\u00deg T \u00bc v\u0304\u00f0t\u00de 2 RN is the normalised displacement and N is the total number of degrees of freedom in the system. If the mass and damping in this system are given as independent uniform variables mi \u00bc m; ci \u00bc c \u00f0i \u00bc 1; . . . ; n\u00de, then we obtain the equation from Eq. (16) for the nodal variables \u20ac\u0304v1\u00f0t\u00de; _\u0304v1\u00f0t\u00de; v\u03041\u00f0t\u00de and this results in the following equation: m \u20ac\u0304v1\u00f0t\u00de \u00fe c _\u0304v1\u00f0t\u00de \u00fe F1 v\u03041\u00f0t\u00de;F2\u00f0v\u03041\u00f0t\u00de\u00de\u00f0 \u00de \u00bc 0. (26) We solve this equation using the dynamic numerical method [5,6] with both incremental load and incremental displacement. The nodes for the FEM model are shown in Fig. 1(c) and Table 2. The extensional stiffness of all members in the model have a normalised value of EA \u00bc 1. The value for geometric stiffness parameter b is given as b \u00bc EA \u00f01\u00fe g2\u00de3=2 \u00bc 1 \u00f01\u00fe 0:6792\u00de3=2 \u00bc 0:566, ARTICLE IN PRESS I. Ario, A. Watson / Journal of Sound and Vibration 324 (2009) 263\u2013282 271 where g \u00bc h1=L \u00bc 0:163 0:240 \u00bc 0:679. For the contact problem we introduce an additional virtual element between nodes 1 and 3; nodes 2 and 4 (and nodes 6 and 5). We assume a constant level of damping, c, and mass, m, for all elements"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001138_03043790701276809-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001138_03043790701276809-Figure1-1.png",
        "caption": "Figure 1. Mechanics problem.",
        "texts": [
            " from theoretical and \u201cclosed\u201d to practical and \u201copen\u201d. And give assignments in different subjects embodying the same concept, principle or method, thus training discernment and at the same time demonstrating relations between theoretical and more practical subjects. Keywords: Discernment; Polyparadigmatic curriculum; Restructuring; Transdisciplinary variation In a sophomore level course in engineering mechanics at MIT\u2014an engineering science course ordinarily of the \u201ctheoretical\u201d type\u2014students are assigned the exercise shown in figure 1. In this context, the students know that they need not worry about the relationship of this abstract representation of this particular problem to anything out there in the so-called real world. They quickly come to understand that the words and figure are to be read, tested, and interpreted only in so far as they relate to finding the (single) answer to the stated problem. The world the student must connect up with is the world of Static Equilibrium, the abstract world of Force Equilibrium, Free Body Diagrams, Force as a Vector, and a bit of Vector Algebra and Trigonometry"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002811_1754337112441112-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002811_1754337112441112-Figure1-1.png",
        "caption": "Figure 1. Instrumentation diagram of the ball dynamic stiffness test apparatus.",
        "texts": [
            "11,14 To reduce the effect of environmental conditions, the temperature and humidity of the following tests were regulated by an environmentally controlled chamber set to conditioning parameters defined in ASTM F2845 (22.2 6 2 C, 50 6 10% RH).15 Balls were acclimatized in the chamber for three weeks prior to testing. To eliminate variation in ball weight, size, and material properties, one ball model was studied (DeMarini A9044). Instrumented ball impacts were performed on a dynamic stiffness test apparatus as shown in Figure 1.11 The test consisted of firing a ball from an air cannon against a fixed solid-steel half cylinder (57mm diameter) and a flat plate. Load cells (PCB, model 208C03) were placed between the impact surface and a massive support where measurements were taken at a sample rate of 100 kHz. While inside the cannon barrel, the ball was cradled by a sabot which controlled speed and launched the ball with no spin. An arrester plate captured the sabot at the end of the cannon, while the ball continued forward"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002930_978-3-642-31988-4_26-Figure3-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002930_978-3-642-31988-4_26-Figure3-1.png",
        "caption": "Fig. 3 Line modeling of a cable",
        "texts": [
            " 5 can be written as: dp dx + \u03c1gE A0 H \u221a 1 + p2 H \u221a 1 + p2 + E A0 = 0 (6) Therefore, x = \u2212 H \u03c1g sh\u22121 ( dz dx ) \u2212 H2 \u03c1gE A0 dz dx + c (7) Where, sh\u22121(x) = ln ( x + \u221a 1 + x2 ) , x \u2208 (\u2212\u221e,+\u221e) (8) x = \u2212 H \u03c1g ln \u239b \u239d dz dx + \u221a 1 + ( dz dx )2 \u239e \u23a0 \u2212 H2 \u03c1gE A0 dz dx + c (9) Integrating and applying the boundary conditions as follows: x = 0, z = 0 x = L , z = h (10) The length of cable is l, the unstrained length of the cable is l0, and \u0394l represents the strain of the cable. The relationship can be expressed as: l = l0 + \u0394l. l = \u222b l 0 \u221a 1 + ( dz dx 2) dx (11) l0 = \u222b 1 T E A0 + 1 dl (12) A line is shown in Fig. 3. A cable under its mass cannot remain straight. The idealization would be possible if the ends of the cable were subjected to tensions that are predominantly larger than the effect of the cable mass or the accuracy requirement is not high. In Fig. 3, line equation with elastic deformation of a cable can be easily derived as follows: l = ( h2 + L2 )1/2 (13) T = ( V 2 + H2 )1/2 (14) L = l H T (15) h = l V T (16) \u0394l = T l E A0 (17) where \u0394l is elastic deformation of the cable. In Fig. 4, for studying the largest radio telescope FAST, a similarity model of FAST is set up in Beijing. The related geometric parameters of the six-cable driven parallel manipulator in the similarity model are given in Table 1 [9]. In Fig. 5, two coordinates are set up for the six-cable driven parallel manipulator: an inertial frame : O \u2212 XY Z is located at the center of the reflectors bottom"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000388_14644193jmbd97-Figure10-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000388_14644193jmbd97-Figure10-1.png",
        "caption": "Fig. 10 Non-linear elastic model of the bearing",
        "texts": [
            " The deflections and contact angles for the RHS bearing are, respectively \u00f0di\u00deR \u00bc Bdb cos\u00f0a0\u00de \u00fe x cos\u00f0ui\u00de \u00fe y sin\u00f0ui\u00de b1\u00f0fi\u00de cos\u00f0ui\u00de \u00feb1\u00f0ci\u00de cos\u00f0fi\u00de sin\u00f0ui\u00de \u00feR\u00f01 cos\u00f0fi\u00de\u00de cos\u00f0ui\u00de \u00feR\u00f01 cos\u00f0ci\u00de\u00de sin\u00f0ui\u00de 0 BBBBBB@ 1 CCCCCCA 2 \u00fe Bdb sin\u00f0a0\u00de \u00fe z0 \u00fe z \u00fe b1\u00f01 cos\u00f0fi\u00de\u00de cos\u00f0ui\u00de \u00fe b1\u00f01 cos\u00f0ci\u00de\u00de cos\u00f0fi\u00de sin\u00f0ui\u00de \u00f0fi\u00deR cos\u00f0ui\u00de \u00fe \u00f0ci\u00deR sin\u00f0ui\u00de 0 BBB@ 1 CCCA 2 2 66666666666666664 3 77777777777777775 1=2 Bdb \u00f021\u00de \u00f0ai\u00deR \u00bc tan 1 Bdb sin\u00f0a0\u00de \u00fe z0 \u00fe z \u00fe b1\u00f01 cos\u00f0fi\u00de\u00de cos\u00f0ui\u00de \u00fe b1\u00f01 cos\u00f0ci\u00de\u00de cos\u00f0fi\u00de sin\u00f0ui\u00de \u00f0fi\u00deR cos\u00f0ui\u00de \u00fe \u00f0ci\u00deR sin\u00f0ui\u00de 0 BB@ 1 CCA Bdb cos\u00f0a0\u00de \u00fe x cos\u00f0ui\u00de \u00fe y sin\u00f0ui\u00de b1\u00f0fi\u00de cos\u00f0ui\u00de b1\u00f0ci\u00de cos\u00f0fi\u00de sin\u00f0ui\u00de \u00feR\u00f01 cos\u00f0fi\u00de\u00de cos\u00f0ui\u00de \u00feR\u00f01 cos\u00f0ci\u00de\u00de sin\u00f0ui\u00de 0 BBBB@ 1 CCCCA 2 666666666666664 3 777777777777775 \u00f022\u00de The deflection for the ith ball was calculated in the previous section. The force due to this deflection for a single ball can be found and then the total force in the bearing can be calculated as this ball rotates around the inner ring. In the model, the contacts of balls to the inner and outer races are represented by non-linear contact springs, i.e. balls act as massless springs. The elastic model of the bearing is represented in Fig. 10. To calculate the total contact force on the ith ball in Fig. 1, the reference axes should be determined and the total deflection, and hence the forces with respect to these axes, should be calculated. The deflection for each ball can be calculated as described earlier. Having calculated the deflection for the ith ball in its contact direction, the force in the same direction Proc. IMechE Vol. 222 Part K: J. Multi-body Dynamics JMBD97 # IMechE 2008 at UNIV PRINCE EDWARD ISLAND on August 5, 2015pik"
        ],
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    },
    {
        "image_filename": "designv11_12_0003603_0954406212466479-Figure13-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003603_0954406212466479-Figure13-1.png",
        "caption": "Figure 13. Geometry for determination of interference.",
        "texts": [
            " Then, the problem here becomes a solution for the smallest tooth number of the internal gear, which insures that there is no interference in the gear pair. Figure 12 presents a limiting case to make the interference analysis, in which the tip point of the pinion tooth, A, just touches the tip point of the internal gear tooth, K. Obviously, the trajectory of point A in S2 passes through point K, which is the intersect point between the work profile and the addendum circle of the internal gear (Figure 13). The trajectory of point A can be obtained from equation (34) xtra\u00bc cos\u00f0\u20191 \u20192\u00dexa sin\u00f0\u20191 \u20192\u00deya\u00fe\u00f0r2 r1\u00desin\u20192 ytra\u00bc sin\u00f0\u20191 \u20192\u00dexa\u00fe cos\u00f0\u20191 \u20192\u00deya\u00fe\u00f0r2 r1\u00decos\u20192 \u00f055\u00de Therefore xk\u00bc cos\u00f0\u20191 \u20192\u00dexa sin\u00f0\u20191 \u20192\u00deya\u00fe\u00f0r2 r1\u00desin\u20192 yk\u00bc sin\u00f0\u20191 \u20192\u00dexa\u00fe cos\u00f0\u20191 \u20192\u00deya\u00fe\u00f0r2 r1\u00decos\u20192 \u00f056\u00de where (xk, yk) is the coordinate of point K in S2. at University of Bristol Library on January 6, 2015pic.sagepub.comDownloaded from Since point K is the intersect point between the work profile and the addendum circle, the equation of the work profile should be derived first",
            " By substituting equation (7) into equation (34), the equation of the work profile can be obtained as follows x2\u00bccos\u00f0\u20191 \u20192\u00dex1 sin\u00f0\u20191 \u20192\u00de\u00f0kx1\u00feb\u00de\u00fe\u00f0r2 r1\u00desin\u20192 y2\u00bcsin\u00f0\u20191 \u20192\u00dex1\u00fecos\u00f0\u20191 \u20192\u00de\u00f0kx1\u00feb\u00de\u00fe\u00f0r2 r1\u00decos\u20192 ( \u00f057\u00de Equations (9) and (57) yield another form of the work profile y2 \u00bc cot\u00f0\u20191 \u20192 \u00fe \u00dex2 \u00fe b sin \u00fe \u00f0r2 r1\u00de sin\u00f0\u20191 \u00fe \u00de sin\u00f0\u20191 \u20192 \u00fe \u00de \u00f058\u00de Here cot\u00f0\u20191 \u20192 \u00fe \u00de \u00bc 1 tan\u00f0\u20191 \u20192\u00de tan tan\u00f0\u20191 \u20192\u00de \u00fe tan \u00f059\u00de \u20191 \u20192 \u00bc z2 z1 z2 \u20191 \u00f060\u00de Equation (58) shows that the work profile of the internal gear has a similar form to that of the straight line. However, its slope and intercept are not constant, as illustrated by equations (59) and (60). Since the tooth number difference is very small compared to the tooth number of the internal gear, an approximation can be obtained cot\u00f0\u20191 \u20192 \u00fe \u00de cot \u00bc k \u00f061\u00de Then, we replace the work profile by an approximate line, as represented by equation (62) y2 \u00bc kx2 \u00fe b0 \u00f062\u00de As shown in Figure 13, the approximate line intersects the addendum circle at point K0, which is used to replace point K in the solution process. Then, we get an approximate profile, as shown in Figure 14. According to the definition of the gear with straightline profile, the approximate line passes through point C and a relation can be obtained xc \u00bc r2 sin c yc \u00bc r2 cos c \u00f063\u00de where (xc, yc) is the coordinate of point C in S2. Here c \u00bc s 2r2 \u00f064\u00de Equations (9) and (62) to (64) yield b0 \u00bc r2 sin sin s 2r2 \u00fe \u00f065\u00de The coordinate of point K0 can be solved by equation (66) yk0 \u00bc kxk0 \u00fe b0 x2 k0 \u00fe y2 k0 \u00bc r2 a2 \u00f066\u00de where (xk0 , yk0 ) is the coordinate of point K0 in S2"
        ],
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    },
    {
        "image_filename": "designv11_12_0001196_0094-114x(72)90004-3-Figure3-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001196_0094-114x(72)90004-3-Figure3-1.png",
        "caption": "Figure 3. Joint i in isometric projection.",
        "texts": [
            " The unit vector y~,~ is defined as Yi+l = Z~+l \u00d7 xi+l. (2) The twist angle, a~, is measured from +z~ to +zi\u00f7, in a positive sense about \u00f7x~+, according to the right-hand rule. The link length, ai, is measured along x,\u00f7l from its intersection with z~ tO its intersection with ~+t. This length can be negative if xt+l has the opposite sense. The rotation, 0i, is measured from +x, to +xf+l in a positive sense about z,, and the translation, s~, is measured along z~ from xi to x~+~ in a manner similar to that of a~. Figure 3 is an isometric drawing of joint i (axis z~) and the two links, i - 1 and i, which contain joint i. The figure illustrates the quantities 0~ and s~. The following restrictions on the values of the chain parameters are required to insure the existence of one and only one mathematical model for every linkage of the R C C C chain. 1. Three twist angles must be ~< 90 deg. 2. The fourth twist angle must be ~< 90 deg or be the twist angle closest to 90 deg if it is i> 90 deg. 3. The fixed translation, Sl, must be positive"
        ],
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    },
    {
        "image_filename": "designv11_12_0000410_0094-114x(80)90020-8-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000410_0094-114x(80)90020-8-Figure2-1.png",
        "caption": "Figure 2.",
        "texts": [
            " For the line coincident with $2 we may write the unit line vector I~ = (A2, ~'~), where Hence, ti2 = h2cb2 + v~. Of significance in what follows will be the algebraic condition for two lines, given by Ii and 12 say, to intersect. We may state it as (o1\" v2 + cb2\" vl =0. (1.1) This condition is equivalent to that for two screws of zero pitch to be reciprocal, the more general case being governed by &l\"/12 + d~2\" ill = 0. (1.2) The form of the closure equations to be used is well-known, the standard symbols being most conveniently defined by means of Fig. 2. In addition, we use the abbreviations s for sine and c for cosine. For the linkage we choose as our leading example, the 6-bar closure equations are the relevant ones. A basic set of twelve equations is given in the Appendix. It is now a common ploy to obtain alternative equations by advancing subscripts in those listed. A second, less common, example will afterwards be used to further illustrate the flexibility and power of the theory. 2. Discussion We choose to explain the detailed application of screw system algebra to special configurations by means of the R-S-C-R- chain, for two reasons"
        ],
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    },
    {
        "image_filename": "designv11_12_0002733_0954410012464002-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002733_0954410012464002-Figure2-1.png",
        "caption": "Figure 2. Level banked turn.",
        "texts": [
            "35\u201339 For maneuvering flight, the aircraft linear velocities are (Figure 1) u v w T \u00bc VA cos\u00f0 F\u00decos\u00f0 F\u00de VA sin\u00f0 F\u00de VA sin\u00f0 F\u00decos\u00f0 F\u00de T \u00f02\u00de where fuselage angle of attack, F, and sideslip, F, are given by F \u00bc tan 1 w=u\u00f0 \u00de; F \u00bc sin 1 v=VA\u00f0 \u00de \u00f03\u00de Level banked turn is a maneuver in which the helicopter banks towards the center of the turning circle. For helicopters, the fuselage roll angle, A, is in general slightly different than the bank angle, B. For coordinated banked turn A \u00bc B. A picture describing these angles for a particular case ( A \u00bc 0) is given in Figure 2, where Fresultant is the sum of the gravitational force (W) and the centrifugal force (Fcf). Helical turn is a maneuver in which the helicopter moves along a helix with constant speed (Figure 3). In a helical turn, the flight path angle is different than zero being given by sin\u00f0 FP\u00de \u00bc sin\u00f0 A\u00de cos\u00f0 F\u00de cos\u00f0 F\u00de sin\u00f0 A\u00de cos\u00f0 A\u00de sin\u00f0 F\u00de cos\u00f0 A\u00de cos\u00f0 A\u00de sin\u00f0 F\u00de cos\u00f0 F\u00de \u00f04\u00de A picture describing the flight path angle for a particular case, A \u00bc 0, F \u00bc 0, is given in Figure 4. Note that _ A 4 0 is a clockwise turn and _ A 5 0 is a counterclockwise turn (viewed from the top) while FP 4 0 is referring to the ascending flight and FP 5 0 is referring to the descending flight"
        ],
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    },
    {
        "image_filename": "designv11_12_0003548_1.4006667-Figure10-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003548_1.4006667-Figure10-1.png",
        "caption": "Fig. 10 Sector Mesh of the 24-bladed Blisk",
        "texts": [
            " Note that, since the bounds are set according to the probability density function of the absolute error, tighter set of bounds have larger probability of missing the global optimum if the bounds are fixed. Therefore, such a relaxation procedure increases the flexibility of the method by altering the solution domain even if the initial solution domain does not include the global optimum. 4.2 Case Study III: OptID-Genetic Optimization. In this case study, to address a relatively more realistic problem, a 24 bladed blisk finite element model is created, a sector of which is given in Fig. 10, and its ROM is structured according to the component mode synthesis presented in Sec. 2.2. Ten modes for each blade and for each nodal diameter of the disk for the first modefamily are involved in the ROM. The resulting number of degrees of freedom is 658 in the ROM; whereas for the full finite element model it is 10,368. Mistuning is applied through blade cantilever natural frequency perturbations, dk n in Sec. 2.2, and a fleet of 45,000 disks are simulated with random mistuning using normal distribution with zero mean and 1"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003214_eeeic-2.2013.6737911-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003214_eeeic-2.2013.6737911-Figure1-1.png",
        "caption": "Figure 1. SRM with 6x4 topology.",
        "texts": [
            " Torque production occurs by the tendency of alignment of stator and rotor poles when the respective phase is energized, position of minimum reluctance for the established magnetic circuit (and maximum inductance). On the other hand, if mechanical energy is applied to the machine shaft, phase energization provides restorative torque resulting in an additive back-electromotive force, thereby generating electric power. For the SRM\u2019s, mechanical energy applied to the machine shaft by an external agent is converted to electric power by the exercised force to dishevel the stator and rotor poles. Fig. 1 illustrates the transversal cut of a 6x4 topology SRM representing the winding of a single phase. The mathematical modeling presented in subsection A indicates how the instants of current application determine the machine operation as motor or generator. The operation is related to the inductance variation as function of rotor position, as illustrated in Fig. 2. 978-1-4799-2803-3/13/$31.00 \u00a92013 IEEE The RL circuit of a SRM can be represented mathematically as: v = Ri+ L di dt + i\u03c9 dL d\u03b8 (1) where: (v) is the applied voltage to a phase winding, (i) is the winding current, (L) is the phase inductance, (\u03c9) is the rotor speed and (\u03b8) is the rotor instantaneous angular position",
            " 2, if current circulates in a phase winding during positive variation of its inductance, the produced torque is positive. On the other hand, if current circulates in a phase winding during negative variation of its inductance, the produced torque is negative, which means that this torque is restorative and is converted to backelectromotive force, allowing a generator operation. Considering the friction coefficient (D) and inertia moment (J) the torque can be written as: Tm = Temag \u2212 J d\u03c9 dt \u2212D\u03c9 (4) For the three phases SRM illustrated in Fig. 1, the electromagnetic torque (Temag) can be rewritten as function of the three currents: Temag = 1 2 (i2a dLa d\u03b8 + i2b dLb d\u03b8 + i2c dLc d\u03b8 ) (5) Considering the rotor speed: \u03c9 = d\u03b8 dt (6) Thus, the mathematical model comtemplating the three phases of the 6x4 SRM is [7]: \u03b1 = \u03b2 \u00b7 \u03b3 + \u03b4 \u00b7 \u03b6 (7) Being: \u03b1 =  va vb vc Tm 0 , \u03b3 =  ia ib ic \u03c9 \u03b8 , \u03b6 =  i\u0307a i\u0307b i\u0307c \u03c9\u0307 \u03b8\u0307 , \u03b2 =  Ra 0 0 0 0 0 Rb 0 0 0 0 0 Rc 0 0 r1ia r2ib r3ic D 0 0 0 0 \u22121 0  and \u03b4 =  La 0 0 0 ia dLa d\u03b8 0 Lb 0 0 ib dLb d\u03b8 0 0 Lc 0 ic dLc d\u03b8 0 0 0 J 0 0 0 0 0 1 "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000673_bibe.2008.4696658-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000673_bibe.2008.4696658-Figure1-1.png",
        "caption": "Fig. 1. Display of runs, tumbles and the respective parameters.",
        "texts": [
            "eyword: E. coli; Chemotaxis; Random motility coefficient; Random walk; I. INTRODUCTION OF E. coli AND CHEMOTAXIS If we look at a single E. coli\u2019s movement without chemical stimulus gradient through the microscope, we can observe two typical types of movements: run and tumble. The track of the E. coli movement can be seen as a series of runs and tumbles, shown in Fig. 1. The run is the smooth segment of the random walk. During runs, the cell keeps on a reasonable smooth track. During tumbles, the cell orients itself and selects a new direction to start another smooth run. It is the tumble that gives the cell a nearly random reorientation from which to begin the next run.[1] Motility of E. coli is determined by the rotation mode of the flagellar filaments, each of which is driven by a reversible rotary motor located at the base of it. During a run, the flagella rotate in the counterclockwise direction (as viewed from the distal end) and form a coordinated bundle according to the flagellar left-handed helicity, and as a result form a nearly constant propulsion and drive the cell in a smooth track"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001517_j.mechmachtheory.2009.12.004-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001517_j.mechmachtheory.2009.12.004-Figure1-1.png",
        "caption": "Fig. 1. A Stephenson six-bar linkage.",
        "texts": [
            " Generally speaking, among the various linkage mobility issues, the branch should be identified and rectified first [10,11]. A sub-branch refers to a linkage configuration space, in which transformation between configurations may be accomplished without reaching a singularity or dead center position, where the linkage may lose control [10\u201312]. In a sub-branch, an input value corresponds to one and only one linkage configuration and the order of motion of a linkage is determined by the magnitude order of the input value [12]. A Stephenson six-bar linkage, as shown in Fig. 1, consists of a four-bar loop ABCDA and a five-bar loop ABEFGA. A Watt six-bar linkage, as shown in Fig. 2, consists of two four-bar loops. A Watt six-bar linkage may also be regarded as one comprised of a four-bar loop ABCDA and a six-bar loop ABCEFGA or even a five-bar loop by the stretch and rotation of the second four-bar loop [11]. Two other six-bar linkages are shown in Figs. 3 and 4. In the following discussion, any link in the linkage may be used as the reference link. Hence, the treatment is valid for any linkage inversion",
            " The two dead center positions in a branch separate the branch into two sub-branches (Fig. 7b) represented by the two equations in Eq. (9). The dimensions for the four-bar linkages in Figs. 1 and 2 are a1 = 5.0, a2 = 2.82, a3 = 5.0, a4 = 2.2 and a = 0 for the plotting in Fig. 7. One may note that branch formation is irrelevant to the choice of the input joint while the sub-branch formation is affected. The discriminant and the relevant equations in the form of Eqs. (6)\u2013(9) can be derived from any bimodal linkage. RRRRR loop: A Stephenson six-bar chain (Fig. 1) contains a five-bar loop ABEFGA and its mobility is affected by it. The loop equation for the chain ABEFGA in Fig. 1 can be expressed as a2eih2 \u00fe a9ei\u00f0h3\u00feb\u00de a7 a5eih5 \u00bc a6eih6 : \u00f013\u00de Eliminating eih6 and using the tangent half-angle formula x5 \u00bc tan\u00f0h5=2\u00de: \u00f014\u00de Eq. (13) can be written as P2x2 5 \u00fe Q2x5 \u00fe R2 \u00bc 0; \u00f015\u00de where P2 \u00bc a2 2 a2 5 \u00fe a2 6 \u00fe a2 7 \u00fe a2 9 2a6a7 \u00fe 2\u00f0a6a9 a7a9\u00de cos\u00f0h3 \u00fe b\u00de \u00fe 2a2a9 cos\u00f0h2 h3 b\u00de \u00fe 2\u00f0a2a6 a2a7\u00de cos h2; \u00f016a\u00de Q 2 \u00bc 4a2a6 sin h2 4a6a9 sin\u00f0h3 b\u00de; \u00f016b\u00de R2 \u00bc a2 2 a2 5 \u00fe a2 6 \u00fe a2 7 \u00fe a2 9 \u00fe 2a6a7 2\u00f0a6a9 \u00fe a7a9\u00de cos\u00f0h3 \u00fe b\u00de \u00fe 2a2a9 cos\u00f0h2 h3 b\u00de 2\u00f0a2a6 \u00fe a2a7\u00de cos h2: \u00f016c\u00de Eq. (15) is a quadratic equation",
            " (9a) and (20a); Sub-branch n-4: Segment 4-n with h2 e [49.7 , 173.6 ] and Eqs. (9a) and (20b); Sub-branch n-3: Segment n-3 with h2 e [150.5 , 173.6 ] and Eqs. (9b) and (20a); Sub-branch 3-n: Segment 3-n with h2 e [150.5 , 173.6 ] and Eqs. (9b) and (20b). Branch 6-5: This branch contains segment 6-5. Sub-branch 6-5: Segment 6-5 with h2 e [ 25.3 , 21.8 ] and Eqs. (9a) and (20a); Sub-branch 5-6: Segment 5-6 with h2 e [ 25.3 , 21.8 ] and Eqs. (9a) and (20b). Example 2. Given the dimensions for the Stephenson six-bar linkage in Fig. 1 as follows, a1 = 4.35, a2 = 4.8, a3 = 2.5, a4 = 1.56, a5 = 3.62, a6 = 4.8, a7 = 10.0, a9 = 4.16, a10 = 2.09, a = 34.9 . With the above given dimensions, the plots for the Stephenson six-bar linkage are shown in Fig. 11, where the four-bar curve is drawn from Eq. (4) and the motion domain (shade area) is drawn from Eq. (17). The mobility analysis of Stephenson linkage with the proposed method above can be carried out as follows, (1) Branch identification of four-bar chain: From Eq. (8) or Eq. (10) with D = 0, there are four dead center positions e( 17"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000669_1.2410018-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000669_1.2410018-Figure1-1.png",
        "caption": "Fig. 1. A block sliding on the surface of a stationary sphere of radius R. the block is approximated as a point particle, and its motion parametrized by the polar angle .",
        "texts": [
            " A comparison of the two solutions identifies a parameter range where the perturbation series accurately represents the motion of the particle and another range where the perturbative solution fails qualitatively to describe the motion of the particle. \u00a9 2007 American Association of Physics Teachers. DOI: 10.1119/1.2410018 I. INTRODUCTION A classic problem in introductory mechanics1,2 is to determine the release point for a block of mass m that slides without friction on the surface of a sphere or cylinder of radius R starting from rest that is, with an infinitesimal velocity at the top see Fig. 1 . The answer is that the release point, the point at which the block first loses contact with the sphere, occurs when it has fallen a vertical distance h=R /3. The result is interesting because it is independent of both the particle mass m and the gravitational acceleration g. It is useful to extend this analysis to include frictional effects, which we model by the frictional force Ff = FN opposing the motion in the tangent plane of the sphere; FN is the normal force and is the coefficient of kinetic friction"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000154_iros.2007.4399278-Figure7-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000154_iros.2007.4399278-Figure7-1.png",
        "caption": "Fig. 7. Yaw Control.",
        "texts": [
            " This controller stabilizes the platform pitch angle system because the roots of the characteristic equation 0.1s3+5s2+908.8s+ 4098 are located at \u221222.7 \u00b1 91.45 and \u22124.62 which are in the left hand side of the complex plane. Now, to control the vehicle yaw position, it is assumed that the pitch and roll angles are stabilized, then the roll rate and pitch rate vanish, then equation (9) can be written as follows: \u03c8\u0308 = n/Jz (28) where n is the vehicle yaw moment. Notice that n is used to control yaw during hovering flight and to control roll during forward flight as shown in Figure 7. Under this assumption, the yaw moment can be approximated by the following expression n = \u03c1V 2SbCn/4 (29) where b is the wing span and Cn is the yawing moment coefficient given by Cn = Cn\u03c8\u0307 \u03c8\u0307 + Cn\u03b4e \u03b4e. Then, (28) can be rewritten as \u03c8\u0308 = (\u03c1V 2Sb)(Cn\u03c8\u0307 \u03c8\u0307 + Cn\u03b4e \u03b4e)/4Jz (30) where Cn\u03b4e = 0.19 represents the variation of the yaw moment with respect to the ailerons positions. Cn\u03c8\u0307 = 0.19 is the yaw damping derivative. Applying Laplace transform and using numerical values yields: \u03c8(s) \u03b4e(s) = 20 s2 + 20s (31) Then, using the actuator dynamics given before, a closedloop control system with a proportional derivative controller can be proposed, where Kp = 68 and Kd = 17"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000597_978-3-540-73129-0-Figure6.104-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000597_978-3-540-73129-0-Figure6.104-1.png",
        "caption": "Fig. 6.104. Schematic drawing of a modified Mc Kibben actuator",
        "texts": [
            " A method to transfer and amplify the radial strain of a CP fiber into an axial strain has been proposed, inspired to a Mc Kibben actuator [245]. In the classical version of this latter device, a cylindrical rubber bladder is covered by a braid mesh, made of flexible, but not extensible, threads. Both ends of the bladder are connected to the mesh. By changing the force applied to the free end of the mesh and the pressure inside the bladder, the mesh shape change dimensions: its diameter increases and its length decreases. In the CP version of the Mc Kibben actuator (sketched in Fig. 6.104), the bladder is substituted with a bundle of conducting polymer hollow fibers. In the center of each hollow fiber a rigid metal wire works as a counterelectrode. A filling liquid electrolyte completes the system. The actuation mechanics of such a device has been studied, by performing an electro-chemo-mechanical analysis [246]. According to results of that study, this type of structure might enable axial strains of different magnitude, ranging from 25% up to 80%, depending on the inclination angle of the mesh"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002855_pbce077e_ch14-Figure14.2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002855_pbce077e_ch14-Figure14.2-1.png",
        "caption": "Figure 14.2 (a) The mechanical design of the coupling between the module and the vehicle. (b) The first payload, developed by NUS, carrying a Wi-Fi antenna and the NUS acoustic modem. (c) The first eFolaga",
        "texts": [
            " The module should not change its electrical, mechanical or hydrodynamic properties during operation in any way that could cause the eFolaga to malfunction or to compromise control. Vehicle safety (emergency sensors and watchdogs) must remain in charge of the native eFolaga system. The first step in the project has been the definition of mechanical modularity to separate the vehicle into forward and aft sections to interpose an extra piece of hull (the module) containing added functionality between the two (Figure 14.2). It was envisioned that multiple modules could be stacked together. As the eFolaga hull is made of GRP filament wound, it is not possible to directly fashion a watertight interface between mating hull sections. Hence, it has been decided to employ machined inserts, male on one section and female on the other, bonded into the GRP hull, which could then be sealed against each other with O-rings. The design of these coupling inserts was based on the end bulkheads already in use. Notable in the design of the coupling inserts is the long penetration distance (over one-third of the diameter), which allows one section to support the other without putting the three radial locking screws under load"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000213_iecon.2006.348006-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000213_iecon.2006.348006-Figure2-1.png",
        "caption": "Fig. 2- Stator and rotor winding axes (s1 open).",
        "texts": [
            " Here, because of the space restrictions, we give only the model under fault conditions. To do this, we begin by the following stator and rotor voltage equations: [ ] [ ][ ] [ ][ ] [ ][ ])..(. rsrssssss ILIL dt dIRV ++= (1-a) [ ] [ ][ ] [ ][ ] [ ][ ])..(. srsrrrrrr ILIL dt dIRV ++= (1-b) In which the current and voltage vectors are: [ ] [ ] [ ] [ ] [ ] [ ] [ ] [ ]Trrrrrrr T r T ssssss T ssssss iiiiiiI V iiiiiI vvvvvV 654321 65432 65432 000000 = = = = (1-c) 13631-4244-0136-4/06/$20.00 '2006 IEEE In (1), we supposed that the phase s1 is opened (Fig. 2)1. In these equations, [Rs], [Rr], [Lss], [Lrr], [Lsr] and [Lrs] are parameters of SPIM given in [2]. It must be noted that contrarily to the resistance matrices [Rs] and [Rr], the inductance matrices are not diagonal. This means that the model (1) is highly coupled. As can be seen from (1), the SPIM is a five dimensional system in the stator when one of the stator phases is opened. Meanwhile, the rotor will be modeled as a six dimensional system. In order to obtain a block-decoupled model, it is shown in [2] that two transformation matrices are needed"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002420_s10237-009-0187-9-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002420_s10237-009-0187-9-Figure1-1.png",
        "caption": "Fig. 1 Decomposition of the deformation gradient as F = Fa Fe. The reference, intermediate and deformed configurations are indicated by \u03a90,\u03a9 and\u03a9t , respectively. Other variables associated to each configuration and map are indicated in the figure and described in the text",
        "texts": [
            "1 Kinematics We analyse in this section the deformation of a body B with reference configuration \u03a90 \u2208 R 3 and material coordinates X0 into a deformed configuration \u03a9t \u2208 R 3 where the material points are located at x = \u03c7(X0, t), with \u03c7(X0, t) : R 3 \u00d7 R \u2192 R 3 the map of the whole motion. For clarity, we remove the dependence on the time variable t and the position X0, and simply denote by x the map \u03c7 . The notation employed here and throughout the article has been listed in Table 1. As it is customary in biomechanics since the seminal work of Rodriguez et al. (1994), we use a multiplicative decomposition of the deformation gradient F = \u2202x \u2202X0 into F = Fe Fa (see Fig. 1). In the present case, the active deformation gradient Fa = \u2202X \u2202X0 is due to the growth process in the cell, which in turn is due to the mechanotransduction of the genetically regulated chemical reactions that take place in the cytoskeleton. On the other hand, the tensor Fe = \u2202x \u2202X represents the passive elastic deformation due to the elastic response of either the cytoskeleton and the cytoplasm. The intermediate configuration \u03a9 is the one obtained after removing the elastic deformation from \u03a9t , which may give rise to material incompatibilities (tears and overlappings, see Fig. 1). Therefore, the two tensors Fa and Fe may be discontinuous, while F is continuous. From the decomposition of the deformation gradient, and setting J = det F, Ja = det Fa and Je = det Fe, we have that J = Ja Je. We denote by \u03c10, \u03c1 and \u03c1t the densities in the reference, intermediate and deformed configurations, respectively. Analogously, we denote by (dM0, dV0), (dM, dV) and (dm, dv) the pairs of differentials of mass and volume in the same configurations (see Fig. 1), which are related by dM0 = \u03c10dV0, dM = \u03c1dV and dm = \u03c1t dv. 2.2 Balance equations Throughout the paper, we use the following assumptions: \u2022 The active deformations do not introduce any change of density, (i.e. \u03c1 = \u03c10,\u2200t), although changes of mass (growth) may occur when passing from configuration\u03a90 to \u03a9 , (i.e. \u03c10dV0 = \u03c1dV in general). \u2022 The elastic deformation may introduce density changes when passing from the configuration\u03a9 to\u03a9t , (i.e.\u03c1 = \u03c1t in general), although this transformation preserves the mass, (i"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002129_s0263574709005426-Figure6-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002129_s0263574709005426-Figure6-1.png",
        "caption": "Fig. 6. PPR Assur virtual chain.",
        "texts": [
            " a kinematic subchain with null mobility that when connected to another kinematic chain preserves mobility. Due to the analogy with Assur groups and to avoid possible confusions with other unrelated virtual chains in the literature, we coined the name Assur virtual chain. 3.4.1. The orthogonal PPR Assur virtual chain. The PPR virtual chain is composed of two virtual links (C1, C2) connected by two prismatic joints, whose movements are in the x and y orthogonal directions, and a rotational joint, whose the movement is in the z direction (see Fig. 6). The prismatic joints are called px and py, and the rotative joint is called rz. The first prismatic joint (px) and the rotative joint (rz) are attached to the chain to be analysed (real chain). The joint px connects the link R1 with the virtual link C1, the joint py connects the virtual link C1 with the virtual link C2 and the joint rz connect the virtual link C3 with the real link R2. Let the twist $px represent the movement of the link C1 in relation to the link R1, twist $py represent the movement of the link C2 in relation to the link C1 and twist $rz represent the movement of the link R2 in relation to the link C2"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003713_978-1-4614-2419-2-Figure16.2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003713_978-1-4614-2419-2-Figure16.2-1.png",
        "caption": "Fig. 16.2 Extruded view of FE model",
        "texts": [
            " where M is the moment, krot is the rotational stiffness, and y is the rotation of the joint in the direction of interest). Three springs are used for each beam to capture rotations in all three directions. The stiffness coefficient of these springs is calibrated with experimental data to obtain the highest fidelity model. The translational degrees of freedom for each connection are coupled so that there is no relative motion between translation of the beam and translation of the column at the connection. Base connections are modeled with similar rotational springs. extruded view of the model is shown in Fig. 16.2. 16 Damage Detection in Steel Structures Using Bayesian Calibration Techniques 181 values (10,000,000,000 in lb) to represent an essentially fixed connection. Since a dynamic analysis is desired, the level of discretization necessary for the mesh is determined by looking for the convergence of the natural frequencies of the mode shapes. Reference [7] describes that since the essence of finite element analysis is the discretization of elements in an attempt to capture the behavior of a continuous system, one must verify the convergence of the solution"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001763_robot.2009.5152405-Figure8-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001763_robot.2009.5152405-Figure8-1.png",
        "caption": "Fig. 8. Unexpected pheromone potential field",
        "texts": [
            " the reason why the pheromone potential field obtained from the experiment is not smooth. Although the pheromone potential field is not smooth as that in the simulation results, the experiment shows that our model can be used for pheromone potential field construction in real hardware that consider random data carrier layout and communication range problem. We did additional experiments to verify the condition showed in Fig.5 by varying \u03be in the constant \u03b6 = \u03b7 = 0.125. The value for \u03be is 0.5 if we follow Fig. 4 for stability. From Fig. 8 we can see that the constructed pheromone potential fields are in the unexpected condition if we set \u03be below or above 0.5. We can say that there is a boundary range between \u03be = 0.4 and \u03be = 0.6, where the system goes to stable if near \u03be = 0.5 and goes to unstable if far from \u03be = 0.5. In this paper, we discussed an artificial pheromone system for pheromone potential field construction. On the basis of these results, we can propose a navigation system in the constructed pheromone potential field or investigate the network properties in the constructed pheromone potential field as the next research topic"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001927_nme.1620030405-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001927_nme.1620030405-Figure1-1.png",
        "caption": "Figure 1 . Geometry of displacements of a curved element",
        "texts": [
            " This is achieved at the expense of but a slightly lengthier process of solution than that used in the previous approach. The basic method of solution of axisymmetrically loaded elastic-plastic thin shells of revolution has been presented previously24 and will not be repeated. Here discussion is confined to the main features in the development of the refined element. This is followed by three illustrative examples which show the superiority of the new formulation. GEOMETRY AND DISPLACEMENT PATTERN Geometry of a curved element The geometry of an axisymmetric curved shell element is illustrated in Figure 1. As has been its meridional curve in dimensionless co-ordinates f and 7 may be expressed as where a, = tan IS, a, = tan p4 + 472\u201d a, = - ( 5 tan pi + 4 tan p,) + $7;- 75 a, = 3(tan pi + tan &) + $($- 7;) and I = cord length d2 I d t 2 r, C O S ~ p7\u201d = = -___ Note that the curve given by equation (1) satisfies the requirements of continuity of slopes and curvatures at the nodal circles. A REFINED CURVED ELEMENT 497 Displacement pattern co-ordinates, Figure 1, is assumed over each discrete element: In this development the following displacement model, expressed in terms of local Cartesian (2) 1 u1= a l + ~ . $ + a 3 . $ 2 + a p p us= a,+a65+Ly,p+a8p where a\u2019s are the generalized co-ordinates. The number of these generalized co-ordinates is equal to the total number of internal and external degrees of freedom of the element. The six degrees of freedom at the nodes i and j , Figure 2, are the external degrees of freedom and the two at the nodes m and n are the internal degrees of freedom"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001901_1.3591479-Figure7-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001901_1.3591479-Figure7-1.png",
        "caption": "Fig. 7 Bennett mechanism",
        "texts": [
            ", it is parallel to the vector cross product BC X BF and the coordinates of E(x, y, z) are: a' = \u2014b sin 6 + l(m'n\" \u2014 n'l\" cos 6 + n'm\" sin 0) y = -b cos 6 + ( ( - \u00bb ' / \" sin 6 - n'm\" cos 6 - 1'n\") z = / + ((IT cos 6 - I'm\" sin 8 + m'l\" sin 8 + m'ft\" cos 8) (31) In equation (31), /', ))(', ft' are known from case (i) and I is a real parameter. Upon substituting for /', ill', ft.' from case (i), the working is similar to case (ii) and the result is the same: The curve is of order S and genus 1. It seems, therefore, that many of our spatial curves (discussed so far) have the same order and genus. Bennett Mechanism This mechanism has been the subject of numerous investigations, including [1, 2, 7, 12, 17, 19, 26]. The mechanism is shown in Fig. 7. It is a spatial four-link mechanism with 4 turning joints, whose axes, shown as 1, 2, 3, 4 are skew. The common perpendiculars between adjacent sets of axes intersect (at A, B', A', B). AB'A'B is called a skew isogram; opposite sides are equal (to a or b). The angles between adjacent axes (4, 1), (2, 3) are equal (to a) , and the angles between the adjacent axes (1, 2), (3, 4) are equal (to (3). Furthermore, a/b = \u00b1sin a/sin j3. If link AB' is fixed, for example, Bennett has shown that the motion of the mechanism is determinate (and possible) and in effect that the input and output rotations are related as the rotations of a pair of elliptical gears. The latter can be derived also by imposing the special proportions, which reduce the skew fourbar mechanism of case (i) to the Bennett mechanism. Referring to Fig. 7 and following Bennett, z is the line, which perpendicularly bisects the diagonals AA' and BB'. Let P' be a point on the coupler A'B and P the \"corresponding\" point on the fixed link, coincident with P' after rotating the moving link a half turn about z. As shown by Bennett, z is a generator (of one system) of a hyperboloid, Z. We investigate the coupler curve generated by P'. Since the midpoint of PP' lies on z, and P is a fixed point, P' lies on another hyperboloid, H, which is the result of \"blowing up\" Z from P in the ratio 2:1"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001164_pesafr.2007.4498051-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001164_pesafr.2007.4498051-Figure2-1.png",
        "caption": "Figure 2: Lab setup",
        "texts": [],
        "surrounding_texts": [
            "The CSA-390-1, IEEE 112-B and IEC 61972-1 follow the following main testing procedures and calculations: Temperature test: To accurately compute the losses of a motor, the temperature has to be taken into account. The tested machine is loaded at the rated load and is run until the temperature (of the stator windings) does not change for more than a degree Celsius between measurements (every 30 min). This test allows for the temperature correction of stator winding resistance and therefore correction on the stator losses. Variable load test: The machine is run at the rated conditions; the machine is loaded with six-load point ranging from 150% down to 25%. The winding temperature before the commencement of test has to be more than 10 degree Celsius for the IEEE 112 and CSA390, and 5 degree Celsius for the IEC 61972 method during testing. From this test, the stator and rotor losses are calculated (equation in Appendix A). No-load test The test is done with the motor uncoupled from the loading device (or dynamometer). The tested motor is then run with the supply at rated frequency and voltage (some motors might require a number of hours to stabilize the bearings). The IEEE and CSA-390 uses variable voltages ranging from 125% to the point where current begins to increase (due to loss of voltage and increase in slip) while the IEC uses a minimum of four voltages between 125% and 60% of rated voltage and three or more between 50% and 20% of rated voltage are used. The windage and friction, and core losses are obtained from this test (equations in Appendix A). The SLLs, in the three standards, are calculated using equation 1. This is an indirect method of acquiring SLLs. PTtrayloss (Piput - Poutput ) - Pwindageftiction - Pstator - (Pro,or ) X S - Pcore (2) The calculated SLLs are used in a linear regression method and the relationship of the regression determines the quality of testing. The variations between the standards are mainly found in the calculation of the efficiencies. These differences are discussed in Section 5 with reference to obtained efficiencies. The inclusion of equation 2 in the IEC standard is another reason for the lower efficiency values. The reduced voltages calculated from the equation produce smaller core losses values."
        ]
    },
    {
        "image_filename": "designv11_12_0002059_j.mechmachtheory.2009.04.005-Figure3-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002059_j.mechmachtheory.2009.04.005-Figure3-1.png",
        "caption": "Fig. 3. Example 1.",
        "texts": [
            " The process finishes when the final configuration is reached. Three criteria (see Fig. 2) have been introduced to obtain the cost function c\u00f0p\u00de used to select a new configuration to branch out: Table 2 Example 1. Obstacle location (in m). 0 Obstacle 1st Spherical obstacle 2nd Spherical obstacle 3rd Spherical obstacle Centre cSO 1 \u00bc \u00f0 0:85; 0:4; 0:5\u00de cSO 2 \u00bc \u00f0 0:75; 0; 0:5\u00de cSO 3 \u00bc \u00f0 0:75;0; 0:5\u00de Radius rSO 1 \u00bc 0:15 rSO 2 \u00bc 0:15 rSO 3 \u00bc 0:15 Table 1 Example 1. Configurations to be joined and path obtained (black line). See Fig. 3. Joint No. Initial configuration Joint No. Final configuration 6 0 6 0 7 0 mm 7 0 mm Table 4 Example 2. Obstacle location (in m). 0 Obstacles 1st Cylindrical obstacle 2nd Cylindrical obstacle 3rd Cylindrical obstacle Centre 1 cSO 1 \u00bc \u00f0 0:85; 0:4; 0:5\u00de cSO 2 \u00bc \u00f0 0:75;0; 0:5\u00de cSO 3 \u00bc \u00f0 0:75;0; 0:5\u00de Centre 2 cSO 1 \u00bc \u00f0 0:85; 0:4; 0:5\u00de cSO 2 \u00bc \u00f0 0:75;0; 0:5\u00de cSO 3 \u00bc \u00f0 0:75;0; 0:5\u00de Radius rSO 1 \u00bc 0:15 rSO 2 \u00bc 0:15 rSO 3 \u00bc 0:15 1st Prismatic obstacle 2nd Prismatic obstacle 3rd Prismatic obstacle Point 1 aPO 1 \u00bc \u00f0 0:7; 0:35;0\u00de aPO 2 \u00bc \u00f0 0:5;0; 0\u00de aPO 3 \u00bc \u00f0 0:5;0:3; 0\u00de Point 2 qPO 11 \u00bc \u00f0 0:7; 0:35;2\u00de qPO 12 \u00bc \u00f0 0:5; 0;2\u00de qPO 13 \u00bc \u00f0 0:5; 0:3;2\u00de Point 3 qPO 21 \u00bc \u00f0 1:5; 0:35;0\u00de qPO 22 \u00bc \u00f0 1:3; 0;0\u00de qPO 23 \u00bc \u00f0 1:3; 0:3;0\u00de Point 4 qPO 3l \u00bc \u00f0 0:7; 0:45;0\u00de qPO 32 \u00bc \u00f0 0:5; 0:15;0\u00de qPO 33 \u00bc \u00f0 0:5; 0:45;0\u00de Fig",
            " The results presented correspond to 20 examples which had different initial and final configurations with a different number of obstacles. Each example has been solved using the harmonic function as the interpolation function with three different cost functions: uniform cost (UC), greedy (G) and algorithm A* (A*) given in all 60 results. Next, we show two of the examples that have been solved. In Tables 1 and 2, we give the information about the robot configuration and the position of obstacles for example 1 (Fig. 3). In Tables 3 and 4, we give the information about the robot configuration and the position of obstacles for example 2 (Fig. 4). In the next four Tables 5\u20138, we show the numerical results corresponding to 5 of the examples resolved. The graphs correspond to all of them (1 and 2 are included). In Graph 5 it can be seen that the worst execution time is obtained when the greedy time function is used, whereas the best times correspond to the use of A* followed by the uniform cost time function. In Graph 6 it can be seen that the worst computational time is obtained when the A* and uniform cost time function is used, whereas the best times correspond to the use of the greedy time function"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003882_9781118181249-Figure2.9-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003882_9781118181249-Figure2.9-1.png",
        "caption": "Figure 2.9 Target analyte separation in a microfl uidic channel facilitated by pH - responsive magnetic nanoparticles (mNPs) under isothermal conditions. The channel contains two fl ow streams. The left stream (green) is the sample that has been preincubated with mNPs. mNP aggregation is triggered by the lower pH buffer in this sample fl ow stream. The pH of the right stream (pink) is chosen to reverse mNP aggregation. A magnet provides a suffi cient fi eld to pull the aggregates laterally into the higher pH fl ow stream. The conjugate aggregates move out of the sample fl ow stream and in to the higher pH stream, where they return to a dispersed state, carrying the bound target analyte with them. Movement of other molecules across this interface is limited by diffusion due to the laminar fl ow conditions. Reproduced from Lai et al. 44 by permission from the Royal Society of Chemistry. (See color insert.)",
        "texts": [
            " For example, the Elecsys \u00ae immunoassay system by Roche utilizes batch mode separation of magnetic microbeads conjugated to antibodies and currently has \u223c 100 approved biomarker tests worldwide. In addition to batch mode magnetic separation, continuous - fl ow magnetic separation systems have been used to achieve continuous separation of cells and proteins in microfl udic channels. Lai et al. 44 demonstrated how continuous - fl ow magnetic separation could be used in conjunction with dual pH/ temperature - responsive smart mNPs to achieve continuous separation of fl uorescently labeled proteins, as shown in Figure 2.9 . The authors combined an H - fi lter microfl uidic device with a smart pH - responsive mNP reagent that could be reversibly aggregated by lowering the solution pH. The system was designed to pull aggregated mNPs and their bound biomarker targets from one laminar fl ow stream across the laminar fl uid interface into an adjacent fl ow stream. The adjacent fl ow stream with an alkaline pH caused the mNPs to disaggregate. Instead of being caught at the surface of the magnet, the disaggregated mNPs with low magnetophoretic mobility fl owed freely by the magnet via the second laminar fl ow stream to the second fork of the H - fi lter device, enabling further downstream processing"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003702_cjme.2013.04.801-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003702_cjme.2013.04.801-Figure1-1.png",
        "caption": "Fig. 1. Mechanism of electrically driven heavy-duty six-legged robot",
        "texts": [
            " Generally speaking, the feet-ground contact forces are more valid in a static simulation analysis than in the dynamic simulation analysis in the ADAMS software. Based on the problems mentioned above and the previous basic research on the transmission modes of a joint[22], an electrically driven heavy-duty six-legged robot is considered as an example in this paper to present an analysis method for the articulated torques. The mechanism of the electrically driven heavy-duty six-legged robot is shown in Fig. 1. This paper is divided into five sections. In section 2, an analysis model is used to determine which leg bears the maximum normal contact force on a slope. In section 3, based on the relevant tripod gait, the formulas of classical mechanics and the MATLAB software are employed to solve the normal contact forces, the foothold frictions, and the static torques of the joints. Under hypothetical parameters for the robot, variable tendency charts and extreme curves are obtained for the static articulated torques with changes in the joint angles"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003713_978-1-4614-2419-2-Figure28.8-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003713_978-1-4614-2419-2-Figure28.8-1.png",
        "caption": "Fig. 28.8 Modal analysis test setup",
        "texts": [
            " The starting value of the orientation is calculated with the laser tracer oriented to the z-axes at the origin i.e. (a0,b0,g0) \u00bc (0,0,atan( y0, x0)). We calculate the arctangent-function range of ( p,p). Additionally, we set constrains for angles a and b to be within the interval of ( p/2; p/2) as the used TI mounting platform does not allow higher inclinations anyway. 28 Modal Testing Using Tracking-Interferometers 291 To demonstrate the performance of the tracking interferometer we use the machining center in Fig. 28.8 for a modal test. The machine tool is approximated by only eight points as shown in the Fig. 28.8. Points 1, 2 and 3 represent the traveling column. Points 4, 5 and 6 are attached to the spindle housing. The table is represented as points 7 and 8. The modal analysis was conducted with both accelerometers and the TI. We use a hydraulic exciter to apply forces relatively between the tool and the work table. The results are shown for the first two modes at 17 and 34 Hz. Higher modes show poor repeatability and a large mismatch between tracking-interferometer and accelerometer measurements in this setup due to vibrations coupling into the tracking-interferometer\u2019s platform. This will be discussed in the outlook. The test shows specifically that the coordinate transformation works properly in three dimensions, too. The first mode shown in Fig. 28.8 is an oscillation vibration of the entire machine around the y-axes. The machine table and the spindle vibrate in phase but with different amplitudes. The highest amplitudes are detected at the top of the travelling column. The main relative displacement between the tool center point and the worktable in the second mode at 34 Hz is in x-direction as well. The tool center point and the worktable vibrate in antiphase. Though still in x-direction, this vibration mode is caused by a rotation of the traveling column and the z-slide with the spindle around the z-axes"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003352_imece2013-62193-Figure7-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003352_imece2013-62193-Figure7-1.png",
        "caption": "Fig. 7 Calculated temperature distribution in oC, molten pool area, and 3-D view of activation element for generation the inclined molten surface.",
        "texts": [
            " In this study, a 3-D transient heat transfer model was developed to estimate the temperature history of depositing H13 steel on mild steel A36 by HPDDL. The modeled temperature was coupled with thermokinetic empirical relations to predict the hardness distribution in the coating layer. In order to validate the computed results for temperature history and hardness distribution, a series of experiments were performed. The simulation result for temperature distribution in the process domain is shown in Figure 7. This figure presents the shape of the molten pool based on elliptical cylinder model for the activation element (Eq.7). The isotherm boundary at the melting temperature of a clad H13 (1480 oC) is regarded as the molten pool edge, as shown in Fig. 7.presented by dashed lines. The minimum discrepancies between modeled and measured results were achieved when the thermal conductivity was enhanced by multiplying it with a factor of 2.6 along the x and y directions at temperatures higher than the melting point. The peak temperature at the position of TC2 (550 oC) is higher than TC1 (500 oC), and it could be attributed to laser energy accumulation that increased the substrate temperature gradually during LC. 7 Copyright \u00a9 2013 by ASME Downloaded From: http://proceedings"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003645_optim.2012.6231768-Figure8-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003645_optim.2012.6231768-Figure8-1.png",
        "caption": "Fig. 8: Testbench. A load machine is coupled by a torque transducer to the PMSM",
        "texts": [
            " model 2 (see section III) from the r identified inductances, again a SLE has to be solved:                                                                                                                                                                               0 0 0 \u02c6 \u02c6 \u02c6 0,0 0,0 0,0 , , , , , , , , , , 2, 1, , 2, 1, , 2, 1, 2 1 ,, 2,2, 1,1, ,, 2,2, 1,1, ,, 2,2, 1,1, rdq dq dq rq q q rd d d l Eco q d rqrdLdq qdLdq qdLdq rqrdLq qdLq qdLq rqrdLd qdLd qdLd L L L L L L L L L B B B iiB iiB iiB iiB iiB iiB iiB iiB iiB           (39) The absolute flux linkage and co-energy values can not be identified from the measurement and are forced to be zero at idq = 0 in eq. (39) in order to get a matrix with full rank. Experimental results are shown in section VII. VI. IDENTIFICATION BY STATIONARY OPERATION ON A TESTBENCH ANALYSING VOLTAGE AND TORQUE The PMSM is operated on a testbench and driven by a load machine. A torque transducer is used to measure the shaft torque (Fig. 8). The test machine is still operated in field oriented current control. Special requirements for the dynamics of the control are not necessary: All measurement data will be taken, after the currents have reached stationary values. Different values of the motor speed are forced by the load machine. Thus, each measurement point is defined by the currents id, iq and the speed n. Additionally, if required, the motor winding temperature can be considered. For each point [id iq n]T, the recorded quantities are:  Average value of the shaft torque T, measured by the torque transducer"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000801_gt2007-28042-Figure9-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000801_gt2007-28042-Figure9-1.png",
        "caption": "Figure 9: High speed brush seal test rig",
        "texts": [
            " This phenomenon is one of the concerns in brush seal designs, especially for high pressure environments, and was the subject of many studies such as Wood and Jones [20], Arora et al. [23], to name a few. In all of these studies, torque measurements were utilized to establish the effect of pressure on tip force increase, not to estimate the seal tip force at a set pressure. In order to study the effect of pressure on the seal tip force, as well as friction and heat generation properties of a brush seal, a high speed seal test rig, shown in Figure 9, capable of reaching up Copyright \u00a9 2007 by ASME 5 ?url=/data/conferences/gt2007/71687/ on 05/01/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use Down to 40000 rpm, has been designed for and utilized in this study. A detailed description of the test rig is provided by Demiroglu [19]. The torque tests were conducted for seals #1, #5 and #8 (See Table 3 for details). Two labyrinth teeth seals with radial clearance of 0.025 mm were used as dummy seals to create a pressurized cavity without touching the rotor"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000802_1.2908921-Figure10-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000802_1.2908921-Figure10-1.png",
        "caption": "Fig. 10 Contact forces between the ro",
        "texts": [
            " Figure 9 presents a vertical displacement of the rotor as a function of time at the locations of the retainer bearings in example g . In the figure, it can be seen that the cylindrical mode of the whirling has occurred due to the misalignment of the retainer bearings. From the simulated scenarios, it can be noticed that the horizontal misalignment causes the cylindrical whirling motion of the rotor. In the case of the horizontal misalignment, the first contact of the rotor with the misaligned retainer bearing will produce a force component in the rotor transverse direction. This behavior is depicted in Fig. 10 a , in which the forces Frx and Fry are the components of the contact force Fr. The force Frx in the rotor transverse direction increases the rotor motion in the circumference direction. As a consequence, one end of the rotor may experience backward whirling motion. As a result of the backward whirling motion in the one end of the rotor, translational kinetic energy of the rotor increases. This, in turn, leads to backward whirling motion at both ends of the rotor. Finally, friction forces compel the Table 5 Misalignments of the retainer bearings in the simulations Case Retainer bearing 1 Retainer bearing 2 X-direction Y-direction X-direction Y-direction a 0 m \u2212180 m 0 m 180 m b 0 m \u2212190 m 0 m 190 m c 0 m \u2212210 m 0 m 210 m d 0 m \u2212220 m 0 m 220 m e \u2212180 m 0 m 180 m 0 m f \u2212190 m 0 m 190 m 0 m g \u2212210 m 0 m 210 m 0 m h \u2212220 m 0 m 220 m 0 m rotor into a cylindrical whirling motion see Fig. 9 . In the case of Transactions of the ASME ata/journals/jotre9/28757/ on 02/06/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use t t t B s t t ring J Downloaded Fr he vertically misaligned retainer bearing see Fig. 10 b , the ransverse contact force component is negligible. As a result, in he simulations, the behavior of the rotor remains oscillatory. ased on the above presented results, it can be concluded that the ufficiently high friction force effecting on the rotor can lead to he whirling motion of the rotor. This is a consequence of the fact hat the friction force increases the total transversal force compo- ournal of Tribology om: http://tribology.asmedigitalcollection.asme.org/pdfaccess.ashx?url=/d nent of the rotor during the contacts"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001758_tmag.1972.1067544-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001758_tmag.1972.1067544-Figure1-1.png",
        "caption": "Fig. 1. Coordinate system.",
        "texts": [
            "NTRODUCTION A one-dimensional Bloch wall is assumed to exist in a bulk specimen of ferromagnetic material described by a uniaxial anisotropy constant K , an exchange constant A, and a saturation magnetization M . With the coordinate system shown in Fig. 1, the following torque equations may be derived from the vector equation of motion containing a viscous damping parameter a!: + 2K sin 8 cos 0 + 2A sin e cos e (&q2 - 2A v2e. ( 1b) The fields Hy and Hz permit the introduction of stray and applied fields and y is the gyromagnetic ratio. The z axis is the easy direction. In the absence of an applied field in the hard or y direction, H = - - sin 8 sin cp. M Po The following trial function is introduced: In (tan e / 2 ) = c2 ( t ) [y -S t V ( T ) dT] ( 2 4 and cp is assumed to be a function of time only and not a function of y during the transient as well as during steady-state velocity conditions",
            " Feldtkeller [6] recognized the initial stages of this nonlinearity and developed an approximate heoretical description although a previously developed theory was available [ 71 . In Walker\u2019s analytic solution of a planar Bloch wall in a bulk material, the steady-state wall velocity is of the form with 0 < < 1, where h, = po H,/Ms is the normalized drive field, hk = po H k / M , is the normalized anisotropy field (Hk = 2K/Ms), and M s is the saturation magnetization. The exchange and anisotropy constants are A and K, the phenomenological wall damping parameter is 01, and y is the gyromagnetic ratio. The velocity-field relationship of (1) is shown in Fig. 1 for various anisotropy conditions. For low hk materials a prominent peak in the velocity together with a region of negative differential mobility ( d V / d H , < 0) is predicted. The coordinates of the velocity peak are given as hz IVp k = o r h k 1 / 4 ( 1 + h k j ~ [ ( l + h k ) % - h k l / 2 ] . (2b) The negative differential mobility region occurs for fields hZIVpk <hz GyO1. 1 With typical Permalloy values, M s = 1 Wb/m2, a = 0.01, A = lo-\u2019 \u2019 J/M and K = 200 J/M3 in a sample calculation Manuscript received May 22, 1972; revised September 18, 1972"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000685_j.triboint.2007.02.012-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000685_j.triboint.2007.02.012-Figure2-1.png",
        "caption": "Fig. 2. Influence factor of surface displacement.",
        "texts": [
            " Thus, the influence factor Dc ji i0j;jj j0j of surface displacement of coated film is calculated by using a threedimensional axis-symmetrical finite element method to simplify the calculations [17], and then surface displacement is calculated from Eq. (5b) in our paper. In the analysis of the influence factor, calculations are performed by using the model of unit force load on the equivalent circles with rectangular region to analyze by using threedimensional axis-symmetrical finite element method. Dc ji i0 j;jj j0 j is given by the displacement at point (xi, yj) which is located at a distance of r from the point \u00f0xi0 ; yj0 \u00de where unit force load as shown in Fig. 2. In the analysis of stresses in the coated film and the substrate, the normal stress components sx, sy, sz and the shear stress components txy, tyz, tzx into the elements of coated film and the substrate are calculated by using the three-dimensional finite element analysis [18]. The stress components are calculated from the displacement of the coated film and the substrate. The maximum shear stress in the xz plane, denoted tmax, is obtained by the Eq. (7) [19]. tmax \u00bc ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi \u00f0sx sz\u00de 2 2 \u00fe t2zx s "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003961_s1068366613040119-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003961_s1068366613040119-Figure2-1.png",
        "caption": "Fig. 2. Calculation model: 1. antifriction element, 2. rigid shaft.",
        "texts": [
            " General theses of the algorithm for the solution of wear contact problems using the method of triboele ments with ANSYS package are discussed in [6]. In order to solve wear contact problems in a spatial set ting, one should take into consideration that, in this case, a cubic spline of the worn surface is constructed. At each iteration of wear determination, a cubic spline of the worn surface is constructed, the geometry of which takes into account the wear of the previous step. The calculation scheme allowed us to build a parame terized calculation model in the preprocessor of the ANSYS package (Fig. 2). The contact interaction of elements of the tribo joint was simulated as a rigid\u2013compliant contact. Shaft 2 (Fig. 2), which is rigid, was taken as a target surface. Antifriction layer 1 was taken as the contact surface. On the contact surface of the antifriction layer, the geometric position of the triboelements was determined. SOLID186 twenty node spatial elements were used to generate a mesh of the finite element model of the antifriction layer. Surface\u2013surface con 1R Keywords: plain bearing, wear, profile formation, simulation, triboelement method DOI: 10.3103/S1068366613040119 JOURNAL OF FRICTION AND WEAR Vol"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000140_s00542-007-0415-0-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000140_s00542-007-0415-0-Figure1-1.png",
        "caption": "Fig. 1 Schematic of impact between the slider and disk: a contact of slider corner with disk; b simplification of the slider corner as a sphere",
        "texts": [
            " The contact between slider corner and disk has also been experimentally observed in slider-disk contact in dynamic load/unload processes (Liu and Ma 2003). Moreover, the slider corners have certain degree of radius of curvature, which could be artificially increased to reduce the contact stress and damage (Suk et al. 1998). The size and mass of the slider are significantly smaller than those of the disk, therefore the contact between the slider and the disk could be treated as a sphere impacting into a semi-infinite plane, as shown schematically in Fig. 1. In this study, we consider the case of a spherical corner with a radius of 1\u201310 lm, which is a range for a typical sharp corner and a \u2018\u2018rounded\u2019\u2019 corner, respectively (Suk and Gillis 1998; Polycarpou et al. 1999). The analysis of a spherical slider corner impacting a disk, presented in this paper, could readily be implemented in an air-bearing simulation thus being able to predict the exact pitching and dynamic motions of the slider due to the slider corner impact. In this paper, the mechanical response of slider-disk impact was first considered using Hertzian elastic impact theory, with emphasis on the mean contact pressure and the difference between NiP/Al substrate disk and glass substrate disk, at a practical range of impact velocities and slider corner radii"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001257_1.3071973-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001257_1.3071973-Figure2-1.png",
        "caption": "Fig. 2 Geometrical model of a laser textured surface: \u201ea\u2026 schematic of the model, \u201eb\u2026 zoom in on a group of dimples, and \u201ec\u2026 a single column of dimples in the x1 direction",
        "texts": [
            " In oth cases, during sliding the mating surfaces are completely eparated by a fluid film with a thickness distribution h x1 ,x3 . It should also be noted that dimples on the moving textured urface enter and leave periodically the boundaries of the fluid lm zone, which may lead to transient elastomer deformation and ossible squeeze action close to the elastomer edges. However, ecause a relatively large number of dimples between 10 and 15 re present in the fluid film zone at any given time, this edge ffect can be neglected as was shown in Refs. 10,16 . Figure 2 presents the LST details. The dimples are uniformly istributed on the counterpart surface see Fig. 2 b with an area l2 l1 hp x3 x2 x1 c pa pa Undeformed elastomer Rigid LST counterpart Rigid foundation Deformed elastomer U \u03b4(x1,x3) h(x1,x3) ig. 1 Schematic of the model showing the initially undeormed elastomer \u201edashed line\u2026 at zero sliding velocity, and the nal deformed elastomer \u201esolid line\u2026 at full sliding velocity ensity Sp. These dimples are regularly arranged in axial columns 21503-2 / Vol. 131, APRIL 2009 om: http://tribology.asmedigitalcollection.asme.org/ on 01/28/2016 Terms as shown in Figs. 2 a and 2 c . Curvature, in the case of a rigid LST rod, can be neglected due to the small clearance in the x2 radial direction compared with the radius of the rod. In this case, the problem domain is infinitely long in the x3 circumferential direction and has finite length in the x1 axial direction from x1 =0 to x1= l1 see Fig. 1 . Each dimple is located in the center of an imaginary square cell of sides 2r1 2r1 see Fig. 2 b and has the shape of a spherical segment with a base radius rp and maximum depth hp. The dimple area density and dimple base radius are related to the square cell side by r1 = rp 2 Sp 1 Modeling of the SEHL problem requires simultaneous solution of the hydrodynamic lubrication between the sliding surfaces and the deformation of the elastomer see Ref. 16 . 2.1 The Hydrodynamic Lubrication Model. All the assumptions and governing equations for the hydrodynamic model used in Ref. 16 are adopted here too. A special subroutine was written using finite difference method to solve the Reynolds equation for the hydrodynamic pressure distribution in the fluid film assuming viscous incompressible Newtonian fluid in laminar flow squeeze effect is neglected . x1 h3 p x1 + x3 h3 p x3 = 6 U h x1 2 Because of the periodicity of the surface texturing in the x3 direction see Fig. 2 b and the symmetry of each axial dimple column of width 2r1 about its longitudinal axis x3=0 see Fig. 2 c , it is sufficient to consider the pressure distribution within just one-half of a single dimple column confined between x3=0 and x3=r1 . Thus, the boundary conditions for the Reynolds equation are as follows. Symmetry of the pressure distribution about x3=0, periodicity, and smoothness of the pressure distribution in the x3 direction p x3 x1,x3 = 0 = p x3 x1,x3 = r1 = 0 3a Ambient pressure at the elastomer axial boundaries see Fig. 1 Transactions of the ASME of Use: http://www.asme.org/about-asme/terms-of-use T t t t w d t s w a b T n d c t c c t n e t w f t t a G d t w t = d t t t J Downloaded Fr p x1 = 0,x3 = p x1 = l1,x3 = pa 3b he Reynolds cavitation condition was applied in potential cavi- ation areas assuming a cavitation pressure pa"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001114_17543371jset42-Figure9-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001114_17543371jset42-Figure9-1.png",
        "caption": "Fig. 9 3D visualization of the global model kinematic. Picture from 3D animation obtained with the ADAMS postprocessor",
        "texts": [
            " This fact shows clearly the limit of the harmonic approach of the radiation problem for rowing and the convolution approach will be preferred in the future. Since the coupled hydrodynamic coefficients were neglected and the influence of the secondary movements on the resistance was not taken into account, the boat velocity is weakly modified by the secondary movements. This is another point which must be accounted for. The powerful graphic interface of ADAMS is a feature that has to be discussed with the coaching staff and rowers. For example, Fig. 9 shows an instantaneous picture picked up during an animation and shows the possibilities offered by the ADAMS environment to render visually the calculated results, particularly the well known and typical absolute motion of the blade centre [17]. The level of agreement between the computed results and the measured data gives credence to the methodology and the realism of the model. The release degree of freedom in the model is not a mechanical problem when a powerful mechanical modeller such as ADAMS-LifeMOD is used but it Table 4 Mean velocity versus lower torso rotation angle Mean boat velocity (m/s) 3"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002400_1.4006324-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002400_1.4006324-Figure1-1.png",
        "caption": "Fig. 1 Face-contact model used for loaded tooth contact analysis of gears",
        "texts": [
            " We have also found that the effects of the web angle on the tooth contact stresses and the root bending stresses of the thinrimmed right web gears are different from those of the left and center web gears. This difference can be used to improve the tooth contact patterns of gears with misalignment errors. In previous researches, face-contact models of contacting teeth have been suggested for conducting LTCA, deformation, and stress calculations of a pair of spur gears [11\u201315]. This model is also used here to conduct LTCA, deformation, and stress calculations of thin-rimmed inclined web gears. Figure 1 shows the face-contact model used for the LTCA of a pair of spur gears. This face-contact model is different from a line-contact model in that the contact of the teeth is assumed to be on a reference face with a contact width WH on the tooth surface as shown in Fig. 1(a). The line-contact model only assumes that the tooth contact is on the geometric contact line as shown in Fig. 1(a). The face contact of the contacting teeth on the reference face is considered to be the contact area of many pairs of points on the reference faces as shown in Fig. 1(b). In Fig. 1(b), (k\u2013k0) is an optional pair of points assumed to be in contact with each other on the reference faces. k is the point on the reference face of gear , and k0 is the point on the reference face of gear . ek is the gap between the pair of points (k\u2013k0). This face-contact model is in agreement with the real contact state of teeth with deformations; therefore, this model can analyze the real surface contact stress as well as the root bending stress of the contact teeth, especially in the case where the gears have machining errors, assembly errors, and tooth modifications",
            " 2, when these gears have both straight webs and inclined webs, and they are engaged with the solid mating gear at the highest point of the single pair tooth contact. The tooth contact stresses are calculated under a torque load of 294 N m for all of the cases in the paper. 4.1 Tooth Contact Stress Distributions of the Thin-Rimmed Left Web Gears. At the time when the webs are located at the left side of the tooth, the tooth contact stresses are calculated. Figure 4 shows the results of the tooth contact stresses of the left web gears. In Fig. 4, the abscissa is the tooth longitudinal dimension and the ordinate is the contact width WH as shown in Fig. 1. The web position and the geometrical contact line are also shown in Fig. 4(a). Figures 4(a)\u20134(d); are the contour line graphs of the contact stresses of the gears when the web angles are 0, 15, 30, and 45 deg, respectively. The maximum contact stresses of the teeth are indicated in Fig. 4, and the resulting relationship between the maximum contact stress and the web angle is summarized in Table 1 and is shown in Fig. 5. In Table 1 and Fig. 5, the contact stress when web angle is 0 deg is used as the standard value, and the rate of change is defined as the change in the contact stress when the web is inclined relative to this standard value"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000020_jjap.45.4241-Figure9-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000020_jjap.45.4241-Figure9-1.png",
        "caption": "Fig. 9. Configuration of nonliquid-junction type reference electrode.",
        "texts": [
            " In the case of the liquid-junction-type, the cross section of the liquid junction region is responsible for the lifetime and response time. This can be achieved by careful control of the cross section with microfabrication technology. The nonliquid-junction type was fabricated only by placing a gas-permeable membrane on the electrode. However, both membrane quality and thickness are important to provide a long lifetime and a quick response. In this work, we have studied the fabrication of the non liquidjunction type reference electrode with PVC (polyvinyl chloride) as the membrane for our ion sensor. Figure 9, it shows the structure of the new prototype nonliquid-junction type reference electrode. In the fabrication process, SiO2/Si is used as a substrate, an 800 nm Ag layer is introduced and sputter-deposited on adhesive 50 nm Cr layer and chlorinated with 0.5M FeCl3 to form a Ag/ AgCl electrode, which suggests a high sensitivity in comparison with the mixture solution of K2Cr2O7 and HCl. Figure 10 shows the time dependence of the output voltages of the micro-reference electrode in 0.1M PBS (phosphate buffer solution) for various PVC film thicknesses"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000597_978-3-540-73129-0-Figure6.126-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000597_978-3-540-73129-0-Figure6.126-1.png",
        "caption": "Fig. 6.126. Micromotor with integrated gear box: a schematic of the micromotor, b assembled planetary gear system [329]",
        "texts": [
            " The motors have been brought to spin at rates higher than 10 000min\u22121. The most severe problem is the occurrence of friction and high wear in the bushing, severely limiting the practical lifetime of such a motor. Up to now, these motors have mainly been used to demonstrate the capability of the technology but are still waiting for real world applications. The Institut fu\u0308r Mikrotechnik, Mainz in Germany, has developed a highly reliable micromotor [329]. A synchronous motor scheme with a rotating permanent magnet has been employed (Fig. 6.126a). The micromotor featur- ing an outer diameter of 1.9mm exhibits a maximum measured torque of 5 \u00b5Nm in continuous operation. The lifetime is considerably longer than 6 months at 10 000min\u22121. A micro gear box of the Wolfrom type has been developed using individually modified involute tooth profiles, whose components are fabricated in metal and polymer materials by means of the LIGA process (Fig. 6.126b). The motor with integrated gear box increases the available torque and helps to open new fields of application \u2013 for example, communication and information technology as well as consumer electronics. Hybrid concepts make use of the most suitable material and the most appropriate process in the fabrication of each component. Such a heteromorphic construction is typical of many microsystems and also demonstrates a broad need for efficient construction, connection and microassembly techniques, and standardized electrical and mechanical interfaces"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000399_s10999-008-9077-z-Figure8-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000399_s10999-008-9077-z-Figure8-1.png",
        "caption": "Fig. 8 Characteristics of a cubic spline segment",
        "texts": [
            " On the other hand, only C0-continuity of the contact surface violates the assumptions of smoothness, which are essential to the proofs of uniqueness and convexity of contact problems (Kikuchi and Oden 1988). To overcome these inconsistencies, contact problems commonly involve a disproportionately fine mesh at the vicinity of the contact region. This has motivated the use of splines passing through the FE nodes and possessing at least C1continuity. The resulting normal vector is uniquely defined at all point belonging to the selected surface (Fig. 7). 4.1 Cubic spline interpolation of surfaces Figure 8 shows a smooth cubic spline segment connecting two points in a 2D space. Assuming a parametric representation, the coordinates of any point on this curve can be expressed in terms of an independent variable u, such that: x u\u00f0 \u00de \u00bc x u\u00f0 \u00de y u\u00f0 \u00de ; 0 u 1 \u00f029\u00de For each curve segment a, the interpolation function passes through the end points p0 and p3, while the intermediate points p1 and p2 dictate the shape of the curve (Fig. 9). The segment can be described as follows: xa u\u00f0 \u00de \u00bc X3 i\u00bc0 B3 i u\u00f0 \u00de pa i \u00f030\u00de where Bm i u\u00f0 \u00de \u00bc m i ui \u00f01 u\u00dem i \u00f031\u00de are the Bernstein polynomials"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001531_s11012-009-9232-0-Figure10-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001531_s11012-009-9232-0-Figure10-1.png",
        "caption": "Fig. 10 Intermediary position of the mechanism",
        "texts": [
            " The consequence is that, when the joints O1, O3 and O5 are constrained to make displacements to the long of the heights of the triangle O1 1O1 3O1 5 , the other three joints O2,O4 and O6 are always in perpendicular planes to the horizontal plane. More precisely, when b = 0, all the six joints are in the initial position O1 1O1 2O1 3O1 4O1 5O1 6 in the hori- zontal plane, when the input angle \u03b8 has zero degrees. When b increases, which means when the three \u201cactuated\u201d joints O1, O3 and O5 follow the heights of the initial triangle O1 1O1 3O1 5 , the other three joints O2, O4 and O6 are in perpendicular planes to the horizontal plane, and the input angle \u03b8 increases. In Fig. 10 the mechanism is represented in an intermediary position. For the clarity of figure, the joint number 3 is not represented. The distance between the horizontal plane and the translated plane is noted H = OT . Continuing the movement, the mechanism reaches its final position O3 1O3 2O3 3O3 4O3 5O3 6 , where all the six joints are again in the fixed horizontal plane but in a different position considering the initial one. In this final position, the input angles has its maximal value \u03b8 = 2\u03c0/3. In the triangles O2 1O2 6M1 and O2 1OM1 we can write the next expression: O2 1M1 = a \u00b7 cos \u03b8 2 = 2a \u2212 b \u221a 3 2 (27) We obtain the dependence of the variable distance b according to the input angle \u03b8 : b = 2a\u221a 3 \u00b7 ( 1 \u2212 cos \u03b8 2 ) (28) We can obtain the variation between the two planes, the fixed one and the translated plane [15]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000871_jsen.2008.917482-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000871_jsen.2008.917482-Figure1-1.png",
        "caption": "Fig. 1. Schematic diagram of the homemade cell.",
        "texts": [
            "020 g of the reagent in 100 ml of water\u2014ethanol solution (1:1) in a 100-ml standard flask. A UV-Vis-double beam spectrophotometer, Jasco Model V-570, with 2 nm spectral resolution was used for recording the 1530-437X/$25.00 \u00a9 2008 IEEE visible spectra and for absorbance measurements. The sensing membrane was placed and fixed in a disposable plastic cuvette and all measurements were performed in a batch mode. A homemade cell holder (constructed from plastic) was used with a special frame with a size of 9 mm 40 mm, as shown in Fig. 1. The film was placed between the two stabilizers (in the slit) and the cell is directly mounted in the spectrophotometer. A Shimadzu (Model AA-680G) furnace atomic absorption spectrometer furnished with an Hg-hollow cathode lamp was used. The instrument was set at a wavelength of 228.8 nm and all of other parameters were adjusted according to the standard recommendation as: palladium solution (45 g/mL) as a modifier for determination of Hg(II), with heating program as: 1) 20 s as a drying time at 60 C; 2) 3 s as an atomization time at 1450 C; and 3) 2 s as a cleaning time at 2500 C",
            " The immobilized indicator on the triacetylcellulose membrane was prepared according to the following procedure: the transparent triacetylcellulose membranes were produced from waste photographic film tapes that were previously treated with commercial sodium hypochlorite for several seconds in order to remove colored gelatinous layers. The films were treated with a solution of PAN for 4 h at ambient temperature. Then, the membranes were washed with water for removing loosely trapped indicator. The prepared sensor had a thickness (path length) of 0.14 mm. The membranes with immobilized indicator were put vertically inside the cuvette using a special frame. The size of its opening was 8.5 mm 35 mm (Fig. 1). The control sample against which the measurement was performed consisted of triacetylcellulose film treated in the same way but without indicator. The control sample was put in the same way inside the cuvette using a frame of the same size. The membrane exposed to a buffer solution, acetate buffer (pH 6.0), dried in air, was used as a blank. The same membrane exposed to sample solution at the same pH (pH 6.0) for 15 min and washed with water, dried in air and was used as a sample. Then, the absorbance of the sample membrane was measured against the blank at 560 nm"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002739_icra.2011.5980299-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002739_icra.2011.5980299-Figure2-1.png",
        "caption": "Fig. 2. The execution of power-grasp behavior and the final object range image. The arrow shows the corresponding approach direction (\u03b1).",
        "texts": [
            " 1 (a)), equipped with a range camera, and its physics-based simulator is used as the experimental platform (Fig. 1 (d)). The robot platform consists of a five fingered 16 DOF Gifu robot hand and 7 DOF PA-10 robot arm. For robot perception, SwissRanger SR-4000 infrared range camera, with 176\u00d7144 pixel array, 0.23\u25e6 angular resolution and 1 cm distance accuracy was used. Along with the range image, the camera also provides grayscale image of the scene that enables us to differentiate the robot hand from objects. The simulator (Fig. 2, 3), developed using the Open Dynamics Engine (ODE) library, is used during the exploration phase. The range camera is simulated by sending a 176\u00d7144 ray array from camera center with 0.23\u25e6 angular intervals. A. Interactions What type of interactions the robot can perform on the objects depend on the diversity of its behavior repertoire. In this work, five different behaviors, that are assumed to be learned in a previous developmental stage, are used to manipulate the objects in the environment"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003117_acc.2013.6580066-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003117_acc.2013.6580066-Figure1-1.png",
        "caption": "Fig. 1. Geometry of virtual structure formation in an inertial frame",
        "texts": [
            " The resulting formation is a rigid equilateral polygon, and its center represents the motion of the entire formation. The polygon center can be regarded as a virtual leader of the UAV team. In this work, this virtual UAV leader, denoted as UAV0, is described by (1) as well, and its acceleration, linear speed and turning rate are denoted by u0, v0 and \u03c90, respectively. The motion of the UAV team as a rigid formation can be controlled by specifying u0 and \u03c90 for a designated mission, such as path following and target tracking. The geometry between UAV0 and UAVi in the formation is depicted in Fig. 1, and the relative motion between them can be derived as l\u0307i = vi sin(\u03c8i \u2212 \u03bbi)\u2212 v0 sin(\u03c80 \u2212 \u03bbi) (2) \u03bb\u0307i = \u2212vi cos(\u03c8i \u2212 \u03bbi) + v0 cos(\u03c80 \u2212 \u03bbi) li +R (3) where \u03bbi = \u03b8i \u2212 \u03c0/2 and \u03b8i \u2208 [0, 2\u03c0). In this work, it is assumed that each UAV can sense its states and has the state information of the virtual UAV. The virtual UAV can be maintained by one UAV within the group and broadcast to the rest of the group. It can also be maintained by a ground station and transmitted to the UAV group. The variables \u03bbi (i = 1, 2, "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002281_robio.2012.6491125-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002281_robio.2012.6491125-Figure2-1.png",
        "caption": "Fig. 2. The quadrotor system",
        "texts": [
            "1, if the speed of each rotor increases, the quadrotor achieves lifting; varying the speed of rotor 1 and rotor 3 could realize pitching while changing the speed of rotor 2 and rotor 4 could achieve rolling. Yaw motion is more subtle as it results from the difference of the clockwise torque and anticlockwise torque. B. The Small Quadrotor Prototype for Experiments This section describes the composition of the quadrotor prototype for experiments and its hardware of the control system. The quadrotor experimental system is composed of three parts: R/C command sender device, the quadrotor aircraft with on board control system and ground station, as shown in Fig. 2. R/C is a 6 channels wireless remote controller working at frequency of 72MHz, which can be assigned control commands such as pitch, roll, yaw, thrust and emergency stop, respectively. The on board control system of the quadrotor captures the PPM signals from the R/C receivers to get the desired inputs and samples the inertia sensor signals to get the attitude of the aircraft. The control algorithm is implemented to keep the attitude stable. Ground station is designed to test the control system and save the flight data"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003218_j.matcom.2011.07.010-Figure5-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003218_j.matcom.2011.07.010-Figure5-1.png",
        "caption": "Fig. 5. Modeled structure of the beam.",
        "texts": [
            " For Case 2, the estimation error converges toward zero, because the change of parameter q is made to get a smaller cost function (J(qi, k3) < J(qii, k3)) where the speed of convergence is even faster than with t o t t 4 s a w o d T m t i t c v he parameter q(k3) = qii. This discussion can be extended to the general case of API-Observer with uncertain number f parameters q for choosing. Considering the shaded area in Fig. 4 as all possible values of J, the final trajectory of he cost function from an API-Observer with uncertain number of parameters q for choosing will be the gray solid line hat ensures the convergence of the estimation error toward zero. . Simulation example An elastic beam example [8] shown in Fig. 5 is given here to illustrate the proposed approach. The elastic beam system is modeled using Finite Element Method. The length of each element is 98 mm, the crossectional area is 125 mm2. The displacements zi and the angles \u03b8i (i = 1, . . ., 5) as well as the corresponding velocities nd angular velocities are considered as the system states. The system model can be described in the standard state space form x\u0307(t) = Ax(t) + b(t) + Nd(x, t), (31) y(t) = Cx(t) + Chh(t), (32) ith the state vector x(t) = [ z1 \u03b81 \u00b7 \u00b7 \u00b7 z5 \u03b85 z\u03071 \u03b8\u03071 \u00b7 \u00b7 \u00b7 z\u03075 \u03b8\u03075 ]T , (33) ne known input b(t), one unknown input d(x, t) acting at the moment of contact between vibrating beam and contact evice"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000546_pes.2007.386135-Figure4-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000546_pes.2007.386135-Figure4-1.png",
        "caption": "Fig. 4. Planes in suspension force generation.",
        "texts": [
            " In the bearingless motors with full-pitch motor winding, the principle of suspension force generation is well known as written in [2]. This paper explores the principle of a bearingless motor with a short-pitch winding. Fig. 3 shows the cross section of 12 slots, 8-pole bearingless motor with brushless DC structure as described in Fig. 2. The principle of rotational torque generation is the same as that of conventional brushless DC motors. In this type bearingless motor, three u-, v- and w-planes are defined as shown in Fig. 4, and a plane is selected depending on the rotor angular position to generate the suspension force [5]. From Fig. 2 and Fig. 4 (a), the MMF directions of suspension windings su1 and su2 are defined as u1- and u2-axes, and the u-plane is composed by the u1- and u2-axes. The u1- and u2-axes are corresponding to the x- and y-axes, respectively, as shown in Fig. 3. Similarly, the MMF directions of suspension windings sv1 and sv2 are defined as v1- and v2-axes, and the v-plane is composed by the v1- and v2-axes, which are rotated by 120 [deg] in the u1- and u2-axes as described in Fig. 4 (b). The MMF directions of suspension windings sw1 and sw2 are defined as w1- and w2-axes, and the w-plane is composed by the w1- and w2-axes, which are rotated by 240 [deg] in the u1- and u2-axes as described in Fig. 4 (c). Fig. 3 shows the suspension force when the rotor is positioned during 0 and 15 [deg] and the suspension windings sv1 and sv2 are excited using the v-plane. The flux densities at the air-gaps 1 and 2 are unbalanced by the superposition of suspension flux on the PM field excitation; as a result, the suspension force Fv1 is generated in the negative direction in the v1-axis. The flux densities at the air-gaps 3 and 4 are also unbalanced; as a result, the suspension force Fv2 is generated in the negative direction in the v2-axis",
            " In the angular position during 0-15 [deg] and 45-60 [deg], the suspension windings sv1 and sv2 are excited in the vplane. The current commands isv1* and isv2* in the suspension windings sv1 and sv2 are determined from the suspension force commands Fx* and Fy* as, \u23a5 \u23a5 \u23a6 \u23a4 \u23a2 \u23a2 \u23a3 \u23a1 \u23a5 \u23a6 \u23a4 \u23a2 \u23a3 \u23a1 \u2212 =\u23a5 \u23a6 \u23a4 \u23a2 \u23a3 \u23a1 =\u23a5 \u23a6 \u23a4 \u23a2 \u23a3 \u23a1 \u2217 \u2217 \u2217 \u2217 \u2217 \u2217 y x v v sv sv F F K F F K i i oo oo 120cos120sin 120sin120cos 2 1 2 1 (1) where K is a current constant. Note that the v1- and v2-axes (v-plane) are rotated by 120 [deg] in the x- and y-axes as described in Fig. 4 (b). In the angular position during 15-30 [deg] and 60-75 [deg], the suspension windings su1 and su2 are excited in the uplane. The current commands isu1* and isu2* in the suspension windings su1 and su2 are determined from the suspension force commands Fx* and Fy* as, \u23a5 \u23a5 \u23a6 \u23a4 \u23a2 \u23a2 \u23a3 \u23a1 =\u23a5 \u23a6 \u23a4 \u23a2 \u23a3 \u23a1 =\u23a5 \u23a6 \u23a4 \u23a2 \u23a3 \u23a1 \u2217 \u2217 \u2217 \u2217 \u2217 \u2217 y x u u su su F F K F F K i i 2 1 2 1 (2) Note that the u1- and u2-axes (u-plane) are corresponding to the x- and y-axes. In the angular position during 30-45 [deg] and 75-90 [deg], the suspension windings sw1 and sw2 are excited in the wplane"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001868_elan.200800014-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001868_elan.200800014-Figure1-1.png",
        "caption": "Fig. 1. Schematic representation of a biofuel cell, at the anode biocatalysts oxidise the fuel and at the cathode, other catalysts reduce an oxidant, usually oxygen.",
        "texts": [
            " The values of the cell voltage can also decrease dramatically as a result of kinetic limitations, electron transfer rate at the electrode interfaces, ohmic resistances and concentration gradients [7]. The electrode size also has an important role in determining cell current, as well as the ion permeability and transport rate across the membrane separating the anodic and cathodic compartments. All these parameters make research on biofuel cells a fascinating multidisciplinary field and improvement of the limiting components is the key to the achievement of high current densities and long lasting biofuel cells. Figure 1 is a simple description of the components needed to make a biofuel cell. At the anode, chemical agents or biocatalysts oxidise the fuel and the electrons that are now present in the reduced form of the biocatalyst are shuttled to the electrode. Meanwhile, in the other half-cell, a further reaction takes place in which another biocatalyst reduces an oxidant, usually, dissolved oxygen. The electron transfer between the anode and the cathode produces the cell s power output. The attractions of biofuel cells are several: regardless of their kinetic disadvantages relative to conventional fuel cells, biocatalyst-driven devices can be used for more complex fuels like blood glucose or lactate"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003251_s12555-012-0070-9-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003251_s12555-012-0070-9-Figure1-1.png",
        "caption": "Fig. 1. Phase paths of the second-order system.",
        "texts": [
            "t\u03c3 Then the trajectory enters the halfplane ( ) 0t\u03c3 < (quadrant III). When ( ) 0,t\u03c3 < we have 1 2 ( ) ( ) ( )t l t l t k\u03c3 \u03c3 \u03c3+ + = and obtain its equivalent point as 2 ( , ) ( / ,0).k l\u03c3 \u03c3 = Since the system is stable, it follows that the system will hit the axis ( ) 0t\u03c3 = in finite time. Therefore, its solutions cross the axis ( ) 0t\u03c3 = from quadrant II to quadrant I, and from quadrant IV to quadrant III. After gluing these paths along the line ( ) 0,t\u03c3 = we obtain the phase portrait of the system, as shown in Fig. 1. Then we choose a Lyapunov function as 22 2 ( )( ) ( ) ( ) 2 2 l tt V t k t \u03c3\u03c3 \u03c3= + + , and then obtain its time derivative as ( )( ) ( ) 1 2 2 2 1 ( ) ( ) ( ) ( ) ( ) ( ) ( ) ( ) ( ) ( ). V t t l t l t ksign t l t t ksign t t l t \u03c3 \u03c3 \u03c3 \u03c3 \u03c3 \u03c3 \u03c3 \u03c3 \u03c3 = \u2212 \u2212 \u2212 + + = \u2212 From the above equation, we can obtain that the variables ( )t\u03c3 and ( )t\u03c3 asymptotically converge to zero. The proof of the lemma is finished. Theorem 1: Consider system (2) with satisfying (3). If the parameters l1 and l2, and the gain k are chosen to satisfy the following conditions: 2 1 2 4 l l < and ,k \u03b7> (5) then the two variables ( )t\u03c3 and ( )t\u03c3 asymptotically converge to zero"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003501_2041302510394742-Figure3-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003501_2041302510394742-Figure3-1.png",
        "caption": "Fig. 3 Transition to forward flight from hover and vice versa",
        "texts": [
            " From reference [23], rotating the free wings, the portions of propeller wash will not be blocked by the free wings and propeller wash only operates over the fixed-wing root portions and fuselage. The fixed-wing root portions in horizontal flight mode perform as the wings by generating lift in association with the free wings. Therefore, the normalized aerodynamic lift force and drag force can be obtained even during flying modes transition. Moreover, a backstepping controller is designed to control the hybrid aircraft to keep the flying height invariant during modes transition. The flying modes transition is shown in Fig. 3. 2 DESIGN OF ROTOR-FIXED WING HYBRID AIRCRAFT A designed rotor-fixed hybrid aircraft is shown in Figs 4(a) to (e). When the aircraft is in VTOL flight or in hover (see Fig. 4(b)), the thrust generated by the co-axial counterrotating propellers 1 provide lift force. The fixed-wing root portion 3 advantageously remains in the slip stream and dynamic pressure acting thereon tends to provide some degree of directional stability. From reference [23], rotating the free wings 2, the portions of propeller wash will not be blocked by the free wings 2 and propeller wash only operate over the fixed-wing root portion 3 and fuselage 4"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001052_physreve.75.011707-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001052_physreve.75.011707-Figure1-1.png",
        "caption": "FIG. 1. Geometry of the system. The coil illustrates the cholesteric helix emerging from the twist of the director orientation in the ground state.",
        "texts": [
            " 2 , there arise the following conditions of thermodynamic stability for the material parameters: c1 0, c1+c2 0, D1 0, D2 2 4c1D1, K1 0, K2 0, and K3 0. As can be seen from the definitions of the variables that enter Eq. 2 , in our linearized model there are five independent variables that completely describe the state of our system. They coincide with small deviations from the ground state. Those variables are the three components of the displacement field u r and the two independent components of n r , characterized by the condition n\u0302 \u00b7 n=0. The setup of the system we have in mind is sketched in Fig. 1, where the undistorted cholesteric helix given by n\u03020 from Eq. 1 is indicated, and its axis is parallel to the z\u0302 direction. If we denote the general form of n\u0302 r with the help of two angles, we obtain two independent variables that completely define n r : One is the angle that gives the tilting of the director out of the respective plane perpendicular to z\u0302, which we will call nz r , the other is the angle that determines the change of the phase of the director rotation around the helix axis, which will be referred to as r ",
            " Finding solutions to this set of equations then guides us to the actual state of the system. During the derivation of equations A1 \u2013 A5 , we thereby neglected energetic contributions of the sample surfaces and concentrated on the energetic contributions of the bulk. This is justified by the dimensions of the systems under consideration. We will include the effect of the surfaces by the boundary conditions we will impose. In the following, we will set the origin of our coordinate system as indicated in Fig. 1. The bottom of the sample is located at z=0 and the top at z=d. We will assume strong anchoring of the director at the bottom and the top of the sample, where in the next two sections the director is supposed to be fixed in the x\u0302 direction, n\u0302 z = 0 n\u0302 z = d 1 0 0 . 8 This assumption is certainly justified for the films synthesized by photo-cross-linking see 6 , because during the cross-linking process the polymer network gets covalently bound to the substrate, if supported films are produced. If free-standing films are synthesized, the polymer network gets covalently bound to a sacrificial layer, by which the substrate is coated and most of which is dissolved in water afterwards in order to separate the film from the substrate. In the ground state of the system, the adjusted cholesteric pitch 2L see Fig. 1 for homogeneous structures then fulfills the condition d=mL with m=1,2 ,3 , . . . . The wave number for the rotation of the cholesteric helix can therefore be written as q0 = m d with m = \u00b1 1, \u00b1 2, \u00b1 3, . . . . 9 What we want to point out at this point is that the equations A1 \u2013 A5 can be solved by an ansatz that separates the z dependence of the solution from the lateral dependencies, which are the x and y dependencies, ux r = cos kxx + kyy + u\u0303x z , uy r = cos kxx + kyy + u\u0303y z , uz r = sin kxx + kyy + u\u0303z z , nz r = cos kxx + kyy + n\u0303z z , r = sin kxx + kyy + \u0303 z ",
            " We emphasize that our results were obtained for the case of an isotropic elastic behavior of the polymer network. If, however, an anisotropic elastic response of the polymer network prevails, the resulting effects could quantitatively exceed the anisotropic behavior caused by the coupling to the director. As the absolute value of D2 can be quite small compared to the elastic coefficients, this must be taken into account when a corresponding experiment is performed. In this section, we want to investigate the geometry we already depicted in the beginning in Fig. 1. No external mechanical forces are imposed on the system, however a static external electric field is applied parallel to the cholesteric helix axis. We suppose the material to be a perfect electric insulator, and that a 0 in Eq. 2 , which means that the director tends to align parallel to the electric field. First we want to treat the problem as only z-dependent, which means that we are looking for solutions that are homogeneous over the whole sample in lateral directions, that is, in the x\u0302 and y\u0302 directions",
            " In this connection, it has to be noted that our calculations were performed assuming an isotropic elastic behavior of the polymer network, and that an anisotropic elastic response of solely the polymer network can mask this effect. Compression and dilation of the system in both directions perpendicular to the helix axis by equal force densities lead to a homogeneous stretching or compression of the cholesteric axis, which is consistent with the experiments already performed 7,8 . In the case of an external electric field applied parallel to the helix axis Fig. 1 and Sec. V , we detected a threshold value of the field amplitude at which the ground-state conformation of the system becomes unstable with respect to an instability. At onset we found two qualitatively different types of instabilities, one of which takes place homogeneously over the whole sample in the directions perpendicular to the cholesteric helix axis Sec. V A while the other one corresponds to undulations in these directions Sec. V B . As a reorientation of the director, the first of these two instabilities only contains a tilting of the director out of the planes perpendicular to the cholesteric helix axis"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001051_10402000903097361-Figure7-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001051_10402000903097361-Figure7-1.png",
        "caption": "Fig. 7\u2014Oil drop test; figure concept from Mu\u0308ller and Nru (20).",
        "texts": [
            " The temperature and therefore the viscosity changes at the interface are potentially higher as oil moves through and/or is pumped out of this region. Two types of tests are conducted\u2014one for seals with patterns expected to reverse pump, i.e., seal, and one for seals with patterns expected to forward pump. Forward pumping is defined as an enhanced leakage that pumps the lubricant out of the oil bath and through the interface to the environment side of the seal and can be used to exclude contaminants. If the seal reverse pumps, then a series of oil drop tests are performed as shown in Fig. 7. Five of the six surface types tested in this study are evaluated using oil drop tests where known amounts of room-temperature oil (250, 500, 1000, 1500 \u00b5L) are alternately injected on top of the elastomer/shaft interface using a digital pipette and the time required for the seal to pump this oil through the interface into the oil bath is recorded. The result is a reverse pumping rate. A spec- ified amount of time is allowed to pass between the recovery of the system from one oil drop and the injection of the next"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003154_icrera.2013.6749809-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003154_icrera.2013.6749809-Figure1-1.png",
        "caption": "Figure 1. 1/6th cross-sectional view of the 18 stator slot 12 rotor pole PM BLDC motor",
        "texts": [
            " In this paper magnet pole shaping technique has been used by varying the PM offset for cogging torque and torque ripple reduction [4]-[6]. The PM offset is varied from 0mm (base model) to 30mm and various performance parameters like cogging torque, back EMF, developed torque, torque ripple and efficiency are computed and analyzed using 2D FE analysis. From this it is observed that there is substantial reduction in torque ripple with extremely less reduction in efficiency. II. MAGNET POLE SHAPING PM offset variation has been used for magnet pole shaping [5]. Fig. 1 shows the 1/6th cross-sectional view of the 18 stator slot 12 rotor pole PM BLDC motor with only one magnet visible. Pole arc radius is used to create the magnet outer radius using the pole arc center, where pole arc radius is equal to magnet outer radius (without offset i.e. 0mm) minus PM offset. For example, if PM offset is 20mm and magnet outer radius (without offset i.e. 0mm) is 39mm, then pole arc radius would be 19mm. The PM offset is varied from 0mm to 30mm in the steps of 5mm therefore providing seven models i"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002855_pbce077e_ch14-Figure14.1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002855_pbce077e_ch14-Figure14.1-1.png",
        "caption": "Figure 14.1 (a) The Folaga III vehicle, consisting in a single cylinder with two wet ends. (b) The eFolaga concept, with separation at mid-vehicle to allow insertion of a universal payload",
        "texts": [
            " The interest here stems from the fact that the UAN module had to not only act as interface between the vehicle and a payload (an acoustic modem) but also host the vehicle mission supervision and management system. The interface between mission management and the eFolaga Guidance, Navigation and Control (GNC) system has been designed in order to maintain modularity also at the software level. The chapter is organized in three sections, one for each of the above topics, and a conclusion section. The eFolaga is an evolution of the Folaga III vehicle (Caffaz et al., 2010) purposely designed to be capable of carrying different kinds of sensors. The Folaga III (Figure 14.1(a)) was a torpedo-like vehicle, consisting in a fibre-glass waterproof cylinder connected to two wet ends hosting jet pumps for steering and propulsion in the surge direction. Inside the cylinder, a ballast chamber was present, controlling the vehicle buoyancy, while internal displacement of the battery pack was actuated through a wormscrew mechanism, allowing for pitch control. The Folaga III was essentially a long pipe, closed at both ends by waterproof bulkheads. All the fittings were going through the end bulkheads",
            " The use of NiMH batteries of total weight of approximately 13 kg was matched by the use of lightweight GRP (Glass Reinforced Polyester) hull such that the vehicle required minimal or no extra ballast to provide roll stability. Other vehicles with lightweight carbon hull of roughly the same diameter and lighter lithium\u2013ion batteries have to take external ballast strapped to the outside bottom on the hull. The basic idea at the start of the eFolaga project was to keep the Folaga III vehicle characteristics while allowing insertion of a \u2018universal\u2019 payload module at mid-vehicle (Figure 14.1(b)). Moreover, it was decided that the project should not develop payloads, but should define mechanical, electrical and electronics interfaces so that any third party could develop its own payload module, as appropriate, and mount it on the eFolaga. This approach has hence led to a series of constraints to which any payload module must conform in order to be mounted on eFolaga vehicles. Obviously, from a functional point of view, the requirement to the payload module is that it should be self-sufficient in its operation, booting up autonomously when energized from the vehicle power supply, and shutting down normally when disconnected, or under active control via communication with the eFolaga control computer"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003086_j.elecom.2012.05.002-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003086_j.elecom.2012.05.002-Figure1-1.png",
        "caption": "Fig. 1. Recurrent voltammetries (A) of a saturated solution of carbon dioxide; Scan",
        "texts": [
            " However, there is no certitude that all the carboxylates are functionalized when observing this procedure that remains qualitative. Carbon dioxide is known to poorly react electrochemically. Woks available in literature principally report electrolyses at lead cathodes [11,12] with formation of oxalate by coupling of CO2 anion radicals. The proton availability appears to be a factor that governs the product distribution (carbon monoxide, formate, glyoxylate, and glycolate were also reported to be formed). At a GC electrode (Fig. 1, curves A) in DMF containing TMABF4, it is found that a saturated solution of carbon dioxide exhibits a large step (Ep=\u22122.2 V) that progressively vanishes upon recurrent scans and disappears after ten cycles. The use of more bulky tetraalkylammonium cations gives very small currents in voltammetry. This fits with the fact that cathodic charge of HOPG is strongly favored by small size ammonium ions like TMA+ [10]. The behavior described abovemakes expecting that (i) the global electrochemical process is governed by the carbon cathodic charge (formation of a poly-nucleophilic species) and that (ii) these species can react with CO2. Carboxylation of electrogenerated carbanions issued from the electrochemical reduction of \u03c0-acceptors (naphthalene, anthracene) in organic aprotic media is now a well documented field [13,14]. Therefore, the chemical carboxylation of GC is expected to occur (Scheme 1, (I)) providing carbon covered by carboxylate functions, at least superficially. The surface coverage of such interfaces was checked by the reduction of a \u03c0-acceptor, like chloranil, and was found out to be quite large (Fig. 1, B). When the modified carbon is allowed to react with a bulky electrophile such a long chain RI (like 1-iodotetradecane) dissolved in DMF for a few minutes, the blocking of the surface may happen as depicted in Fig. 1, C. This does confirm the reactivity of CO2 toward GC under electron transfer. For this, several methods were used. First, the formation of an activated ester (example: reaction of 3-nitrophenol onto the carboxylated GC) that is expected to lead to a two-electron step (Ep=0.98 V, ester cleavage with expulsion of nitrobenzene reduced afterwards (Ep=\u22121.32 V). The current integration (ester reduction) leads to a global immobilization of 1.0\u00d710\u22128 mol cm\u22122. Secondly, the addition of a redox marked electrophile such as \u03c9-iodohexylferrocene (already used as a redox probe [15]) that successfully permits checking the immobilization of ferrocene (Ep=0"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000695_00423117508968470-Figure3-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000695_00423117508968470-Figure3-1.png",
        "caption": "Fig. 3. Tractor and trailer lateral motions and forces",
        "texts": [
            " , n) are functions of the vehicle motion and may be written as: Eliminating Zi,i+, (i= 1, . . . n-1) from equations ( I ) and (2) by using (3) gives: and for 1 <i<n where the vertical motion of the ith component is: Equations (4), (5) and (6) are the equations of vertical motion for the n-component vehicle train. D ow nl oa de d by [ U ni ve rs ity o f A ri zo na ] at 0 8: 45 0 7 A ug us t 2 01 3 The equations of lateral motion are obtained by a similar process, but related to a moving axis system located at the centre of gravity of the tractor (x fwd, y stbd, z down). Figure 3 gives the lateral motions and forces for the tractor unit and the i th trailer. Equating rates of change of linear and angular momentum to the applied forces and moments gives the following equations of motion for the lateral motions of the tractor unit and i th trailer: The lateral interaction equation for the ith trailer unit are: front y,+, = (mi((; + t (rl t rj) 9 - 7.) j =1 rear , = - (rnj($ + (rl t rj)U) - j= i+l The tyre lateral are functions of the motion of the appropriate component of the vehicle train and may be written: D ow nl oa de d by [ U ni ve rs ity o f A ri zo na ] at 0 8: 45 0 7 A ug us t 2 01 3 108 F"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002736_med.2011.5983178-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002736_med.2011.5983178-Figure1-1.png",
        "caption": "Fig. 1. Components of the Twin Rotor MIMO System",
        "texts": [],
        "surrounding_texts": [
            "A. LPV modelling of TRMS An accurate dynamic model of the system, that takes into account its high nonlinearity and cross coupling between its two axes, is required to achieve control objectives satisfactorily. In [10], a linear parameter-varying system (LPV) approach is suggested. A LPV system is defined as a linear time-varying plant whose state-space matrices are fixed functions of some vector of varying parameters \u03c8t . They can be described by time state-space equations of the form: x\u0307(t) = A(\u03c8t)x(t)+B(\u03c8t)u(t) y(t) = C (\u03c8t)x(t)+D(\u03c8t)u(t) (7) 978-1-4577-0123-8/11/$26.00 \u00a92011 IEEE 1083 where x(t) \u2208 R nx represents the state vector, u(t) \u2208 R nu denotes the control inputs and y(t) \u2208 R ny are the sensor outputs. Here A(\u03c8t) \u2208 R nx\u00d7nx , B(\u03c8t) \u2208 R nx\u00d7nu and C(\u03c8t) \u2208 R ny\u00d7nx are time-variant matrices scheduled by the vector \u03c8t of varying parameters. The order of the non-linear model (1)-(6) can be reduced neglecting the dynamics of the currents iah/v. Then, following the non-linear embedding in the parameters approach proposed in [12], the reduced model can be expressed in an LPV form by introducing the following set of time-varying parameters \u03c8t = [a11, a21, a22, a23, a24, a25, a51, a52, a54, a55, a56, b22]T as follows:\u23a1 \u23a2\u23a2\u23a2\u23a2\u23a2\u23a3 \u03c9\u0307h(t) \u03a9\u0307h(t) \u03b8\u0307h(t) \u03c9\u0307v(t) \u03a9\u0307v(t) \u03b8\u0307v(t) \u23a4 \u23a5\u23a5\u23a5\u23a5\u23a5\u23a6= Aabs(\u03c8t) \u23a1 \u23a2\u23a2\u23a2\u23a2\u23a2\u23a3 \u03c9h(t) \u03a9h(t) \u03b8h(t) \u03c9v(t) \u03a9v(t) \u03b8v(t) \u23a4 \u23a5\u23a5\u23a5\u23a5\u23a5\u23a6+ \u23a1 \u23a2\u23a2\u23a2\u23a2\u23a2\u23a3 b11 0 0 b22(pt) 0 0 0 b42 b51 0 0 0 \u23a4 \u23a5\u23a5\u23a5\u23a5\u23a5\u23a6 [ uh(t) uv(t) ] (8) where: Aabs(\u03c8t)= \u23a1 \u23a2\u23a2\u23a2\u23a2\u23a3 a11(pt) 0 0 0 0 0 a21(pt) a22(pt) a23(pt) a24(pt) a25(pt) 0 0 1 0 0 0 0 0 0 0 a44(pt) 0 0 a51(pt) a52(pt) 0 a54 a55(pt) a56(pt) 0 0 0 0 1 0 \u23a4 \u23a5\u23a5\u23a5\u23a5\u23a6 with: a11(pt) = \u2212 k2 ah\u03d52 h /Rah+Btr+ f1(\u03c9h)/\u03c9h Jtr , a21(pt) = lt cos\u03b8v f2(\u03c9h)/\u03c9h Dcos2 \u03b8v+E sin2 \u03b8v+F a22(pt) = \u2212 koh Dcos2 \u03b8v+E sin2 \u03b8v+F , a23(pt) = \u2212 f3(\u03b8h)/\u03b8h Dcos2 \u03b8v+E sin2 \u03b8v+F a24(pt) = \u2212 km(Bmr+k2 av\u03d52 v /Rav+ f4(\u03c9v)/\u03c9v)cos\u03b8v (Dcos2 \u03b8v+E sin2 \u03b8v+F)Jmr a25(pt) = km\u03c9v sin\u03b8v(Dcos2 \u03b8v\u2212E sin2 \u03b8v\u2212F\u22122E cos2 \u03b8v) (Dcos2 \u03b8v+E sin2 \u03b8v+F)2 a44(pt) = \u2212 k2 av\u03d52 v Rav +Bmr+ f4(\u03c9v) \u03c9v Jmr , a51(pt) = \u2212 kt(Btr+ f1(\u03c9h)/\u03c9h+k2 ah/Rah) JvJtr a52(pt) = kg cos(\u03b8v) f5(\u03c9v)/\u03c9v\u22120.5\u03a9hH sin2\u03b8v Jv , a54(pt) = lm f5(\u03c9v)/\u03c9v Jv a55 = \u2212 kov Jv , a56(pt) = g((A\u2212B)cos(\u03b8v)\u2212C sin\u03b8v) Jv\u03b8v b11 = kah\u03d5hk1 JtrRah , b22(pt) = km cos(\u03b8v)kav\u03d5vk2 (Dcos2 \u03b8v+E sin2 \u03b8v+F)RavJmr b42 = kav\u03d5vk2 JmrRav , b51 = kah\u03d5hk1kt JvJtrRah The scheduling variables used in the model (8) are pt = [ \u03b8h \u03b8v \u03c9h \u03c9v \u03a9h ]T (9) B. TRMS Shamma Representation A LPV model for the TRMS can be obtained by means of state transformation from (8) by considering: x1 = [ \u03b8h \u03b8v ]T ; x2 = [ \u03c9h \u03a9h \u03c9v \u03a9v ]T By defining new states and inputs, a new LPV model of TRMS can be obtained. Then, this model can be reshaped by changing the order of the state variables, obtaining:\u23a1 \u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a3 \u0307\u03c9h(t) \u03a9\u0307h(t) \u03b8\u0307h(t) \u0307\u03c9v(t) \u03a9\u0307(t) \u03b8\u0307v(t) \u23a4 \u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a6 = Aabs(\u03c8t) \u23a1 \u23a2\u23a2\u23a2\u23a2\u23a2\u23a3 \u03c9h(t) \u03a9h(t) \u03b8h(t) \u03c9v(t) \u03a9v(t) \u03b8v(t) \u23a4 \u23a5\u23a5\u23a5\u23a5\u23a5\u23a6+ \u23a1 \u23a2\u23a2\u23a2\u23a2\u23a2\u23a3 b11 0 0 b22(pt) 0 0 0 b42 b51 0 0 0 \u23a4 \u23a5\u23a5\u23a5\u23a5\u23a5\u23a6 [ uh(t) uv(t) ] (10) Aabs(\u03c8t) = \u23a1 \u23a2\u23a2\u23a2\u23a2\u23a2\u23a3 a11 \u2212 d\u03c9eq h d\u03b8h 0 0 \u2212 d\u03c9eq h d\u03b8v 0 a21 a22 0 a24 a25 0 0 1 0 0 0 0 0 0 0 a44 \u2212 d\u03c9eq v d\u03b8v 0 a51 a52 0 a54 a55 0 0 0 0 0 1 0 \u23a4 \u23a5\u23a5\u23a5\u23a5\u23a5\u23a6 where \u03c9h/v(t) = \u03c9h/v(t)\u2212\u03c9eq h/v (t), uh/v(t) = uh/v(t)\u2212u eq h/v (t), \u03c9eq h = f3(\u03b8h) lt cos(\u03b8v) f2(\u03c9h) , \u03c9eq v = g\u03c9v [C sin(\u03b8v)+(B\u2212A)cos(\u03b8v)] lm f5(\u03c9v) and u eq h/v = kah/v\u03d5h/v +Rah/v(Btr/mr + f1/4(\u03c9h/v)/\u03c9h/v)/(kah/v\u03d5h/v) k1/2 \u03c9eq h/v d\u03c9eq h d\u03b8h = \u23a7\u23aa\u23aa\u23a8 \u23aa\u23aa\u23a9 kchp\u03c9h lt cos\u03b8v f2(\u03c9h) i f \u03b8h 0 kchn\u03c9h lt cos\u03b8v f2(\u03c9h) i f \u03b8h < 0 (11) d\u03c9eq h d\u03b8v = f3(\u03b8h)\u03c9h sin\u03b8v lt f2(\u03c9h)cos2 \u03b8v (12) d\u03c9eq v d\u03b8v = g\u03c9v [C cos\u03b8v +(A\u2212B)sin\u03b8v] lm f5(\u03c9v) (13) For a sake of simplicity, the dependance on the pt vector of the matrix Aabs elements has been omitted. A discrete-time LPV model can be obtained using Euler approximation with a sample time Ts to allow a digital implementation of the observer/controller:\u23a1 \u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a3 \u03c9h(k +1) \u03a9h(k +1) \u03b8h(k +1) \u03c9v(k +1) \u03a9v(k +1) \u03b8v(k +1) \u23a4 \u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a6 = Aabsd(\u03c8k) \u23a1 \u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a3 \u03c9h(k) \u03a9h(k) \u03b8h(k) \u03c9v(k) \u03a9v(k) \u03b8v(k) \u23a4 \u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a6 + Babsd(\u03c8k) [ uh(k) uv(k) ] (14) where Aabsd(\u03c8k) = \u23a1 \u23a2\u23a2\u23a2\u23a2\u23a2\u23a3 1+a11Ts \u2212 d\u03c9eq h d\u03b8h Ts 0 0 \u2212 d\u03c9eq h d\u03b8v Ts 0 a21Ts 1+a22Ts 0 a24Ts a25Ts 0 0 Ts 1 0 0 0 0 0 0 1+a44Ts \u2212 d\u03c9eq v d\u03b8v Ts 0 a51Ts a52Ts 0 a54Ts 1+a55Ts 0 0 0 0 0 Ts 1 \u23a4 \u23a5\u23a5\u23a5\u23a5\u23a5\u23a6 Babsd(\u03c8k) = \u23a1 \u23a2\u23a2\u23a2\u23a2\u23a3 b11Ts 0 0 b22Ts 0 0 0 b42Ts b51Ts 0 0 0 \u23a4 \u23a5\u23a5\u23a5\u23a5\u23a6 C. TRMS Polytopic LPV Model A common approach to design a controller/observer for the LPV system (7) is to approximate it by a polytopic LPV system. This kind of LPV systems corresponds to those whose time-varying parameter vector varies within a polytope. More precisely, the state-space matrices range in a polytope of matrices defined as the convex hull of a finite number of matrices N. Each polytope vertex corresponds to a particular value of scheduling variable. In other words:( A(\u03c8k) B(\u03c8k) C(\u03c8k) D(\u03c8k) ) \u2208Co {( A j B j Cj D j ) , j = 1...N } 978-1-4577-0123-8/11/$26.00 \u00a92011 IEEE 1084 : = \u2211N j=1 \u03b1 j k (\u03c8k) ( A j B j Cj D j ) (15) with \u03b1 j k (\u03c8k) \u2265 0 and \u2211N j=1 \u03b1 j k (\u03c8k) = 1 , where each jth model is called a vertex system. Because of this property, this type of LPV systems is referred as polytopic. The simplest polytopic approximation relies on bounding each LPV parameter by an interval. This approximation is known as bounding box approach and it leads to an interval LPV model. An alternative approach known as small hull approach proposed by [13] is based on the fact that not only the range of varying parameters is known, but also intelligent choices can be made about the intermediate values that the varying parameter can take. From this set of parameter values, the convex hull that contain them is obtained using the quick hull algorithm [14]. Proceeding with one of these two approaches, the polytopic LPV system (15) can be expressed as follows: x(k +1) = N \u2211 j=1 \u03b1 j k (\u03c8k) [A jx(k)+B ju(k)] y(k) = N \u2211 j=1 \u03b1 j k (\u03c8k)Cjx(k) (16) where x(k) \u2208 R nx represents the state vector, u(k) \u2208 R nu denotes the control inputs and y(k) \u2208 R ny are the sensor outputs. Here A j \u2208 Rnx\u00d7nx , B j \u2208 R nx\u00d7nu and Cj \u2208 R ny\u00d7nx are time-invariant matrices defined for jth model. The polytopic system is scheduled through functions designed as follows: \u03b1 j k (\u03c8k) , \u2200 j \u2208 [1, ...,N] that lie in a convex set: \u03a9 = { \u03b1 j k (\u03c8k) \u2208 R N ,\u03b1k(\u03c8k) = [ \u03b11 k (\u03c8k), . . . ,\u03b1N k (\u03c8k) ]T , \u03b1 j k (\u03c8k) \u2265 0, \u2200 j, N \u2211 j=1 \u03b1 j k (\u03c8k) = 1 } (17)"
        ]
    },
    {
        "image_filename": "designv11_12_0002273_1.3680609-Figure3-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002273_1.3680609-Figure3-1.png",
        "caption": "Fig. 3. Lacrosse type launcher used to throw balls.",
        "texts": [
            " Ball 1 was an unmodified polystyrene ball. The hollow plastic ball was smooth, apart from a small indentation used to inflate the ball. It was manufactured as a child\u2019s basketball and was slightly larger in diameter than an approved soccer ball (218\u2013221 mm). The balls listed in Table I were launched either by hand at relatively low speed and low spin or at higher speed and spin with a homemade lacrosse type ball launcher. The launcher was constructed from a 1.5-m length of 5-mm diameter aluminum rod, bent into the shape shown in Fig. 3, and bolted to a rectangular wood handle. When launching ball 2, the string seam did not come into contact with the launcher so sidespin could be avoided. The lacrosse launcher was swung either by hand or by pivoting it in a frame using an elastic bungee cord to swing it more precisely at controlled and adjustable speeds. No attempt was made to control the ball speed and spin separately, with the result that the spin imparted to the ball was approximately proportional to the launch speed, both when throwing by hand and when using the lacrosse launcher"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001345_robot.2008.4543604-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001345_robot.2008.4543604-Figure2-1.png",
        "caption": "Figure 2. The force sensors distribution",
        "texts": [
            " The paper is organized as follows: section 2 presents the basic principles of a hyperredundant structure with continuum elements; section 3 studies the dynamic model; section 4 discusses the both problem of grasping by coiling, the position control and force control; section 5 verifies by computer simulation the control laws. II. BACKGROUND A. Technological Model The paper studies a class of hyperredundant arms that can achieve any position and orientation in 3D space, and that can perform a coil function for the grasping (Figure 2). The arm is a high degree of freedom structure or a continuum structure. The general form of the arm is shown in Figure 3. It consists of a number (N) of elements, cylinders made of fiber-reinforced rubber. There are four internal chambers in the cylinder, each of them containing the ER fluid with an individual control circuit. The deformation in each cylinder is controlled by an independent electrohydraulic pressure control system combined with the distributed control of the ER fluid. The cylinder can be bent in any direction by appropriately controlling the pressure in the four chambers"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003761_9781118818428-Figure4.1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003761_9781118818428-Figure4.1-1.png",
        "caption": "Figure 4.1 Schematic representation of the three principal configurations. After Finkelstein and Organ (2001), with permission of the American Society of Mechanical Engineers (ASME)",
        "texts": [
            " Once these have been isolated and related to their respective causes, attention can turn to whittling them down. 4 Equivalence conditions for volume variations The cycle of interacting gas processes is driven by periodic variations of the \u2018live\u2019 volumes. The task of setting up a numerical model of the interactions therefore starts with a statement of volume variations as a function of crank angle \u03c6. Many crank mechanisms generate motions close to simple-harmonic. For the opposed piston, or \u2018alpha\u2019, configuration this allows simple expressions forVe(\u03c6) andVc(\u03c6). Employing the notation of Figure 4.1: Ve(\u03c6) = 1\u20442VE(1 + cos \u03c6) (4.1a) Vc(\u03c6) = 1\u20442\u03baVE(1 + cos[\u03c6 \u2212 \u03b1]) (4.1b) Angles \u03b1 and \u03b2 are the kinematic (i.e., crank) phase angles for the respective configuration \u2013 the angle by which the crank pin of the expansion piston (or displacer) leads that of the compression piston. In the case of the opposed-piston machine (only) \u03b1 is also the volume phase angle, or thermodynamic phase angle \u2013 the angle by which the Ve(\u03c6) leads Vc(\u03c6). While Equation 4.1a for expansion volume Ve(\u03c6) extends to \u2018beta\u2019 and \u2018gamma\u2019 types, there is no analogous expression for Vc(\u03c6) because \u03b2 is not the volume phase angle",
            " Its value is now seen to lie in allowing the simple expression for the volume variations of the gamma type to serve both machines and handling the beta option by merely subtracting the numerical value of Vdc + from the compression-end dead space of the equivalent gamma. A gas-process model does not \u2018know\u2019 whether events in the gas path are occurring in an opposed-piston configuration (alpha) or in the displacer (beta or gamma) arrangement. If expansion and compression space volumes of different engines were to follow identical respective cycle histories, then a simulation of one could amount to a simulation of all \u2013 regardless of mechanical layout. The lowermost diagram of Figure 4.1 is the equivalence condition reduced to a \u2018physical\u2019 counterpart. In the two-piston, alpha machine kinematic phase angle is equal to volume phase angle \u2013 the angular difference at the crank between maximum (or minimum) expansion space volume and maximum (or minimum) volume of the compression space. 5 The optimum versus optimization The account byKarabulut et al. (2009) comes as something of awake-up call: A single-cylinder, beta-configuration engine of 296 cm3 swept volume operating between source temperature TE of 200 \u25e6C and sink TC of 27 \u25e6C (temperature ratio NT of 1"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003560_cdc.2012.6425834-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003560_cdc.2012.6425834-Figure2-1.png",
        "caption": "Fig. 2. Coordinate systems: Inertial frame (F i) and Body-fixed frame (F b)",
        "texts": [
            " In order to deal with situations when the image is lost, a switching strategy is proposed, which allows switching between two different controllers (PD and PI). In hover mode, the Quad-plane behaves like a conventional Quad-rotor, then we use the well known mathematical model of the Quad-rotor [2]. For the purpose of control we will use the linear model (1) instead of the nonlinear version, since in general the vehicle operates in areas where |\u03b8| \u2264 \u03c0/2 and |\u03c6| \u2264 \u03c0/2, these constraints are satisfied even when the nonlinear model is used together with a feedback control [2], [7]. The corresponding coordinate system is represented in (Fig.2). x\u03071 = x2 \u03b8\u03071 = \u03b82 x\u03072 = \u2212\u03b8u1 \u03b8\u03072 = u2 y\u03071 = y2 \u03c6\u03071 = \u03c62 y\u03072 = \u03c6u1 \u03c6\u03072 = u3 z\u03071 = z2 \u03c8\u03071 = \u03c82 z\u03072 = 1\u2212 u1 \u03c8\u03072 = u4 (1) IV. VISION-BASED STATES ESTIMATION A sufficiently smooth road is represented in the camera\u2019s image as a group of lines, (Fig. 3). A straight line in the image can be seen as a segment of infinite length and whose center of gravity belongs to the straight line [8]. Using the probabilistic Hough transform method available in OpenCV [9], a straight line will be represented as \u03c1 = x cos \u03b8+y sin \u03b8"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003713_978-1-4614-2419-2-Figure50.11-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003713_978-1-4614-2419-2-Figure50.11-1.png",
        "caption": "Fig. 50.11 Calibration of bending moments",
        "texts": [
            "9 shows the experimental setup used for calibrating the axial force measurement provided by the dynamometric axle. An axial force is imposed moving a crossbeam connected to the axle. The imposed force is measured through a load cell connected one side with the crossbeam the other with the axle. The obtained calibration curve for the rear axle is shown in Fig. 50.10. A linear relation can be clearly noticed (index R2 > 0.99). A similar result has been achieved for the front axle. The test for calibrating the bending moment measurement about the x axis is shown in Fig. 50.11. During the test, skates have been attached to the dynamometric axle in order to reproduce real working conditions. The bending moment is imposed applying calibrated weights (known force F) at the center of the axle. A maximum force F of 2,000 N has been imposed. Although not completely representative of real working conditions, the measurement of the bending moment about the z axis has been calibrated using the same experimental setup previously described, but rotating the axle of 90 . The calibration curve for the bending moments measured in section BB about the x and the z axes are reporter in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000375_02286203.2008.11442485-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000375_02286203.2008.11442485-Figure2-1.png",
        "caption": "Figure 2. View of the harmonic drive test apparatus.",
        "texts": [
            " The harmonic drive gear consists of the mechanical assembly of three components: a rigid circular spline, an elliptical wave generator, and a nonrigid flexible spline or flexspline, which form together a compact, high-torque, high-ratio, in-line gear mechanism as shown in Fig. 1. A harmonic drive test apparatus was designed and built at Rice University as a platform to perform various types of experiments on the harmonic drive and to characterize the different errors inherent in its operation while preventing any external error component from being imposed [8, 16]. The system is shown in Fig. 2. It has its major axis of motion in the vertical plane to avoid the radial loading problem. A special design of vertical support plates and circular steel pipe sections with a highly stable platform was also used to maintain torsional integrity of the system. Effort was also put into making the linkage joining the motor, harmonic drive, and torque sensor very rigid. This aspect is important because the objective was to avoid any torsion in the system produced by elements other than the flexspline and the harmonic drive as a hole"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000532_te.2007.907321-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000532_te.2007.907321-Figure1-1.png",
        "caption": "Fig. 1. Illustrations for applications of the optoelectronic sensing techniques. (a) Direct coupling for optical power measurement. (b) T-type configuration for CD/DVD players. (c) Optical coupling using fibers. (d) Spatial radiance distribution measurement for light sources.",
        "texts": [
            " Diodes are generally discussed in a microelectronics course and the extended idea is not too difficult for students to grasp. Moreover, one may apply the same idea to semiconductor light sources, e.g., light-emitting diodes (LEDs). Spatial radiance distribution measurement, a kind of optoelectronic sensing technique, conveys important information to improve the utilization efficiency (or the characteristic of directionality) of the light sources. Thus, this technique gives rise to much attention in industries [2]. In this paper, the measurement system for the project [Fig. 1(d)] is proposed as a key basis to motivate students learning a variety of optoelectronic sensing techniques. For example, in Fig. 1(a)\u2013(c), this technique can be applied in optical power measurement, source\u2013receiver configuration for CD/DVD players, and optical coupling using fibers, respectively. The objectives of this course (or educational objectives) are given as follows. 1) In the proposed learning activity, the optoelectronic sensing project is to lead students progressively to work with the case studies from basic to advanced ones and further to achieve industrial application goals. 2) Through this project, student teams are to be guided to identify, formulate, and solve the optoelectronic sensing problems",
            " In the experiment, to demonstrate the effectiveness of the noise suppression, the voltage variation (or ripple) generated by use of a power transformer, a full-wave rectifier, and a rectifier filter may be employed to emulate a noise source in realistic situations in the circuit, e.g., temperature variation. In this task, students are required to enhance their ability to identify, formulate, and solve the problems from the stabilizer circuit design. According to the system requirement, the idea behind Task 2 is to guide students to design PD amplifiers progressively from basic to advanced ones, and further to achieve the goals of optoelectronic sensing in the practical applications (Fig. 1). Task 2 contains four subtasks corresponding to Proposals 2-1 to 2-4, as shown in Table II. All of the subtasks are designed to motivate students to learn the related technical materials actively. Through the use of TPs 2-1 to 2-5, students are progressively guided to solve a series of inevitable problems as simply as possible, to achieve the task aims, as indicated in Table II. Subtask 2-1 motivates them in how to design an electronic circuit to convert a tiny photocurrent (about to ampere per lux) to a measurable voltage"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001157_j.optlastec.2008.10.006-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001157_j.optlastec.2008.10.006-Figure1-1.png",
        "caption": "Fig. 1. Schematic of the powder delivery process in nozzle.",
        "texts": [
            " The following assumptions are adopted to simplify the problem: (1) The powder particles are spherical with the same particle size. (2) Only the drag force, gravity, and friction force are considered in present study. (3) The powder particles travel with the same velocity vp in a uniform gas flow in nozzle. (4) The particle collision is not considered. Powder particles group is taken as the analyzed object in this part in order to deal with the friction force conveniently. Provided that the gas-powder mixture flows in a straight nozzle with inside diameter d and incline angle y to the horizontal, as illustrated in Fig. 1. The motion of the powder particles group is governed by an equation that balances the mass-acceleration of the powder particles group with the forces acting on it. Take account of the powder volume with a thickness of Dl (labeled by the dashed lines in Fig. 1), and the governing equation is FD \u00femg sin y Ftp \u00bc m dvp dt (1) where m is the mass of the object. The first term on the left-hand side of Eq. (1) is the drag force FD \u00bc Nf D \u00bc m 6 pd3 prp f D (2) where N is the number of the particles in the analyzed volume, and fD is the drag force acting on each particle which can be described by f D \u00bc rgCD\u00f0vg vp\u00de 2 2 pd2 p 4 (3) There are many empirical models to determine the drag coefficient. For Rep800, CD can be expressed as [15]: CD \u00bc 24 Re \u00f01\u00fe 0:15R0:687 e \u00de (4) where Re is the Reynolds number defined by: Re \u00bc rgdpjvg vpj mg (5) The carrier gas velocity ng can be calculated by: vg \u00bc Qc p\u00f0d=2\u00de2 (6) The third term on the left-hand side of Eq"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003524_j.rcim.2012.02.004-Figure3-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003524_j.rcim.2012.02.004-Figure3-1.png",
        "caption": "Fig. 3",
        "texts": [
            " 2, for a given point is (b cosj, a sinj) located on the surface of oval bar, the projected coordinates in the horizontal plane can be calculated from y\u00bc \u00f0asinj z\u00decot\u00f0b=2\u00de z\u00bc bcosj ( \u00f011\u00de where, b is the half minor axis, a is the half major axis, and j is the eccentric angle. Value of z can be calculated from Eq. (2) and value of b can be obtained from b\u00bc cos 1\u00bd\u00f0R0\u00fe0:5hk asinj\u00de=\u00f0R0\u00fe0:5hk asinj\u00de \u00f012\u00de where, R0 is the minimum radius of rolls, hk is the height of groove. The projected area of horizontal plane Ax can be calculated using numerical integration method. Fig. 3 shows the outline of round groove drawn superimposed on the cross section of an oval work piece. It is easy to get the intersections of the oval curve and the round groove profile curve, and then the project areas of contact in the vertical plan Ay can be calculated through integral operation for the blank area in the center of Fig. 3. After that the value of contact area can be obtained through solving Eq. (10). Contact area calculation for other roll passes can use the same approach. For a given point C located in the deformation zone, Fig. 2 shows the start point of deformation C1, Fig. 3 gives the end point of deformation C2, thus, the length of contact arc related to point C can be presented by x\u00bc asinj y\u00bc b\u00f0R0\u00fehk=2 z\u00de ( \u00f013\u00de For any point located on the surface of deformation zone, the lengths of contact arc and sliding distance have a leading role in its rate of wear. For example, assuming the load is distributed uniformly along the deformation zone, in the flat rolling process, due to the contact arc length as well as the sliding distance of every deformation point equals those of other deformation points, rolls can achieve a uniform wear, and the useful life of roll can be extended"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003046_2013.39761-Figure3-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003046_2013.39761-Figure3-1.png",
        "caption": "FIG. 3 Screw feeder detail.",
        "texts": [
            " According to Brubaker and Pos (1) the coefficient of friction of various grains on cold-rolled steel and galva nized sheet are given in Table 4, in which the range is due to moisture con tent. If the figures (Table 4) hold for the grain on the material of the screw, it is quite evident that few grains would 686 Table 1. Design Factors Affecting a Simple Screw: 1 Outside diameter 2 Inside or shaft diameter 3 Pitch 4 Number of flights 5 Flight thickness 6 Length of flight 7 Material of manufacture give trouble with a full-pitch screw and few would not give trouble on a halfpitch screw. Fig. 3 is a line drawing of the screw at the entrance to the tube or housing. If a notch, large enough to pass the largest particle to be conveyed, is cut in the screw flight just inside the hopper, as shown, jamming can be elim inated. Such a notch might be trou blesome if long stringy material is en countered. In the equipment used, since suitable screws could not readily be obtained commercially, they were turned from a 2-in. solid aluminum bar of one of the harder, free-machining alloys. To date screws of 1, 1%, IV2, 1%, 2 and 21/4-in"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000798_14689360701191931-Figure6-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000798_14689360701191931-Figure6-1.png",
        "caption": "Figure 6. A typical scattering orbit for the 2-body problem with inertia operator A1. The relative momentum 2P \u00bc p1 p2 remains bounded during the close approach.",
        "texts": [
            " A typical such encounter is shown in figure 5. The H3 metric is smoother, but G3(r) is still not four times differentiable at r \u00bc 0, so the above separatrix analysis is still not valid. However, one can check that the 2-body problem does now have a separatrix that divides capture and scattering orbits, although it is not a smooth curve at r \u00bc 0. For Hk (k 4) and H1, the separatrix is as described earlier. Figure 2 shows the phase portrait for the H1 metric. A typical scattering orbit is shown in figure 6; in contrast to the H2 case, the relative momentum 2P \u00bc p1 p2 is bounded over all scattering orbits. In no case are there any periodic orbits in the two body problem. 4.3. An alternative regularization for H1 As was mentioned earlier, there is another possible regularization of the N-particle problem with the H1 metric, which we consider briefly here. This is simply to set the self-induced velocity of each particle to zero, i.e., G\u00f00\u00de \u00bc 0. This seems drastic, when the \u2018correct\u2019 self-induced velocity of a delta-function is infinite, but it still corresponds to a consistent discretization of the PDE (4) in the limit of a large number of particles spaced over a curve or an area"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003362_j.proeng.2013.08.190-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003362_j.proeng.2013.08.190-Figure1-1.png",
        "caption": "Fig. 1. Laser Tracker measuring principle.",
        "texts": [],
        "surrounding_texts": [
            "The importance of accuracy in Laser Tracker Systems (LT) is its function as a standard measurement for a wide range of equipment and facilities. The scope of the present work is to establish a verification procedure and simplified calibration method to correct the errors caused in the LT on a regular measuring range. The procedure will be based both on measurements of patterns with known and calibrated distances and the measurement of a mesh of reflectors placed at unknown locations looking to reduce the time and cost of testing and calibrating of the * Corresponding author. E-mail address: jconte@unizar.es t rs. Published by Elsevier Ltd. electi a eer-review under responsibility of Universi a de Zaragoza, Dpto Ing Dise\u00f1o y Fabricacion equipment. In addition, the mathematical model of the LT will be determined obtaining its kinematic parameters; by one hand considering its model error and by the other throughout uncertainty assessment techniques based on Monte Carlo method and considering the influence of the error sources. This simulation will establish a priori the better conditions for capture points, ie leading to a lower measurement uncertainty in points captured. Because of the large number of sources of error to be considered in this type of equipment, an approximation of these features will, before capture, offer an optimum manner the position of LT in the capture procedure and the sequence and point to check. 2. Laser Tracker measuring principle The LT is a measuring instrument that tracks the movement of a reflector and calculates its position in spherical coordinates. The distance to the reflector (d) can be measured by an interferometer (IF) or by an absolute distance meter (ADM), while the inclination angles ( ) and azimuthal ( ) are measured by two angular encoders. The reflector returns the laser beam, where the beam strikes a position sensor (PSD) that detects any change in position causing the movement of the axes of LT so that the laser beam is incident on the optical center of the reflector. Thus the LT head constantly monitors the position of the reflector. There are also reflectors mounted on rotating devices with two degrees of freedom which can also follow the LT beam, which allow the simultaneous movement of the emitter and reflector so that both seek its correct alignment. This expands the measurement possibilities with moving equipment (machine tools, robots) without need for multiple reflectors or interrupt the measurement process to move or redirect the spotlight. Having the spherical coordinates, it\u2019s possible to obtain the cartesian coordinates of the reflector with respect to the reference source of LT with equations 1 to 3: (1) (2) (3) These are the nominal values corresponding to a perfect LT, equipment calibration will give the correction parameters needed to correct these values by modeling errors."
        ]
    },
    {
        "image_filename": "designv11_12_0001247_0954406jmes321-Figure5-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001247_0954406jmes321-Figure5-1.png",
        "caption": "Fig. 5 Coordinate system S2 projected to coordinate system Sg",
        "texts": [
            " Since this paper aims to design a face gear made by a moulding process \u2013 in which the mould surface is directly or indirectly machined by a computer numerical control (CNC) machine tool \u2013 there is great freedom in designing the tooth surface of the face gear. Thus, this paper proposes a methodology that modifies the tooth surface by superimposing a double crowning on the standard face gear both in the contact path and the instant contact line directions, as shown in the following steps. Step 1: determining the primary contact point P First, a primary contact point P is chosen on the projected plane Sg(Xg, Yg) of the face gear (Fig. 5), usually the highest point on the motion curve and located at the centre of the tooth flank. The coordinate system Sg can then be represented as follows xg = (\u221a x2 2 + y2 2 \u2212 Px ) (16) yg = z2 2 \u2212 Py (17) Step 2: determining the orientation angle \u00b5 of the contact path The instant contact line L1 between the standard face gear and the involute pinion at point P (Fig. 6) inclines with respect to axis Xg at incline angle \u03b6 , which can be determined by solving equations (11) and (15) with the predetermined parameter \u03c61"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000358_icca.2007.4376367-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000358_icca.2007.4376367-Figure1-1.png",
        "caption": "Figure 1. Engagement geometry in three-dimensional space",
        "texts": [
            " Finally, three-dimensional air interception simulation was done. II. SPACE ENGAGEMENT MODEL[91 We first describe mathematically the 3-dimensional pursuit situation. To simplify the dynamic equations of the pursuit situation, we assume that the missile and target are point masses and that the autopilot and seeker dynamics of the missile are fast enough to be neglected. We further assume that the velocities of the missile and target are constant. The engagement geometry situation in three-dimensional space is depicted in Figure. 1, which describes the missile and target relative states. 299 1-4244-0818-0/07/$20.00 (\u00a2 2007 IEEE In Figure. 1, ( OXiyiZi ) Reference (ground) coordinate system; ( OXLYLZL ) Line of Sight (LOS) coordinate system;( OX.Y.Z. ) Missile body coordinate system;( OX%Y,Z,) Target body coordinate system. The engagement geometry situation in Figurel can be represented by the following dynamic vector equations from (1) to (6). dL d(iL) Vi Vmm = QL dt dt At = aytjt + a,tk t = QL X Vt +Q X Vt Am = aj+azmkm = QL X Vm +QmXVm Qt, = V sin 6it + Vf cosoji + 0,k, fm = tkm sin 6mim + eVm cos 6mi + 6mkm 1L qxiL + qyiL + qzkL = V'L sinI9L1L + fL COS9LJL +dLkL (1) (2) (3) (4) (5) (6) Where VM V, velocity of missile and target respectively; 12L LOS rate; \u00a3m Missile to LOS angle rate; \u00a3t Target to LOS angle rate; A, (A,) Missile(Target) acceleration vector; L LOS vector; r Missile-to-target range; ',,q, q 12L vector along iL,jL,kL - Then, before proceeding with the derivation of the new guidance laws, the engagement geometry equations ((7)- (13)), which describes the missile-target, relative kinematics is presented as follows"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003302_s10514-011-9236-1-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003302_s10514-011-9236-1-Figure1-1.png",
        "caption": "Fig. 1 Abstract models of the active and passive modules. The active model has 3 joints and 2 end grippers for 3D motion in a truss environment as well as grasping a passive module which is a simple bar (Yoon 2006)",
        "texts": [
            " We consider a target truss structure composed of modular manipulators and passive bars. Instead of restricting our application to a specific type of structure, we use abstract models for active and passive modules so that the modules can be scaled and specialized according to the target structure. For implementation, this allows us to re-use designs of hardware (3D model and electronic system) and software (kinematic controller and user interface). We propose abstract models for the active and passive modules as shown in Fig. 1. 3.1 Active module Active modules have minimal design for 3D navigation on truss and for grasping passive modules. Three actuated joints enables a 3D configuration of the module, and two grippers are contained for attaching passive modules and locomotion along a fixed truss by successively gripping and swinging (Detweiler et al. 2007). 3.2 Passive module The passive modules are rigid struts. In this paper, we use a straight bar with a fixed length, however, geometry may be application dependent and multiple types of passive parts can be considered according to application. The role of the passive modules is as a medium for connecting multiple active modules as shown in Fig. 1. This gives strength and large workspace to an assembled structure. In addition, it prevents the reduction of the number of DOF when two active modules are connected. By connecting two active modules directly we can generate a 5DOF linkage.1 A 6DOF linkage is obtained by using a truss element as a medium of connection as in Fig. 1.2 3.3 Self-assembled linkage: walking tower Multiple active modules can connect to one another using passive modules to form a larger active structure. The robots become smart joints in the self-assembled structure: they can actuate the structure to travel, bend, twist, and selfreconfigure. Figure 2 shows snapshots of the self-assembly of a truss tower. Twelve active modules and eight passive bars are employed to build a three-dimensional tower that can reconfigure itself by controlling active parts",
            " The maximum deflection when two robots that are connected in a fully stretched configuration is approximately 20 mm. The maximum number of robot modules that can be supported by a robot can be increased by using more powerful motors. The robot can only reach only a specific discrete set of points, that correspond to nodes in the truss graph model (Yun and Rus 2007), and every robot has identical structure and capabilities. Self-assembly can be implemented by Shady3Ds, using a truss element as a medium of connection as in Fig. 1. We have built two fully working Shady3D robots and 5 Shady3d bodies that do not include any electronics, but can be used as obstacles during our experiments, to simulate the presence of up to 7 robots working together on the truss. All algorithms are implemented and tested using this environment. The self-assembly operation requires many grasping steps for object detection and grasping. We choose an approach that embeds beacons in the passive object. Solutions that rely on other sensors such as vision are possible but require more computation"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002488_085106-Figure4-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002488_085106-Figure4-1.png",
        "caption": "Figure 4. Thread profile of the ballscrew.",
        "texts": [
            " (2) Fit and calculate according to the collected data in each screw profile data field to obtain the center coordinate of the ball track in each cycle and further to obtain the center coordinate of the ball. (3) Solve primary geometric parameters of the screw profile, such as effective diameter, thread pitch, ball track runout and ball track cross-section error based on the solved center coordinate of the ball track and the center coordinate of the ball. The thread profile of the ballscrew is shown in figure 4. The collected profile curve is composed of a ball track section, a fillet, a screw outer diameter section and a fillet, which appear periodically as shown in figure 5. Firstly, separate the data of the ball track section: select the data points on a thread pitch axially to find the highest point at the vertical axis in this cycle. The highest point must be located at the screw outer diameter. Then, appoint a data partitioning line along the horizontal axis under the screw outer diameter, while the distance between the partitioning line and the highest point is slightly bigger than the tolerance range of screw outer diameter (normally 0"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000540_amc.2008.4516037-Figure5-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000540_amc.2008.4516037-Figure5-1.png",
        "caption": "Fig. 5. Double inverted pendulum as simplified model",
        "texts": [
            " Attitude Stabilization Control by Internal Force in NullSpace Two-wheel inverted mobile manipulator is under-actuated system. From eq.(18), response of passive joint is given in eq.(29) considering acceleration of active joint q\u0308a as input. q\u0308p = \u2212M\u22121 pp (Mpaq\u0308a + hp + gp) (29) This equation is nonlinear differential equation, and complicated to deal with. So, double inverted pendulum with cart is adopted as simplified model of two wheel-inverted mobile manipulator. A model of double inverted pendulum is shown in Fig.5. This model regards a vehicle body of this robot as the first link of pendulum, and regards length from tip position of the first link to center of gravity (COG) of manipulator part as the second link. And it is regarded that the first link has no mass and the second link has total mass of manipulator part. Position and velocity of COG of two-wheel inverted mobile manipulator is caliculated by following equations. vxg = m0 vx0 + m1 vx1 + m2 vx2 + m3 vx3 m0 + m1 + m2 + m3 (30) vx\u0307g = mv 0x\u03070 + mv 1x\u03071 + mv 2x\u03072 + mv 3x\u03073 m0 + m1 + m2 + m3 (31) = Jgqm (32) , where xi and mi (i=0-3) denote tip position and mass of each link. And a matirx Jg which defined in eq.(32) is called COG Jacobian. In Fig.5, ld and qp denote length of the second link and inclination angle respectively. These are calculated as following equations. ld = \u221a (vxg \u2212 l0 sin qp)2 + (vzg \u2212 l0 cos qp)2 (33) qd = sin\u22121 ( vxg ld ) \u2212 qp (34) From motion equation of double inverted pendulum model, relation between angluar acceleration of passive joint q\u0308p and acceleration of cart wx\u0308cart is given in eq.(35). l0q\u0308p +w x\u0308cart + M mv + M ldq\u0308d + gqp = 0 (35) Desired angular acceleration of passive joint is given in eq.(36) to stabilize inclination angle of platform"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003292_e2013-01943-7-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003292_e2013-01943-7-Figure1-1.png",
        "caption": "Fig. 1. Two-bar truss (von Mises truss).",
        "texts": [
            " The main goal is the vibration reduction, avoiding some critical responses as snap-through behavior. A linear actuator is employed to help this control procedure and therefore, the SMA actuation is not employed for the control purposes. In this regard, we are investigating an SMA structure that needs an appropriate control using external actuators. It is important to highlight that SMA properties are being used to achieve other goals than control. This situation is common in distinct applications that include aerospace systems as self-erectable structures. The two-bar truss is depicted in Fig. 1. This plane, framed structure, is formed by two identical bars, free to rotate around their supports and at the joint. In the present investigation, we consider a shape memory two-bar truss where each bar presents the shape memory and pseudoelastic effects. The two identical bars have length L and cross-sectional area A. They form an angle \u03d5 with a horizontal line and are free to rotate around their supports and at the joint, but only on the plane formed by the two bars (Fig. 1). The critical Euler load of both bars is assumed to be sufficiently large so that buckling will not occur in the simulations reported here. We further assume that the structure\u2019s mass is entirely concentrated at the junction between the two bars. Hence, the structure is divided into segments without mass, connected by nodes with lumped mass that is determined by static considerations. We consider only symmetric motions of the system, which implies that the concentrated mass, m, can only move vertically",
            " Now, the following strain definition is considered, \u03b5 = L L0 \u2212 1 = cos\u03d50 cos\u03d5 \u2212 1 (4) with L0 and \u03d50 representing the nominal values of L and \u03d5, respectively. At this point, we can use the constitutive Eq. (2) together with kinematic Eq. (4) into the equation of motion (1), obtaining the governing equation of the SMA two-bar truss: mX\u0308 + cX\u0307 + 2A L0 X { [a1(T \u2212 TM )\u2212 3a2 + 5a3]+ + [\u2212a1(T \u2212 TM ) + a2 \u2212 a3]L0(X2 +B2)\u22121/2+ + [3a2 \u2212 10a3] 1L0 (X2 +B2)1/2+ + [\u2212a2 + 10a3] 1L20 (X 2 +B2)+ \u2212 5a3 L30 (X2 +B2)3/2 + a3 L40 (X2 +B2)2 } = P (t) (5) where B is the horizontal projection of each truss bar (Fig. 1). Considering a periodic excitation P = P0 sin(\u03c9t), Eq. (5) may be written in non-dimensional form as x\u2032 = y y\u2032 = \u03b3 sin(\u03a9\u03c4)\u2212 \u03bey + x{\u2212 [(\u03b8 \u2212 1)\u2212 3\u03b12 + 5\u03b13]+ + [(\u03b8 \u2212 1)\u2212 \u03b12 + \u03b13](x2 + b2)\u22121/2 \u2212 [3\u03b12 \u2212 10\u03b13](x2 + b2)1/2+ + [\u2212\u03b12 + 10\u03b13](x2 + b2) + 5\u03b13(x2 + b2)3/2 \u2212 \u03b13(x2 + b2)2 } (6) where \u03be is a non-dimensional viscous damping coefficient. The dissipation due to hysteretic effect may be considered by assuming an equivalent viscous damping related to this parameter. Moreover, the following non-dimensional parameters are considered: x = X L , \u03b3 = P0 mL0\u03c9 2 0 , \u03c920 = 2Aa1TM mL0 , \u03a9 = \u03c9 \u03c90 , \u03c4 = \u03c90t, \u03b8 = T TM , \u03b12 = a2 a1TM , \u03b13 = a3 a1TM , b = B L0 and (\u00b7)\u2032 = d(\u00b7) d\u03c4 \u00b7 According to Bessa and Barre\u0302to [22], adaptive fuzzy inference systems could be combined with smooth sliding mode controllers to improve the overall performance of the control system"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000548_j.jsv.2007.08.001-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000548_j.jsv.2007.08.001-Figure1-1.png",
        "caption": "Fig. 1. Analytical model.",
        "texts": [
            " A phenomenon that the pre-existing non-synchronous whirl/whip resulted from the instability of one shaft can activate the onset of oil instability of its neighboring shaft is revealed [23]. In this paper, the bifurcation and chaos characteristics of a flexible rotor system with two unbalanced disks are investigated using the maximum Lyapunov exponent. The results show that the maximum Lyapunov exponent is a valid method in identifying the bifurcation and chaos characteristics for rotor-bearing system. Experimental result carried on a test rig supports partly the theoretical analysis. Fig. 1 shows the model of a flexible rotor system with two unbalanced disks. The mass of flexible shafts is concentrated to the disks and the ends, k1, k2 and k3 denote the bending stiffness of three spans of shaft, respectively. O2 and O3 are geometrical center of the disks 2 and 3, O1 and O4 are center of left and right axle neck, respectively; c2 and c3 are the centers of masses of disks 2 and 3, bearings 1 and 4 are the journal bearings; F is the phase angle between the eccentricity of disks 2 and 3"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003470_j.ijengsci.2012.02.002-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003470_j.ijengsci.2012.02.002-Figure1-1.png",
        "caption": "Fig. 1. Contact of an elastic half-space with a rigid cylinder.",
        "texts": [
            " Section 3 describes the solution procedure that includes reformulation of the boundary-value problem in the bipolar coordinates, application of the Papkovich\u2013Neuber general solution and Fourier integral transform to obtain the distributions of normal and shear contact stresses. Non-adhesive full stick contact problem is discussed in Section 4. Finally, Section 5 provides a simple analytical solution to the adhesive full stick contact problem. Consider an elastic half-space indented by a rigid cylinder of radius R (Fig. 1). The force P applied to the center of the cylinder is normal to the surface of the half-space, and the half-space deforms in the state of plane strain. Assume that the coefficient of friction between two contacting surfaces is large enough that there are no relative displacements of the points on the surfaces in the contact area a 6 x 6 a (i.e., the so-called full stick contact takes place). For finite values of the friction coefficient, two slip zones adjacent to the stick zone will arise, if the elastic properties of the contacting materials are different"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000395_1.2736451-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000395_1.2736451-Figure1-1.png",
        "caption": "Fig. 1 Hole\u2013entry journal bearing system",
        "texts": [
            ", rough bearing and smooth journal) with a transverse pattern was found to partially compensate for the loss in load-carrying capacity due to the thermal and/or non-Newtonian behavior of lubricant effects. It limits 18.86% loss in load-carrying capacity due to the thermal effect to only 1.6% and 33.99% loss due to the combined influence of non-Newtonian lubricant and thermal effect to 16.76%. DOI: 10.1115/1.2736451 Keywords: surface roughness, non-Newtonian lubricant, journal bearing, hydrostatic ntroduction The applications of hole-entry hybrid journal bearings Fig. 1 re quite wide and varied due to their superior performance, such s improved load-carrying capacity at zero and high speed with ow energy consumption and relative simplicity in manufacturing s compared to conventional recessed or pocketed bearings 1 . ver the past several years, the incorporation of many physical ffects into the analysis of fluid-film bearings has provided much ore realistic performance data. In particular, the familiar asumptions of a smooth surface, isothermal operating condition, nd Newtonian behavior of the lubricant can no longer be emloyed to accurately predict the performance of fluid-film bearing ystem"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001361_aero.2009.4839615-Figure4-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001361_aero.2009.4839615-Figure4-1.png",
        "caption": "Figure 4 \u2013The dimensionless forces and moments",
        "texts": [
            " could be obtained by ( ) ( ) ( ) 0 0 1 0 1 \u02c6 l k k k k k W l k k k W W x x x W W = = = + \u0394 \u22c5 + \u0394 + \u0394 \u2211 \u2211 where 0kW , 0kx are the initial weights and C.G. locations, kW\u0394 , kx\u0394 are the updated data. ii. C.G. Identification Based on Neural Networks For active C.G. control aircraft, most of the flight is performed along straight paths, with relatively short periods of heading and airspeed changes [5][12]. Accordingly, the aircraft could be assumed as trimmed, and analysis of trimmed flight can be used to full effect. Figure 4 shows the relations between forces and moments on the aircraft [12]. LC , DC , WC , TC are the dimensionless lift, drag, gravity and propulsion forces. \u03b8 , \u03b1 , T\u03b1 are the pitch attitude angle, the angle of attack and the thrust inclination angle. The following analysis is carried out in the body-fixed reference frame. In a trimmed flight, force and pitch moment on the aircraft satisfy [12] , sin 0 cos 0 0 X W Z W M O Z cg X cg C C C C C C x C z \u03b8 \u03b8 \u2212 = + = + \u2212 = (30) where ( ) , , sin cos cos cos sin sin sin cos X L D T T Z L D T T a M O M O T T T T T T C C C C C C C C C C C x C z \u03b1 \u03b1 \u03b1 \u03b1 \u03b1 \u03b1 \u03b1 \u03b1 = \u2212 + = \u2212 + + = + + in which XC , ZC are the combined forces along x and z body-axis respectively, ,M OC is the combined aerodynamic and propulsion moment about point O"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000150_robot.2005.1570677-Figure4-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000150_robot.2005.1570677-Figure4-1.png",
        "caption": "Fig. 4 Schematic illustration of the electrodes and the coordinate system for the theoretical evaluation",
        "texts": [
            " The first is the effect of the MRI magnetic field on the motor\u2019s output force capability and the second is the effect of the motor\u2019s magnetic field on the MRI. For the purposes of this discussion, the motor is assumed to consist of only one pair of slider and stator films (no stacking), the stator has the same structure and the same size as the slider, and the gap between the slider and the stator is negligible (meaning the slider and the stator occupy the same location). The motor is also assumed to be operating at a constant speed. Figure 4 shows the coordinate system used in this chapter. The coordinate system, called the motor coordinate system, has its origin at the center of the motor. The x-axis is defined along the motor\u2019s operation direction, the y-axis along the lateral direction, and the z-axis along the normal direction. A surrounding magnetic field can affect the slider\u2019s motion in two different ways. It can attract paramagnetic materials within the motor and it can generate Lorenz forces on the currents flowing in the slider",
            " Because the motor contains only minute quantities of materials with magnetic susceptibilities, the former effect is negligible. Therefore, the effect of the Lorenz force is primarily discussed. When currents flow inside a slider that is placed in a magnetic field, the currents induce a Lorenz force on the slider. The total Lorenz force on the slider can be calculated if the current distribution in the slider is known. The electrodes of the slider film, which are located on both sides of the film and are schematically illustrated in Fig. 4, are a combination of three-phase parallel electrodes and feeding electrodes (\u03b1, \u03b2, \u03b31, \u03b32). Due to spatial limitations, the thirdphase electrodes are connected alternately to two feeding electrodes (\u03b31, \u03b32) while the electrodes for the other two phases are connected to identical electrodes (\u03b1, \u03b2), respectively. Therefore, this three-phase electrode pattern repeats every six electrodes. Since the operating frequency of this motor is relatively low (typically, up to 1 kHz) and the transitional signal reflections and propagations can be neglected, the current distribution in each three-phase electrode can be estimated as depicted in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003087_j.proeng.2011.05.057-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003087_j.proeng.2011.05.057-Figure1-1.png",
        "caption": "Fig. 1. Experimental set up in RMIT Industrial Wind Tunnel",
        "texts": [
            " The developed experimental testing methodology will allow evaluating not only the aerodynamic properties of bicycles and the cyclist but also various add-ons including cycling suit and helmet. The developed methodology will allow evaluating the aerodynamic relationship between various cycling garments and the cyclist under a wide range of wind speeds and yaw angles to include crosswinds effects. In order to have a reliable and accurate measurement system for aerodynamic properties, a full scale experimental setup has been developed at RMIT University. The developed setup and the measurement procedure are shown in Figure 1(a). The setup consists of a flat wooden platform (1800 mm 850 mm 30 mm) and a stand to support the bicycle and cyclist with the wooden platform firmly. The gap between the wooden platform and the tunnel floor is around 20 mm in order to avoid any interference between the floor and the wooden platform. A plastic fairing (shown in Figure 1(b)) is used at the front of the platform to minimize the flow separation from the leading edge of the platform. The whole platform is mounted on a 6-component force sensor (type JR3) via a 100 mm diameter strut (see Figure 1(a)) to measure the drag, lift and side forces and their corresponding moments simultaneously. All types of bicycles (recreational, road racing, time trial) along with the cyclist can be experimentally evaluated using this setup. The crosswind effects can also be evaluated using the arrangement. The developed system minimizes error in data recording due to extraneous cyclist movement or variations in weight distribution. The assembly is strong for both static (cyclist with no pedaling) and dynamic (cyclist with pedaling) loading",
            " One of the main difficulties in the full scale testing is to keep the setup and the cyclist\u2019s body position at the same reference point during the test. A small position variation can generate significant errors in data acquisition. In order to address this problem, a video positioning system has also been developed. It consists of two high definition digital video cameras which are installed at 180\u00b0 and 90\u00b0 from the bicycle longitudinal axis, i.e. imaginary axis joining front and back wheel centers (shown in Figure 1(a)) for capturing the front and side views of the entire experimental setup simultaneously. Special software is used for live video monitoring. The cyclist position can be repositioned accurately by overlapping the images taken by these two digital cameras and by adjusting necessary feedback obtained through the monitoring system. It can minimize any error occurred due to the change of positions of the cyclist and equipment as minor position variation can significantly affect the measured aerodynamic data"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003520_j.euromechsol.2011.11.003-Figure9-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003520_j.euromechsol.2011.11.003-Figure9-1.png",
        "caption": "Fig. 9. Maximally regular T2R1-type parallel manipulator with bifurcated planarspatial motion of the moving platform: constraint singularity (a), branch with planar motion (b) and branch with spatial motion (c); limb topology PkRkRkRtRPassPatcstPtkPtRtRSR.",
        "texts": [],
        "surrounding_texts": [
            "The term of constraint singularity (CS) has been recently coined (Zlatanov et al., 2002) to characterize the configuration of lower mobility parallel manipulators in which both the connectivity of the moving platform and the mobility of the parallel mechanism increase their instantaneous values. From a constraint singularity, the mechanism can get out with or without branching (Gogu, 2008b). When branching occurs in a constraint singularity, the mechanism can reach different configurations, called branches, and have different independent motions of the moving platform. In this case, the constraint singularity is also called branching or bifurcation singularity. A branch refers to the free-of-singularity configurations of the mechanism inwhich each structural parameter keeps its value. For this reason, this value is called global of full-cycle value for a branch. Two types of branching in constraint singularity (BCS) have been defined in Gogu (2008b). Branching of type BCS1 occurs when a parallel mechanism F ) G1 . Gj . Gk gets out from a constraint singularity in different branches characterized by the same mobility and the same connectivity of the moving platform but with different bases of the vector space of relative velocities between the moving and fixed platforms. Branching of type BCS2 occurs when a parallel mechanism F ) G1 . Gj . Gk gets out from a constraint singularity in different branches characterized by distinct values of mobility and connectivity of the moving platform. The parallel manipulators presented in this paper have constraint singularities with branching of type BCS1. Branching of type BCS2 occurs in kinematotropic mechanisms (Gogu, 2009c). The finite displacements and the velocities in the actuated joints are denoted by qi and _qi, the linear velocities of the characteristic point H of the moving platform, by v1 \u00bc _x; v2 \u00bc _y and the angular velocity of the moving platform by u \u00bc ua \u00bc _a or u \u00bc ud \u00bc _d: In both branches, the moving platform undergoes two planar translations and one rotation but the rotation axis is different in the two branches. In the first case, the rotation velocity u \u00bc ua \u00bc _a is parallel to the x-axis and the moving platform undergoes a spatial motion. In the second case, the rotation velocity u \u00bc ud \u00bc _d is parallel to the z-axis and the moving platform undergoes a planar motion. With these notations, the linear mapping between the actuated joint velocity space and the end-effector velocity space for a branch the T2R1-type PMs is defined by: 2 4 v1 v2 u 3 5 \u00bc \u00bdJ 2 4 _q1 _q2 _q3 3 5 (8) where J is the Jacobian matrix. In the constraint singularity, the T2R1-type parallel manipulators with bifurcated motions has instantaneously iM \u00bc iSF \u00bc 4 and (iRF)\u00bc(v1, v2, ua, ud). The bifurcation in this constraint singularity can be used to change motion type of the moving platform. In the two distinct branches, the parallel mechanism is characterized by M\u00bc SF \u00bc 3 and (RF)\u00bc(v1, v2,ua) or (RF)\u00bc(v1, v2,ud). In the constraint singularity, both connectivity of moving platform and mobility of parallel manipulator increase their instantaneous values. More details about the behaviour of parallel manipulators in a constraint singularity and the associated structural parameters can be found in Gogu (2009c)."
        ]
    },
    {
        "image_filename": "designv11_12_0001986_1.3428494-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001986_1.3428494-Figure1-1.png",
        "caption": "FIG. 1. An illustration of two cylinders in contact; a the top view x-y and b the front view z-y .",
        "texts": [
            " An image processing analysis is used to determine the size of the contact area under different loading conditions. The approximate JKR theory for elliptical contacts10 is taken as basis for data analysis and its consistency with the experiments is discussed in detail. In this section, the results of the approximate JKR theory for elliptical contacts are given for two cylinders in contact in an opportune form which will be used in the experimental section. When two adhesive cylinders are brought into contact as seen in Fig. 1, normal load W and indentation depth equations in nondimensional form are expressed as follows:10 W\u0304 = 8 3 g 1 \u2212 g 2 \u0304g2 \u2212 \u0304 2 \u0304 \u2212 \u0304g5/2 1 \u2212 g \u2212 1 3 \u0304g2 + \u0304 , \u0304 = 27/2 9 2 2/3 g 1 \u2212 g \u0304g2 \u2212 \u0304 4/3 2K e \u0304 \u2212 \u0304g5/2 1 \u2212 g \u2212 \u0304B e \u2212 g2\u0304D e , 1 where K e , B e , and D e are complete elliptic integrals a Author to whom correspondence should be addressed. Electronic mail: cagdas@cmu.edu. 0021-8979/2010/107 11 /113512/7/$30.00 \u00a9 2010 American Institute of Physics107, 113512-1 [This article is copyrighted as indicated in the article",
            " Therefore, we resort to a method where two identical cylinders can be used for all of the experiments. If two identical cylinders are brought into contact at 90\u00b0 with respect to each other cross cylinders , the profile of the contact area would be circular. In the limiting case, one can arrange the cylinders parallel to each other, where the contact zone takes a rectangular shape. If the cylinders are brought into contact at different angles within the limits above, the contact area is expected to have an elliptical shape as illustrated in Fig. 1 a . To facilitate the approximate elliptical JKR theory, it is necessary to find an expression that relates to the skew angle , the angle between the identical cylinders. Using the Hertz theory for the compression of elastic bodies, the surfaces can be represented by second degree of curvature, which is given by following equations in the vicinity of the point of contact as shown in Fig. 1 b Refs. 1 and 12 z1 = A1x2 + A2xy + A3y2, z2 = B1x2 + B2xy + B3y2. 4 The distance between two points M and N in Fig. 1 b on the contacting bodies is then given as z1 + z2 = A1 + B1 x2 + A2 + B2 xy + A3 + B3 y2. 5 By properly choosing the axes, the xy term can disappear such as z1 + z2 = Ax1 2 + By1 2 = 1 2Ra x1 2 + 1 2Rb y1 2. 6 There is a unique relation between the principal relative radius of curvatures Ra and Rb , and with the principal radii of curvature of the two bodies R1 , R1 , R2 , and R2 as follows1,12 1 Ra + 1 Rb = 1 R1 + 1 R1 + 1 R2 + 1 R2 , 1 Rb \u2212 1 Ra = 1 R1 \u2212 1 R1 2 + 1 R2 \u2212 1 R2 2 + 2 1 R1 \u2212 1 R1 1 R2 \u2212 1 R2 cos 2 1/2 "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003275_robio.2012.6490973-Figure3-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003275_robio.2012.6490973-Figure3-1.png",
        "caption": "Fig. 3. Physical models",
        "texts": [
            " The aerodynamic model (ADM) is a differential equation which shows the motion of the ball flying in the air with rotational velocity. In this section those physical models will be shown briefly. The base coordinate system \u03a3B is set at a corner of the table (see Fig. 2) through this paper. Suppose that the racket strikes the ball with translational velocity V \u2208 R3 and posture of yaw angle \u03b1 \u2208 [\u2212\u03c0 2 , \u03c0 2 ] and pitch angle \u03b2 \u2208 [0,\u03c0]. Denote the ball\u2019s translational and rotational velocities just before and after the racket strikes the ball respectively by v0 \u2208 R3, \u03c90 \u2208 R3 and v1 \u2208 R3, \u03c91 \u2208 R3(see Fig. 3(a)). Then the relation between (v0,\u03c90) and (v1,\u03c91) can be shown as follows.[ v1 \u2212V \u03c91 ] = RRRM(\u03b1 ,\u03b2 ) [ v0 \u2212V \u03c90 ] (1) where RRRM(\u03b1,\u03b2 ) = [ RR 0 0 RR ][ Avv Av\u03c9 A\u03c9v A\u03c9\u03c9 ][ RR 0 0 RR ]T , RR = \u23a1 \u23a3 cos\u03b2 sin\u03b2 sin\u03b1 sin\u03b2 cos\u03b1 0 cos\u03b1 \u2212sin\u03b1 \u2212sin\u03b2 cos\u03b2 sin\u03b1 cos\u03b2 cos\u03b1 \u23a4 \u23a6 , Avv = diag(1\u2212 kv,1\u2212 kv,\u2212er), Av\u03c9 = kvrS12, A\u03c9v =\u2212k\u03c9 rS12, A\u03c9\u03c9 = diag(1\u2212 k\u03c9 r2,1\u2212 k\u03c9 r2,1), and S12 = \u23a1 \u23a3 0 1 0 \u22121 0 0 0 0 0 \u23a4 \u23a6 . Note that RR is a rotational matrix from \u03a3B (the table coordinate) to \u03a3R (the racket coordinate) with yaw \u03b1 and pitch \u03b2 . The parameter r = 2\u00d710\u22122[m] is the radius of the ball, and the parameters kv = 6.15\u00d7 10\u22121, k\u03c9 = 2.57\u00d7 103 and er = 7.3\u00d710\u22121 are obtained by experimental data. See [5], [9] for details about how to identify the parameters and how well RRM works. When a ball flies in the air with rotational velocity \u03c9 (see Fig. 3(b)), the ball motion is given as follows. p\u0308(t) =\u2212g\u2212CD(t) \u03c1 m Sb\u2016p\u0307(t)\u2016p\u0307(t)+CM(t) \u03c1 m Vb\u03c9\u00d7 p\u0307(t) (2) where p(t) \u2208 R3 is the ball\u2019s position at the time t, g = [0,0,g]T with the acceleration of gravity g= 9.8 [m/s2], Sb = 1 2 \u03c0r2, and Vb = 4 3 \u03c0r3. Notice that \u03c9 = [\u03c9x,\u03c9y,\u03c9z] T \u2208 R3 is the ball\u2019s rotational velocity which is here assumed constant during the ball flies in the air. The parameters m and \u03c1 are the ball\u2019s mass and the air density; m = 2.7\u00d710\u22123[kg] and \u03c1 = 1.184[kg/m3](25\u25e6C). CD(t) and CM(t) in (2) are respectively drag coefficient and Magnus coefficient, which are given by CD(t) = aD +bDh(p\u0307(t),\u03c9) (3) CM(t) = aM +bMh(p\u0307(t),\u03c9) (4) where h(p\u0307(t),\u03c9) = 1\u221a 1+ ( p\u03072 x+p\u03072 y)\u03c92 z (p\u0307x\u03c9y\u2212p\u0307y\u03c9x)2 and aD = 0"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001931_iros.2010.5649323-Figure16-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001931_iros.2010.5649323-Figure16-1.png",
        "caption": "Fig. 16. The Flow Field by using the Conformal Transformation",
        "texts": [
            " Before the start of the simulation, the obstacle stands the right side, and the mobile robot stands the left side in the simulation display area. Moreover, the goal is set at the right side. The moving obstacle moves to the left direction and the mobile robot moves to the right direction after starting this simulation. If the distance between two circles is long, it is shown that the velocity is fast. In the intersection on near the center of display area, it can be seen that the mobile robot can avoid the moving obstacle. However, the velocity of the mobile robot accelerates rapidly when it avoids a moving obstacle. Fig. 16 shows the simulation with the new method by using CT. That is, the ellipse flow field is used in this simulation. The initial conditions of Fig. 16 is the same as those of Fig. 15. In comparison of Fig. 15, Fig. 16 and Fig. 17, the difference of their velocities is shown. It can be seen that the maximum velocity of the new method is a half maximum velocity of our previous method. As a result, the effectiveness of our method by using the ellipse flow field can be shown. Fig. 18 shows the simulation with applying the correction function bj(z) in Fig. 15. And Fig. 19 shows the simulation with applying correction function bj(z) in Fig. 16. Similarity, Fig. 20 also shows the difference of their velocities. It can be seen that the case of Fig. 19 is smoother than the case of Fig. 18. The blue line as shown in Fig. 17 has a discontinuous point at the maximum velocity point. However, the blue line as shown in Fig. 20 has no discontinuous point at all time. As a result, the effectiveness of our method by using the new correction function can be shown. Is the robot motion considered? If the robot is just in front of the obstacle, how does the robot move",
            " 20 come from both the ellipse field and the correction function, and they are only influenced by the parameters of m and uj . However, the m and the uj are the same value, respectively in each simulation. The m denotes a sink value and the standard robot velocity comes from the m. That is, the m can not be changed. This can be seen by the vertical constant velocity lines shown in Fig. 17 and Fig. 20. On the other hand, the uj denotes the obstacle velocity. That is, the uj can not be changed too. All distances of many circles of the moving obstacles shown in both Fig. 15 and Fig. 16, and also both Fig. 18 and Fig. 19 are the same, respectively. Therefore, the changing velocities of these simulations come from the other factors. The factors are using both the ellipse field and the correction function. In this paper, we proposed the improved method by applying the Conformal Transformation and the new correction function to our previous Hydrodynamic Potential method for path planning of a mobile robot to avoid the moving obstacle smoothly. A mobile robot can gradually avoid a moving obstacle from further away, and can be safely guided without rapid acceleration"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002400_1.4006324-Figure23-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002400_1.4006324-Figure23-1.png",
        "caption": "Fig. 23 Tooth deformation of the right web gear (web angle 5 0 deg)",
        "texts": [
            " Figures 20(a)\u201320(d) show images of deformed left web gears with web angles of 0, 15, 30, and 45 deg, respectively. Figures 21(a)\u201321(d) show images of the deformed center web gears with web angles of 0, 15, 30, and 45 deg, respectively. Figures 22(a)\u201322(d) show images of the deformed right web gears with web angles of 0, 15, 30, and 45 deg, respectively. To understand the deformation characteristics of the loaded tooth of the thin-rimmed gear well, an image of the deformed right web gear with a web angle of 0 deg is shown in Fig. 23(a), and an enlarged view of the loaded tooth is also shown in Fig. 23(b). Figure 23 aids in explaining the deformation characteristics of the loaded tooth of the thin-rimmed right web gear with a web angle of 0 deg in the following discussion. Based on Fig. 23, the tooth positions before deformation and after deformation are sketched in Fig. 24(a). From Fig. 24(a), the deformation of the loaded tooth can be roughly divided into two types of deformation: one type is an upward and downward deformation of the loaded tooth as shown in Fig. 24(b), and the other type is a rotation deformation of the loaded tooth as shown in Fig. 24(c). In Fig. 24(b), end A of the loaded tooth has a downward deformation, and end B has an upward deformation. This is because end A is further away from the web than end B, so end A is more flexible than end B"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003391_icems.2011.6073503-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003391_icems.2011.6073503-Figure2-1.png",
        "caption": "Fig. 2 Cross section diagram of the radial MFM-BDRM",
        "texts": [
            " It is composed of the radial magnetic-field-modulated brushless double-rotor machine (MFM-BDRM) and the torqueregulating machine, which can provide the speed and torque difference between the ICE and the final gear, respectively. The torque-regulating machine is normally a traditional permanent-magnet machine. And this paper mainly focuses on the radial MFM-BDRM. II. PRINCIPLE ANALYSIS OF THE RADIAL MFM-BDRM A. Theoretical Analysis of Magnetic Field and Speed equation The radial MFM-BDRM consists of three components: the first stator, the modulating ring rotor and first PM rotor, as shown in Fig. 2. It is assumed that the pole-pair number of the first PM rotor is n and that of the rotary magnetic field of the first stator is p with the first-stator windings fed with AC current. Meanwhile, it is supposed that the numbers of magnetic blocks and nonmagnetic blocks in the modulating ring rotor are both q. Suppose the rotary speeds of the first PM rotor, the firststator magnetic field and modulating ring are \u21261, \u21262, \u21263, respectively. The magnetomotive force (MMF) produced by the first PM rotor is given by ( ) ( ) ( )2 2 1 2 1,3,5,"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001345_robot.2008.4543604-Figure9-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001345_robot.2008.4543604-Figure9-1.png",
        "caption": "Figure 9. The position control",
        "texts": [],
        "surrounding_texts": [
            "N i diidiip qtqtte 1\n(27)\n(28) N\ni qiip tetete 1\nThe position control of the arm means the motion control from the initial position to the desired position in order to minimize the error. 0C bC Theorem 1. The closed-loop control system of the position (6), (7), (12) is stable if the fluid pressures control law in the chambers of the elements given by:\ntektekatu i j ii j ijiji 21 (29)\ntektekbtu qi j qiqi j qijiqji 21 , (30)\nwhere ; , with initial conditions: 2,1j Ni ,,2,1\n000 211121 iiiii ekkpp (31)\n000 211121 qiqiqiqiqi ekkpp (32)\n00ie (33) 00qie , (34)Ni ,,2,1\nand the coefficients , , , are positive and verify the conditions\nik qik mn ik mn qik\n2111 ii kk ; (35)2212\nii kk 2111 qiqi kk ; , (36)2212\nqiqi kk Ni ,,2,1 Proof. See Appendix 1.\nB. Force Control The grasping by coiling of the continuum terminal elements offers a very good solution in the fore of uncertainty on the geometry of the contact surface. The contact between an element and the load is presented in Figure 7. It is assumed that the grasping is determined by the chambers in -plane.\nThe relation between the fluid pressure and the grasping forces can be inferred for a steady state from [2],\n8\n~~~~\n21\n0 00 2\n2\ndSpp\ndssTsTsfds s\nsk l s T l\n(37)\nwhere sf is the orthogonal force on the curve ,bC sf is sF in -plane and sFq in q-plane, respectively.\nA spatial discretization is introduced and121 ,,, lsss\nii ss 1 , ii s , 1,,2,1 li (38) For small variation i around the desired position id , in -plane, the dynamic model (6) can be approximated by the following discrete model [11],\neiiidid\ndidiidiiiii FfdqH qHcm , ,, , (39)\nwhere Smi , 1,,2,1 li , did qH , is a nonlinear function defined on the desired position did q, ,\ndiii qcc ,, , ,0ic q, (40) is the viscosity of the fluid in the chambers.\nididii i\ndiddidiidi\nqh H\nqHqH\nd\n,\n,,,\n(41)\nand is the external force due to the environment, the load. The equation (39) becomes, eiF\neiii\nididiidiiii Ffd qhqcm ,,,\n(42)\nThe aim of explicit force control is to exert a desired force . If the contact with load is modeled as a linear spring with constant stiffness , the environment force can be modeled as: idF Lk\niLei kF (43) The error of the force control may be introduced as\nidiefi FFe (44) It may be easily shown that the equation (42) becomes\nidi i iifii i\nfi L\ni fi\nL\ni Fd k h fded k h e k c e k m (45)\nTheorem 2. The closed force control system is asymptotic stable if the control law is\nidiLifiiiLi iL\ni Fdkhemdkh dk\nf 21 (46)\nii mc (47) Proof. See Appendix 2.\nIn this paper, the force error control may be achieved by using the Direct Sliding Mode Control (DSMC) [12]. This method establishes conditions that force the trajectory along the switching line, directly toward the origin. The block scheme of the force control is presented in Figure 8.",
            "A hyperredundan anip lator with eight elements is considered. The m are: linear density\nV. SIMULATION\nt m u echanical parameters\nmkg2.2 and the length of one element is ml 05.0 . The control problem in the -plane will be analyzed.\nThe initial position is the defined by : 00 2sC and or a circular load the grasping function is performed f\ndefined by 0 2 0: ryyxxCb , where22\nyx , represent the coordinates in -plane. A ent with an incrementdiscretisation for each elem 3l is\nintroduced. A control law (29) is used and a MATLAB system is applied. The result is presented in Figure\nA force control for the grasping terminals is simulated. The phase portrait of the force error is presented in Fi\n9.\ngure 10. First, the control (29) is used and then, when the trajectory penetrates the switching line, the viscosity is increased for a damping coefficient 15.1 .\nVI. CONCLUSION\nThe paper treats cont f a hyperredundant robot with contin ele erforms the coli fu\nhe difficulties de\nWe consider the following Lyapunov function [13],\nthe rol problem o uum ments that p\nnction for grasping. The structure of the arm is given by flexible composite materials in conjunction with activecontrollable electro-rheological fluids. The dynamic model of the system is inferred by using Lagrange equations developed for infinite dimensional systems.\nThe grasping problem is divided in two subproblems: the position control and force control. T\ntermined by the complexity of the non-linear integraldifferential equations are avoided by using a very basic energy relationship of this system and energy-based control laws are introduced for the position control problem. The force control is obtained by using the DSMC method in which the evolution of the system on the switching line is controlled by the ER fluid viscosity. Numerical simulations are presented.\nAppendix 1\nqiqiii tektektVtTt 22 2 1 (A.1.1)\nN\ni W 1\nwhere T, V represent the kinetic and potential energies of the system. tW is positive definite because the terms that represent the e ergy T and V are always 0tT ;n 0tV .\nBy using (5), the derivative of this function will be l\nq dstsqtsFtstsFt ,,,,\nN i qiqiqiiii tetektetek\nW\n1\n0 (A.1.2)\nFrom (8)-(11), the relation (A.1.2.) can be rewritten as\niii ttptpSdtW 21\nN\ni qiqiqiiii\nN\ni iqiqi\ntetektetek\ntqtptp\n1\n1\n21\n8 (A.1.3)\nThe control law (29), (30) with the initial conditions (31)- (34) determines the pressures of fluid in the chambers,\ntektektp i j ii j i j i 21 , 2,1j (A.1.4)\ntektektp qi j qiqi j qi j qi 21 , (A.1.5)2,1j\nSubstituting these solutions in (A.1.3) we obtain\nN\ni qiqiqiiii\nN\ni qiqiqiiii\nN\ni qiqiqiqi\nN\ni\ntetektetek\ntekktekkSd\ntetekkSd\nW\n1\n1\n2211122212\n1\n2111\n1\n8\n8 (A.1.6)\nIf\niiii tetekkSdt 2111\n8\n2111 8 iii kkSdk , 2111 8 qiqiqi kkSdk\nerified, the func\nand the\nconditions (35), (36) are v tion tW becomes:",
            "8 221112212 0\n1\n2 N\nqiqiqiiii tekktkkSdt\nAppendix 2 An error measure of the control system is intro the parameter\ni\neW (A.1.7)\nQ.E.D.\nduced by\nee (A.2.1) .and e e\nThe equation of the force error (56) becomes\nidi i ii Fd k h fd L\niii\nL\nmhm k 2\n(A.2.2)\nA Lyapunov function of V is introduced\nL i LL e k d kk ii cm\n2 2 1V ( The derivative of V will be (A.2.4)\nand subs becomes\nA.2.3)\nV tituting the variable from (A.2.2), this relation\nid i\niL\ni\ni ii\nL dkh fd m k\n(A.2.5)\ni\ni\nF mm\nm 2\nBy using the conditions of Theorem 2, (46), (47), this relation will be:\nL i Li e k d km ii hc V 2\ni\ni m c V 2 (A.2.6)\nREFERE\nami, Design of light weight flexible robot arm, Robots 8 Conference Proceedings, Detroit, USA, June\n1640. gne, Ian D. Walker, On the kinematics of\nremotely - actuated co um robots, Proc. 2000 IEEE\n[5] ons, IEEE\n[6] f a\n[7] Shimeura, H. Kobayashi, Direct\n1658.\n[9] lication to robot\n[10] nd Some\n[11] Conf. on\n[12] ,\n[13] bots,\n[14] b. and Aut,\n0V (A.2.7) Q.E.D.\nNCES\n[1] A. Hem\n1984, pp. 1623- [2] Ian A. Grava ntinu Int. Conf. on Robotics and Automation, San Francisco, April 2000, pp. 2544-2550. G. S. Chirikjian, J. W[3] . Burdick, An obstacle avoidance algorithm for hyper-redundant manipulators, Proc. IEEE Int. Conf. on Robotics and Automation, Cincinnati, Ohio, May 1990, pp. 625 - 631. G.S.Chirikjian, A general n[4] umerical method for hyperredundant manipulator inverse kinematics, Proc. IEEE Int. Conf. Rob Aut, Atlanta, May 1993, pp. 107-112. G.S. Chirikjian, J. W. Burdick Kinematically optimal hyperredundant manipulator configurati\nTrans. Robotics and Automation, vol. 11, no. 6, Dec. 1995, pp. 794 - 798. H. Mochiyama, H. Kobayashi, The shape Jacobian o manipulator with hyper degrees of freedom, Proc. 1999 IEEE Int. Conf. on Robotics and Automation, Detroit, May 1999, pp. 2837- 2842. H. Mochiyama, E. kinematics of manipulators with hyper degrees of freedom and Serret-Frenet formula, Proc. 1998 IEEE Int. Conf. on Robotics and Automation, Leuven, Belgium, May 1998, pp. 1653- [8] G. Robinson, J.B.C. Davies, Continuum robots \u2013 a state of the art, Proc. 1999 IEEE Int. Conf. on Rob and Aut, Detroit, Michigan, May 1999, pp. 2849-2854. K. Suzumori, S. Iikura, H. Tanaka, Development of flexible microactuator and its app mechanisms, IEEE Conf. on Robotics and Automation, Sacramento CA, April 1991, pp 1564 - 1573. S.K. Singh, D.O. Popa, An Analysis a Fundamental Problems in Adaptive Control of Force, IEEE Trans. Rob and Aut., Vol. 11, No 6, pp 912-922. M. Ivanescu, V. Stoian, A variable structure controller for a tentacle manipulator, Proc. IEEE Int. Robotics and Aut., Nagoya, pp. 3155-3160, 1995. S.Chiaverini, B.Siciliano, L.Villani, Force and Position Tracking: Parallel Control with Stiffness Adaptation IEEE Contr.Syst, Vol. 18, No 1, pp. 27-33. S.S.Ge, T.H.Lee, G.Zhu, Energy-Based Robust Controller Design for Multi-Link Flexible Ro Mechatronics, No 7,Vol. 6, pp. 779-798, 1996 M. Ivanescu, Position dynamic control for a tentacle manipulator, Proc. IEEE Int. Conf. on Ro Washington, A1-15, May 2002, pp. 1531-1539."
        ]
    },
    {
        "image_filename": "designv11_12_0001336_amc.2008.4516091-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001336_amc.2008.4516091-Figure2-1.png",
        "caption": "Fig. 2. Two wheel vehicle",
        "texts": [],
        "surrounding_texts": [
            "Modeling is described in this section. The two wheel vehicle used in this paper has coaxial two wheeled system. Wheels include in Direct Current (DC) servo motors with encoders which are used to measure wheel angles. On this coaxial wheel system, vehicle body is mounted with three link manipulator. Each link is controlled using a DC servo motor with encoder. Joint between coaxial wheeled link and the body is a passive joint. The body connected through the passive joint, is also assumed as a manipulator link. Therefore, the upper structure is assumed as a four degree of freedom manipulator. The lower part is two wheeled system."
        ]
    },
    {
        "image_filename": "designv11_12_0001901_1.3591479-Figure6-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001901_1.3591479-Figure6-1.png",
        "caption": "Fig. 6 Skew four-bar mechanism wi th one turning pair and one spher ical pair on f i xed l ink",
        "texts": [
            " The corresponding curve in the (/, u) plane is of order 4 with 2 double points, and therefore of genus 1, so the locus of P is also of genus 1. Curves C-l l -C-16 in Table 1 illustrate the foregoing. The motion of points not in line with the centers of the spherical joints is not defined, because of the independent rotability of the coupler link about its axis. I11 special cases, case (i) reduces to spherical four-bar motion, as is well known [16]. Case (i i ) : One Turning Joint and One Spherical Joint F ixed . T h i s mechanism, which is not so well known, is shown in Fig. 6. The origin, 0, is on the axis of the fixed turning joint, which is the z-axis, at a distance b from the fixed spherical joint, B, which lies on the z-axis, xyz being right handed; BC = / . The axis of the moving crank, D, is offset a, a from the fixed-crank axis and its rotation relative to the z-axis is 6. The distance from the common perpendicular of the crank axes to the origin is p. The coupler, DC, has a component of length q along the D-crank axis, and a second component of length e, perpendicular to q; its angular position, <f>, is measured relative to AD"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001265_tmag.2009.2012576-Figure5-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001265_tmag.2009.2012576-Figure5-1.png",
        "caption": "Fig. 5. Magnet sample for measuring contact resistance. (a) Measured magnet and (b) magnet with lead wire.",
        "texts": [
            " The theoretical value of the total eddy current loss of the segmented magnet is obtained by assuming that each piece is completely insulated. The figure denotes that the measured values are almost in agreement with theoretical ones. This fact means that the total eddy current loss of the segmented magnet, of which the compressive force between magnets is a few MPa, can be reduced without insulation. It seems that the eddy current loss could be measured with an acceptable accuracy. We measured the resistance of the contact part between magnets. Lead wires are pasted by a solder near the edges of two magnets as shown in Fig. 5(b). Both magnets are put in a vice as shown in Fig. 6. Dc current is impressed. As shown in Fig. 7, three kinds of voltages between two points including the contact part are measured changing the position of probes. The compressive stress is measured using a load cell (Kyowa Dengyou, LCN-A-10kN). The distances between probes are set at 12 mm, 22 mm, and 32 mm, respectively. It can be assumed that the dc current distribution is uniform in these regions. Measurements are performed ten times and the averaged values are used"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003713_978-1-4614-2419-2-Figure4.2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003713_978-1-4614-2419-2-Figure4.2-1.png",
        "caption": "Fig. 4.2 Unit cell of a honeycomb core with added resonating structure. With (a) a 3D view, (b) a top view, (c) an enlarged section of the top view and (d) a top view indicating the different sections of the resonators: the wings (dark gray) and the centre (light gray)",
        "texts": [
            " Previous research by the authors indicated that stopbands achieved by the addition of a periodic grid of local resonators can show strong attenuation in wave propagation. This section introduces the kind of test panel chosen and investigates the stopband behaviour of such an infinite panel. clear resonance in the low frequency region. The resonator structures are added in the cells of the honeycomb core. 4 Numerical and Experimental Study of Local Cell Resonators. . . 37 The unit cell of the honeycomb core with added resonating structure is shown in Fig. 4.2. Table 4.1 lists the dimensions of the unit cell. The resonating structure comprises two parts with different thickness: the centre and the wings (Fig. 4.2d) The core of the honeycomb has a thickness of 3 mm, the centre of the resonating structure has 4.5 mm of thickness, while the wings have a 9 mm thickness. The resonator structure can be seen as a mass-spring structure where the centre is a three-legged spring and the wings act as mass. The structure is designed to have a first bending mode of the resonator (Fig. 4.3) below 1,000 Hz. When the resonator structure is modelled separately from the honeycomb-core, and the connection points are regarded as clamped, the resulting resonance frequency is 854 Hz when using the material parameters of isopreen (Table 4",
            " The added resonating structure increases the weight of the unit cell by 214%. Although this seems like a high number, one should keep in mind that in a final lay out, also the top and bottom layers of the honeycomb core should be taken into account, which reduces the relative mass addition of the resonator. Furthermore the added resonator structure is deliberately chosen to be heavy, so that the resulting stopband is large as well [6]. Fig. 4.3 First bending resonance mode of the resonating structure (854 Hz) Table 4.1 Dimension of the unit cell in Fig. 4.2 L H a b c d e 30 mm 15 mm 1 15 L 1 8 L 1 5 L 2 15 L 2 15 L 38 C.C. Claeys et al. Unit cell analysis allows the prediction of the stopband behaviour of the test panel. The unit cell is modelled using linear shell elements in an FE method framework (718 Nodes, 648 Elements). From a noise and vibration point of view, the out-of-plane movement of the honeycomb core is the most important, so only the dispersion curves corresponding to outof-plane movement are considered. Figure 4.4 shows the resulting dispersion curves"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000543_s10846-008-9295-5-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000543_s10846-008-9295-5-Figure1-1.png",
        "caption": "Fig. 1 This figure illustrates the helicopter\u2019s body frame, the TPP angles and the thrust vectors of the main and tail rotor",
        "texts": [
            " The second will be the body fixed reference frame defined as FB = {OB, iB, jB, kB} where the center OB is located at the Center of Gravity (CG) of the helicopter. In general the orientation of the orthonormal set of vectors { iB, jB, kB} is standard in most textbooks related with aerodynamics of air vehicles [18, 25] . This paper will follow this convention and iB at the front of the helicopter, jB is pointing at the right side, and kB is facing downwards and it is normal to both iB, jB. The direction of these vectors can be seen in Fig. 1. The linear velocity expressed in the body coordinate frame is vB = [vB x vB y vB z ]T \u2208 R 3. This is the velocity of the helicopter\u2019s CG with respect to the inertia frame and expressed in the coordinates of the body frame. The angular velocity with respect to the body frame is \u03c9B = [\u03c9x \u03c9y \u03c9z]T \u2208 R 3. Positive direction of the angular velocity components refers to the right-hand rule of the respective axis. Following the analysis in [11] we denote F B = [ f B \u03c4 B]T \u2208 R 6 as the external wrench acting on the CG of the helicopter, expressed in the body frame coordinates",
            " The TPP is characterized by two angles, a and b which represent the tilt of the TPP at the longitudinal and lateral axis respectively. The TPP is itself a dynamic system. The work presented in [20] and in [28] provide a simplified model of the TPP dynamics. The TPP first order dynamic equations can be found in [5] as: \u03c4ca\u0307 = \u2212a \u2212 \u03c4cq + Aculon (8) \u03c4cb\u0307 = \u2212b + \u03c4c p + Bculat (9) where \u03c4c is a time constant of the rotor dynamics which includes the effect of the stabilizer bar, and Ac, Bc are just gains. The inclination of the TPP can be seen in Fig. 1. The TPP dynamics are going to be considered very fast in comparison with the rigid body dynamics and only their steady state effect will be regarded. Then, regarding the TPP angles: a = Kaulon (10) b = Kb ulat (11) where Ka, Kb are constant parameters. The magnitude of the main and tail rotor thrust will be considered proportional to the collective control commands, therefore: TM = KMucol (12) TT = KTuped (13) where TM, TT are the magnitude of the forces of the main and tail rotor respectively while KM, KT are constant parameters"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001601_iros.2010.5650421-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001601_iros.2010.5650421-Figure1-1.png",
        "caption": "Fig. 1. Model of the dynamic walker with flat feet and compliant ankles.",
        "texts": [
            " Comparison of walking performance of different gaits could help us better understand the locomotion of real human walking, which is the main motivation of this paper. This paper is organized as follows. Section II presents the dynamic walking model in detail. Section III is devoted to describe the three typical walking gaits. Section IV gives the simulation results. We conclude in Section V. To study the motion characteristics of real human walking, we added flat feet and compliant ankle joints to the dynamic bipedal walking model. As shown in Fig. 1, the twodimensional model consists of two rigid legs interconnected individually through a passive hinge with a rigid upper body (mass added stick) connected at the hip. Each leg includes thigh, shank and foot. The thigh and the shank are connected at the knee joint, while the foot is mounted on the ankle with a torsional spring.A point mass at hip represents the pelvis. The mass of each leg and foot is simplified as point mass added on the Center of Mass (CoM) of the shank, the thigh, and the foot, respectively",
            " To simplify the motion, we have some assumptions, including shanks and thighs suffering no flexible deformation, hip joint and knee joints with no damping or friction, the friction between walker and ground is enough, thus the flat feet do not deform or slip, and strike is modeled as an instantaneous, fully inelastic impact where no slip and no bounce occurs. The passive walker travels forward on level ground with hip actuation. We suppose that the xaxis is along the ground while the y-axis is vertical to the ground upwards. The configuration of the walker is defined by the coordinates of the point mass on hip joint and some angles (swing angles between vertical axis and each thigh and shank, angle between vertical axis and the upper body and the foot angles between horizontal coordinates and each foot) (See Fig. 1), which can be arranged in a generalized vector q = (xh, yh, \u03b81, \u03b82, \u03b83, \u03b82s, \u03b81f , \u03b82f )T . The positive directions of all the angles are counter-clockwise. The model can be defined by the euclidean coordinates x, which can be described by the x-coordinate and y-coordinate of the mass points and the corresponding angles (suppose leg 1 is the stance leg): x = [xh, yh, xc1, yc1, \u03b81, xc2t, yc2t, \u03b82, xc3, yc3, \u03b83, xc2s, yc2s, \u03b82s, xc1f , yc1f , \u03b81f , xc2f , yc2f , \u03b82f ]T (1) The walker can also be described by the generalized coordinates q as mentioned before: q = (xh, yh, \u03b81, \u03b82, \u03b83, \u03b82s, \u03b81f , \u03b82f )T (2) We defined matrix T as follows: T = dq/dx (3) Thus T transfers the velocities of the euclidean coordinates x\u0307 into the independent generalized coordinates q\u0307"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002192_6.2009-2047-Figure4-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002192_6.2009-2047-Figure4-1.png",
        "caption": "Figure 4. Quadrotor frame with motors and landing struts.",
        "texts": [
            " Video Capture and Processing Path Navigation System RC Radio Stabilization Control Manual-Flight Backup System Navigation and Payload Control Communication Layer 1 .3 G H z 2 .4 G H z 7 5 M H z Ground Station Airborne System Figure 3. A high-level system view. American Institute of Aeronautics and Astronautics 3 limited. Cost, weight, performance, and power consumption were therefore carefully considered. The following subsections present in more detail the various system components. A. Vehicle Structure The frame of the OU quadrotor is shown in Figure 4. It consists of a magnesium hub joining four carbon fiber arms. Mounted at the end of each arm is a magnesium motor mount that holds a brushless motor and propeller assembly. Fixed underneath each motor mount is a landing strut made of flexible steel blades. The resulting physical structure weighs approximately 1400 grams and measures 40 * 40 cm (without propellers). The carbon fiber rods used have a tensile strength of 200,000 psi and a density of 0 1.49 gr/cm 3 . The carbon fiber structure is conductive, thus caution was practiced to make sure that no wiring came in contact with the frame"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002390_jmems.2012.2194777-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002390_jmems.2012.2194777-Figure2-1.png",
        "caption": "Fig. 2. Structure of a commercial ID with variable apertures for a liquid lens.",
        "texts": [
            " 1(d) shows the arrangement of one blade positioned on the base plate. By pushing the pin-2 in clockwise direction, the blade will rotate in clockwise direction, as shown in Fig. 1(e). Fig. 1(f) shows the arrangement of five blades positioned on the base plate. For the five blades, their pins pointed to the outside are positioned in the slides of the rotatable disc. By rotating the handle of the rotatable disc, the aperture of the base plate can be changed. To demonstrate a lens, a commercial ID is chosen, as shown in Fig. 2. It exhibits the same operating mechanism of the device as shown in Fig. 1. Fig. 2(a) shows the ID with a larger aperture, and Fig. 2(b) shows that with a smaller aperture. Its aperture can be tuned from 15 to 0.8 mm. The outer diameter of the ID is 22 mm. The ID has ten blades, and each blade is very thin. The surface of overlapped blades is rather smooth. The total thickness of the ID is 5 mm. The cross-sectional structure of the lens cell and its operation mechanism are shown in Fig. 3. From the bottom to the top, it consists of the glass plate, liquid-1 (L-1), liquid-2 (L-2), and the top glass plate. L-1 and L-2 are separated by the blades of the ID except meeting in its central opening area",
            " To demonstrate a liquid lens as shown in Fig. 3, we used BK-7 matching liquid (surface tension \u03b31 \u223c 40 mN/m, refractive index n1 \u223c 1.52, Abbe # \u223c57, Cargille) as L-1 and silicone oil (\u03b32 \u223c 21 mN/m, refractive index n2 \u223c 1.40, Abbe # \u223c87) as L-2. The two materials are highly clear and immiscible. They do not evaporate even at high temperature. Moreover, the color aberration of the lens can be suppressed, and the lens can present high optical performance owing to their large Abbe numbers. The ID as shown in Fig. 2 is chosen for preparing the lens. The fabrication process is shown in Fig. 4. A 0.7-mm-thick glass plate (bottom glass) was used to tightly seal the open area of the ID base plate, as shown in Fig. 4(a). The aperture of the diaphragm was set at 2r \u223c 4 mm, and liquid BK-7 was injected to the bottom chamber. Because the surface tension of the metal blades is fairly small (\u223c20 mN/m), excessive BK-7 will form a convex shape and have a contact angle on the border of the opening aperture, as shown in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000388_14644193jmbd97-Figure8-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000388_14644193jmbd97-Figure8-1.png",
        "caption": "Fig. 8 Effect of rocking angles",
        "texts": [],
        "surrounding_texts": [
            "The curvature difference for a Hertzian contact is found to be a function of the elliptical parameters as follows (for more information see [34]) F\u00f0r\u00de \u00bc \u00f0k 2 \u00fe 1\u00de= 2@ \u00f0k2 1\u00de= \u00f01\u00de where k is the elliptical eccentricity parameters and @ and = are complete elliptic integrals of the first and second kind, respectively. By assuming values of the elliptical eccentricity parameter, a table of k versus F(r) can be obtained and then it can be shown that the local force and deflection relationship of two bodies may be written as follows [34] d \u00bc d 3 2 WP r 1 n21 E1 \u00fe 1 n22 E2 2=3P r 2 \u00f02\u00de where d* is dimensionless deflection and is expressed as d \u00bc 2@ p p 2k2= 1=3 \u00f03\u00de d* is given as a function of F(r) [15]."
        ]
    },
    {
        "image_filename": "designv11_12_0002187_gt2010-22086-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002187_gt2010-22086-Figure1-1.png",
        "caption": "Figure 1- Cross section view of simple radial foil bearing.",
        "texts": [
            " In general, foil bearings are capable of supporting lightly loaded, high-speed rotating shafts. The lubricating fluid-film is generated by the viscous pumping action of the moving shaft or runner surface. The fluid film forms between the moving surface and a thin, flexible sheet metal foil layer that is supported by a series of spring foils. The foil layer facing the moving surface traps the hydrodynamic gas film and the supporting spring foils provide compliance, tolerance to misalignment and distortion and a host of other attributes such as damping [3]. Figure 1 shows a sketch of a typical bump style foil journal bearing. This material is declared a work of the U.S. Government and is not subject to copyright protection in the United States. Approved for public release. Distribution is unlimited. Downloaded From: http://proceedings.asmedigitalcollection.asme.org/pdfaccess.ashx?url=/data/conferences/gt2010/70392/ on 02/21/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use 2 There are two distinct types of foil gas bearings, journal and thrust bearings as depicted in Figure 2"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003082_0022-2569(70)90069-8-Figure3-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003082_0022-2569(70)90069-8-Figure3-1.png",
        "caption": "Figure 3.",
        "texts": [
            " X , Y must lie on the ellipse defined by (3.19), (3.20) and (3.21), and must also lie within the circle X -~ + yz = 1 if an output angle is to exist. We may summarize the possible types of intersection as follows: (a) The ellipse lies wholly within the circle. This means that the input crank rotates completely through 360 deg. The output link may, of course, rotate completely or rock. Also, there will be in general, two possible output angles for one input angle, depending upon the way the mechanism is initially put together. See Fig. 3 and equation (3.1). (b) The ellipse is wholly outside the circle. In this case, it is impossible to assemble the mechanism. (c) The ellipse lies outside of, but jus t touches tile circle. This means that the R - S - S-R linkage can only be assembled in one particular configuration and is a structure. (d) The ellipse intersects the circle in two points. Then the portion of the ellipse within the circle defines a range of possible input angles. (e) The ellipse intersects the circle in four points"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000168_iros.2006.282537-Figure4-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000168_iros.2006.282537-Figure4-1.png",
        "caption": "Fig. 4. Avoidance Manipulability Ellipsoids",
        "texts": [
            " In this paper, the workspace is assumed by two-dimension to make it comprehensive(m = 2). The avoidance manipulability ellipsoid of link2 is the complete avoidance manipulability ellipsoid because the singularvalues of 1M2 denoted by \u03c321 and \u03c322 are not zero, so the avoidance manipulability ellipsoid of link2 is expressed by ellipsoid. The avoidance manipulability ellipsoids of link1 and link3 are the partial avoidance manipulability ellipsoids because the one of singularvalues of 1M1 of them are zero, which are expressed by straight lines shown in Fig.4. However, the avoidance manipulability ellipsoid of manipulator\u2019s each link evaluates merely the mobility of each link and it is not enough to evaluate whole manipulator\u2019s shape. For keeping avoidance ability of whole manipulator\u2019s shape high, we propose an evaluation index of whole manipulator\u2019s shape. That is AMSI(Avoidance Manipulability Shape Index) and the evaluation index of whole manipulator\u2019s shape is defined by, 1E = n\u2211 i=1 1Viai (8) where, a1 = an\u22121 = 1[m\u22121], a2,3,\u00b7\u00b7\u00b7,(n\u22122) = 1[m\u22122] (9) 1V1, 1Vn\u22121[m] denote the length, 1V2,3,\u00b7\u00b7\u00b7,(n\u22122)[m2] denote area"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001855_s00216-010-3866-6-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001855_s00216-010-3866-6-Figure1-1.png",
        "caption": "Fig. 1 Axial symmetric computational domain (not to scale) for the simulation of liquid\u2013liquid interfaces confined within micropores. The rotational symmetry is indicated by the arrow around the axis of symmetry (Z-axis). L denotes the depth of the micropore and M the location of the interface. The incline of the pore wall is given by the angle \u03b1",
        "texts": [
            " The pore depth and the pore wall angle can be adjusted in the preparation of the solidstate micropore arrays, and suggestions for a design optimised for SV analysis are presented. A detailed description of the simulation of CV involving ion transfer across ITIES located within micropores was given recently [18]. Here, we provide a short summary. Transport is considered by diffusion only, with the time-dependent diffusion equation given in cylindrical coordinates with z and r being the space coordinates [19]. The space coordinates are scaled by the radius ra of the micropore opening, so that Z= z/ra and R=r/ra are the dimensionless coordinates (Fig. 1). The normalised radius and depth of the pore are then given by Ra=1 and L=l/ra. By this definition, the dimensionless depth L of the pore is equal to the pore aspect ratio \u03b8, and the parameter L is sufficient to characterise the geometric properties of a single cylindrical pore. Figure 1 illustrates the axial symmetric computational domain. The location of the liquid\u2013liquid interface within the pore is at z=m, and the quantity M=m/ra marks the dimensionless recess of the interface (for convenience, the quantities L and M are assigned to the absolute values of their Z-coordinate). For M=0, the pore is filled with the organic phase and the interface is coplanar with the upper side of the membrane. For M=L, the pore is filled with the aqueous phase, and the interface is coplanar with the lower membrane side (Fig. 1). The concentrations are normalised by C=c/cbulk, where cbulk is the bulk concentration of the analyte ions initially present in the aqueous phase. The time t (in seconds) is transformed to the dimensionless time by \u03c4= tDA/ra 2, where DA is the diffusion coefficient of the transferring species in the aqueous phase, which is greater than or equal to DB, the diffusion coefficient of the same species in the organic phase. A dimensionless scan rate is defined by \u03c3=ra 2Fv/(DART), where v is the potential scan rate (V s\u22121) and F, R, and T have their usual meaning",
            " In the simulations presented, different diffusion coefficients are assumed for the transferring species in the aqueous and organic phase. The ratio of the diffusion coefficient in the organic and aqueous phase, g \u00bc DB DA \u00f01\u00de is used as a parameter in the simulations. Currents are normalised by the limiting current ilim= 4ziFDAc bulkra. The dimensionless current is then given by G \u00bc p 2 Z1 0 @CA R; Z; t\u00f0 \u00de @Z Z\u00bcM RdR \u00f02\u00de The gradient in Eq. 2 is calculated along the line 0 < R \u2264 1 at Z = M (see Fig. 1). Computational details Simulations were performed using the finite element method programme package COMSOL (version 3.5a; COMSOL Ltd., Hertfordshire, UK). Free mesh parameters were used at locations of high-concentration gradients. These locations are the points of the orifices at both sides of the pore and along the boundary line where both phases are in contact. Here, the maximal sizes of the triangular elements were set to 0.0002 with a factor of 1.1 for element expansion. These mesh parameters have been tested previously and shown to deliver acceptable error limits (<1%) in the calculated current [18]. The simulations were performed to assess the impact of the variables tc, \u03b3, M, L, and the incline angle of the pore wall, as described by the angle \u03b1 (Fig. 1). For the cationic analyte, the initial and preconcentration potentials, \u0394w o \u03d5st and \u0394w o \u03d5pc, were set to 0.3 and 1.05 V, respectively, and the formal potential of transfer, \u0394w o \u03d5 00 i , was 0.58 V. These potential values were chosen according to the transfer potential of TEA+, which was used as a probe to characterize the transfer behaviour across microinterface arrays [18] and to investigate the DPSV performance of the microinterface arrays [16]. The scan rate for the linear sweep stripping was 10 mV s-1",
            " Since the shielding effect in both cases is identical, the steady-state current during the preconcentration time is in both cases Gss=0.16. However, for L=4, the stripping peak is |Gp|= 0.03, while with L=8, the peak is |Gp|=0.32. However, compared to |Gp|=2.21 for L=4, M=0, the negative effect of a recessed interface on the stripping peak is obvious. The pore wall angle The above simulations were all performed for cylindrical pores, i.e., for pores with perpendicular walls where the angle \u03b1 is 90\u00b0 (see Fig. 1 for the definition of \u03b1). However, the shape of the pore can be expected to influence diffusion to and from the interface, and hence, the effect of the wall angle was investigated. For \u03b1<90\u00b0, the pore has a conical shape with the pore opening on the aqueous side of the membrane being larger than the pore opening on the organic side, and vice versa for \u03b1>90\u00b0. For the simulation, the dimensionless radius of the pore on the aqueous side and the depth of the pore were kept constant at Ra=1 and L=4, respectively, and the incline angle \u03b1 was varied between 78"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001531_s11012-009-9232-0-Figure7-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001531_s11012-009-9232-0-Figure7-1.png",
        "caption": "Fig. 7 Unity vectors of O6O \u2032, O1O \u2032\u2032, O4O \u2032 and O5O \u2032\u2032 lines",
        "texts": [
            " They are therefore parallel. A second important property of this mechanism is that the intersecting point of even joints axes O \u2032 and odd joints axes O \u2032\u2032 formed a line O \u2032O \u2032\u2032 which is always perpendicular on the planes (O2,O4,O6) and (O1,O3,O5). In order to demonstrate this property we must know the coordinates of the points O \u2032 and O \u2032\u2032. These coordinates will be dependent on the input angle \u03b8 and the twist angle \u03b1. For this, we consider four points A, B , C and D necessary to define the rotation axes equations. In Fig. 7 the A, B , C and D coordinates are written in the mobile reference frame linked to the own joint, x = 0, y = 0 and z = \u22121. The coordinates of the points A,B,C and D in the fixed reference frame R6(O6x6y6z6) are given in the Appendix 3. With these coordinates, we write the equations of the lines O6O \u2032 and O4O \u2032, respectively O1O \u2032\u2032 and O5O \u2032\u2032. The coordinates of the points O \u2032 and O \u2032\u2032 satisfied the two first equations, respectively the two last ones. We finally obtain the coordinates for these two points in the fixed reference frame R6: For the point O \u2032: \u23a7\u23a8 \u23a9 xO \u2032 = 0 yO \u2032 = 0 zO \u2032 = a\u00b7(cos\u03d5+1) sin\u03d5\u00b7sin\u03b1 (22) For the point O \u2032\u2032: \u23a7\u23a8 \u23a9 xO \u2032\u2032 = \u2212a yO \u2032\u2032 = a\u00b7sin \u03b8 1\u2212cos \u03b8 zO \u2032\u2032 = \u2212 a\u00b7(cos \u03b8+1) sin \u03b8 \u00b7 cot\u03b1 (23) By taking account of (22) and (23) relations, the O \u2032O \u2032\u2032 equation can be written: x a = y \u2212 a\u00b7sin \u03b8 1\u2212cos \u03b8 = z \u2212 a\u00b7(cos\u03d5+1) sin\u03d5\u00b7sin\u03b1 a\u00b7(cos \u03b8+1) sin \u03b8 cot\u03b1 + a\u00b7(cos\u03d5+1) sin\u03d5\u00b7sin\u03b1 (24) The unity vector n of this line has the next components: \u23a7\u23aa\u23a8 \u23aa\u23a9 nx = 1 ny = \u2212 sin \u03b8 1\u2212cos \u03b8 nz = (cos \u03b8+1) sin \u03b8 cot\u03b1 + (cos\u03d5+1) sin\u03d5\u00b7sin\u03b1 (25) The unity vector N orthogonal on the plane (O2O4O6) has the components given by the relation (20)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001960_s12239-009-0053-x-Figure6-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001960_s12239-009-0053-x-Figure6-1.png",
        "caption": "Figure 6. Dynamic model of input and output shafts with torsional springs.",
        "texts": [
            " Consequently, the transmitted tangential force (Ft) may be calculated by ADAMS (with the option of a function input) in conjunction with k in equation (3), and it varies as gears rotate. Considering k, the influence of tooth bending deflections can be included in the equivalent model. 2.2. Torsional Stiffness of Shafts In the equivalent model, the torsional stiffness of transmission shafts is represented in ADAMS as several torsional springs with torsional rates at each shaft section: (4) where L is the solid shaft length between two gears, G is the shear modulus, and J is the polar area moment of inertia. Figure 6 represents the locations of the torsional springs on the equivalent model shafts. Another approach, the frequency-based model, has been developed. The natural frequencies and modes of manual transmission input and output shafts (shown in Figures 7 and 8) were calculated based on finite element analyses and were read by an ADAMS model in order to reflect the stiffness of shafts. It takes three to four hours for this model, which is usually three-dimensional, to read and handle all elemental stiffnesses from the finite element shaft meshes"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000506_j.engfracmech.2008.05.004-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000506_j.engfracmech.2008.05.004-Figure2-1.png",
        "caption": "Fig. 2. Mechanics of gear tooth contact, (a) at point of first contact, (b) at pitch point and (c) at last point of contact [9].",
        "texts": [
            " In this study, the Hertz surface pressure distribution on the shift profile of the spur gear was maintained at a constant level in the single and double tooth meshing by making a width modification \u2018\u2018F/b\u201d. With this modification, the pitting damage formed on the single tooth meshing was investigated by minimizing the instantaneous pressure changes on the single and double tooth meshing area. The experiments were conducted under low constant speed and two different torque values. During operation, the teeth of two gears pass through three stages of contact: 1. First phase: includes mating of the first contact point of the dedendum of the driving gear shown in Fig. 2a with the addition of the driven gear. Here, both rolling and sliding movements are present [8,9]. At the same time, the two teeth share the tooth load. The rolling direction is from root to tip (A\u2013E) for the driving gear. For the driven tooth, the rolling direction is from tip to root (E\u2013A). Sliding and rolling orientation for the driving gear at the tooth tip is in the same direction, although it is the opposite in the tooth root. The sliding orientation is always towards the outside direction from the mating line. In contrast with the previous situation, the sliding velocity is from root to tip and towards inside from the mating line. 2. Second phase: At full mesh, when the two teeth meet at their common or \u2018\u2018operating\u201d pitch-line, there is only a rolling motion, and no sliding. However, this stage produces the greatest tooth loading (Fig. 2b), meaning that just one tooth holds the full tooth load. 3. Third phase: Coming out of mesh, the two mating teeth also move with a sliding action, opposite to the initial contact stage (Fig. 2c). In this region, the two teeth share the tooth load [10]. The gear teeth roll and slide on one another. Rolling velocity is beneficial because it entrains lubricant between the contacting areas, increases oil film thickness, and reduces the severity of asperity contacts. Sliding velocity, on the other hand, generates heat from friction and increases asperity distress. Damage due to contact fatigue in gear teeth usually occurs in one of three areas: along the pitch-line, in the addendum (i.e. above the pitch-line), and in the dedendum (i"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002894_icelmach.2010.5607978-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002894_icelmach.2010.5607978-Figure1-1.png",
        "caption": "Fig. 1. Model of an IPMSM with definition of the used coordinates [41]",
        "texts": [
            " These kind of methods do not require a mathematical model of the drive, exhibit good noise rejection properties, can easily be extended and modified, can be robust to parameter variations and are Review of Position Estimation Methods for IPMSM Drives Without a Position Sensor Part I: Nonadaptive Methods O. Benjak, D. Gerling I XIX International Conference on Electrical Machines - ICEM 2010, Rome 978-1-4244-4175-4/10/$25.00 \u00a92010 IEEE computationally less intensive [2] Coordinates, used in this paper, are defined in Fig. 1, which shows the model of a IPMSM. The 3-phase voltages of an IPMSM are given by [50]: d u dt0 0u R is uu d v0 0 .u R is vv dt 0 0u R is ww d w dt \u03c8\u23a1 \u23a4 \u23a2 \u23a5 \u23a1 \u23a4 \u23a1 \u23a4 \u23a1 \u23a4 \u23a2 \u23a5 \u03c8\u23a2 \u23a5 \u23a2 \u23a5 \u23a2 \u23a5 \u23a2 \u23a5 = +\u23a2 \u23a5 \u23a2 \u23a5 \u23a2 \u23a5 \u23a2 \u23a5 \u23a2 \u23a5 \u23a2 \u23a5 \u23a2 \u23a5 \u23a2 \u23a5\u23a3 \u23a6 \u23a3 \u23a6\u23a3 \u23a6 \u03c8\u23a2 \u23a5 \u23a2 \u23a5\u23a3 \u23a6 (1) Where uu, uv, uw, iu, iv, iw, \u03c8u, \u03c8v and \u03c8w are the phase voltages, currents and flux linkages respectively and Rs the phase resistance. (1) can be transformed into the rotor reference dq-frame [50]: d d r q0u iR dtdd s . d0u iR qq s q r ddt \u23a1 \u23a4\u03c8 \u2212 \u03c8\u23a2 \u23a5\u03c9\u23a1 \u23a4 \u23a1 \u23a4\u23a1 \u23a4 \u23a2 \u23a5\u23a2 \u23a5 \u23a2 \u23a5= +\u23a2 \u23a5 \u23a2 \u23a5\u03c8\u23a2 \u23a5\u23a2 \u23a5 \u23a2 \u23a5\u23a3 \u23a6 \u23a3 \u23a6\u23a3 \u23a6 \u23a2 \u23a5+ \u03c8\u03c9 \u23a3 \u23a6 (2) Where ud, uq, id, iq, \u03c8d and \u03c8q are the dq-axis-voltages, currents and flux linkages, respectively"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001112_ejc.15.534-544-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001112_ejc.15.534-544-Figure2-1.png",
        "caption": "Fig. 2. Two inverted pendulum connected by a spring.",
        "texts": [
            " Remark 3: To guarantee the boundedness of the parameters in the presence of the approximation error, which is unavoidable, the proposed adaptive laws (27) is modified it by introducing a modification term as follows: _ i1 \u00bc 1b T i Pieiwi1\u00f0zi\u00de 1 i1 _ i2 \u00bc 2b T i Pieiwi2\u00f0xi\u00de 2 i2 _uir \u00bc uir b T i Piei uiruir _uicom \u00bc uicom bTi Piei uicomuicom _\u0302v0i \u00bc v\u03020 i bTi Piei v\u03020 i v\u03020i \u00f047\u00de 5. Simulation Results In this section, we apply the proposed decentralized fuzzy model reference adaptive controller to a two-inverted pendulum problem [14] in which the pendulums are connected by a spring as shown in Fig. 2. Each pendulum may be positioned by a torque input ui applied by a servomotor and its base. It is assumed that the angular position of pendulum and its angular rate are available and can be used as the controller inputs. The pendulum dynamics are described by the following nonlinear equations. where y1; y2 are the angular displacements of the pendulums from vertical position. m1 \u00bc 2 kg;m2 \u00bc 2:5 kg are the pendulum end masses, j1 \u00bc 0:5 kg; j2 \u00bc 0:62 kg are the moment of inertia, k \u00bc 100N=m is spring constant, r \u00bc 0:5m is the height of the pendulum, g \u00bc 9:81m=s2 shows the gravitational acceleration, l \u00bc 0:5m is the natural length of spring, 1; 2 \u00bc 25 are the control input gains and b \u00bc 0:4m presents distance between the pendulum hinges"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000597_978-3-540-73129-0-Figure6.14-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000597_978-3-540-73129-0-Figure6.14-1.png",
        "caption": "Fig. 6.14. Inverse piezo effect in polarized ceramics. Voltage V is applied in the direction of polarization P . a Longitudinal effect, b transversal effect (cE P stiffness of the piezo material for constant field strength E)",
        "texts": [
            " Piezoceramics, for instance, gradually loose their piezoelectric properties even at operating temperatures far below the Curie temperature (depending on the material 120 . . . 500 \u25e6C, for multilayer ceramics (see below) 80 . . . 220 \u25e6C). Under certain applications when the inverse operating voltage is applied, it may not exceed 20% of the rated voltage, or depolarization may occur. Piezoceramic elements are mainly available as plates or discs with a quadratic, circular or ring-shaped profile and a length from 0.3 upto several millimeters long, with or without metal electrodes. Most are designed to make use of the longitudinal effect (see Fig. 6.14a), which is due to the high d33 value, which is the strongest effect. When making use of the transversal effect the actuator stroke depends also on the dimensions of the material, whereby the influence of the quotient s/l on stiffness and elongation is oppositional (see Fig. 6.14b). Since the 1980s, multilayer ceramics have grown more important. The so-called green and several tens of micrometers thick ceramic foil is cut into pieces and then coated with an electrode paste, similar to multilayer capacitors. The pieces are then placed on top of each other, pressed and sintered. They form a kind of monolithic object that is used as a finished transducer or as a basis for producing stacks (see Fig. 6.15). Multilayer ceramics reach the maximum permissible field strength at a driving voltage of about 100V (low voltage actuators), and achieve therefore the same elongations as ordinary (so-called high voltage) piezoceramics do for a driving voltage in the kilovolt range",
            " In this case, the origin of the diagram does not move, but the maximally achievable elongation is reduced by the factor cP/(cP + cF) (see Fig. 6.17b). In the extreme case cF \u2192 \u221e (fixed clamp support of the transducer), the transducer achieves its maximal force, the socalled clamping force or blocking force which also follows from (6.2), if S = 0. Equations (6.1) and (6.2) show that an ideal piezoelectric transducer input can be considered as an electric capacitor with the capacitance C and its output as a mechanical spring with the stiffness cP. This is illustrated in Fig. 6.14b for the d33 transducer, but the description holds in principle for all piezo transducers. Since C is in reality always lossy and cP always has a mass, the amplitude response |v/F | (sensory operation) has an electrically determined lower cut-off frequency fc and a mechanical eigenfrequency f0. When operated as an actuator, the electrical input is a voltage, that is, C is constantly recharged, so that fc has no effect on the amplitude response |s/v|, as shown in Fig. 6.16b. In contrast to the stack design, the laminar design is based on the piezo constant d31 and the transversal effect. The greater the quotient s/l of the piezoelectric element (see Fig. 6.14b), the bigger the effect. This leads to strip shaped elements with low stiffness. Therefore, several layers of strips are piled up, similar to the stack design, and form a so-called laminate for improving the mechanical stability. Since the transversal effect is applied, the result are flat transducers which shorten proportionally to the applied voltage, as d31 is negative. Bending elements feature the transversal effect as well. They can consist, for instance, of a PZT ceramic mounted onto a piece of spring metal (monomorph)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002449_ijcat.2010.032200-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002449_ijcat.2010.032200-Figure2-1.png",
        "caption": "Figure 2 The illustration of the data acquiring system",
        "texts": [
            " These devices scan without contacting the profile of a physical model by a striped laser beam and CCD cameras capture profile images from which digital data (point cloud data) can be generated by triangulation algorithms. Using this method, hundreds of surface points can be obtained per second. It will take only a few minutes to digitise the worn metal part, no matter how complex its surface geometry is. In this research, a three-dimensional digitiser based on stereo-view has been established by Shenyang Institute of Automation, Chinese Academy of Sciences. Figure 2 shows the illustration of the data acquiring system. For point cloud generated by a three-dimensional digitiser, two-stage modelling is needed to reconstruct the surface (Lai et al., 2001). In the first stage, a topologically triangular mesh is calculated from the point cloud; then, the NURBS surface model follows in the second stage. Before calculating the triangular mesh, data extraction is used to reduce the amount of the data points. Efficient machining of surfaces is very important in many manufacturing industries, such as turbine blades, shoe lasts and dies (Park, 2004)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000597_978-3-540-73129-0-Figure6.16-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000597_978-3-540-73129-0-Figure6.16-1.png",
        "caption": "Fig. 6.16. Piezoelectric stack translator. a Structure, b electromechanical equivalent circuit and amplitude responses of the actuator and sensor transfer behaviour in small signal operation (derived from [5])",
        "texts": [
            " Furthermore, it lends itself to explaining the construction and properties of piezoelectric actuators. Structure. The active part of the transducer consists, for instance, of many 0.3 to 1 mm thin ceramic discs that are mounted with metal electrodes, e. g. made of nickel or copper, for applying the operating voltage. The discs are stacked up in pairs of opposing polarization and glued together. Highly insulating materials seal the stack against external electrical influences. In other designs \u2013 the so-called low-voltage actuators \u2013 the multilayer ceramics described above are used. Figure 6.16 features the electric parallel connection and the mechanical series connection of the stack. Its displacement is the sum of the single element elongations \u0394l. The applied field and the achieved elongation are in line with the polarization, that is, the piezo constant d33 is used (longitudinal effect). The transducer can also handle tensile forces, if prestressed with a slotted cylinder casing as shown in Fig. 6.16 or with an anti-fatigue bolt, as is commonly done. Static and Dynamic Behaviour. The static diagram S(E) in Fig. 6.17 holds for no-load operation (T = 0 in (6.2)). The addend sET in (6.2) takes into account the loaded piezo transducers elastic deformation. Two cases are distinguished: \u2013 The load is constant, e. g. weight FG. In this case, the entire diagram is shifted by sET = \u2212FG/c E P . (6.4) The spring constant cEP follows from the (6.2), if E = 0 (see Fig. 6.17). As long as the maximum permissible load is not exceeded, the original no-load expansion of the piezo substance holds (see Fig",
            " This is illustrated in Fig. 6.14b for the d33 transducer, but the description holds in principle for all piezo transducers. Since C is in reality always lossy and cP always has a mass, the amplitude response |v/F | (sensory operation) has an electrically determined lower cut-off frequency fc and a mechanical eigenfrequency f0. When operated as an actuator, the electrical input is a voltage, that is, C is constantly recharged, so that fc has no effect on the amplitude response |s/v|, as shown in Fig. 6.16b. In contrast to the stack design, the laminar design is based on the piezo constant d31 and the transversal effect. The greater the quotient s/l of the piezoelectric element (see Fig. 6.14b), the bigger the effect. This leads to strip shaped elements with low stiffness. Therefore, several layers of strips are piled up, similar to the stack design, and form a so-called laminate for improving the mechanical stability. Since the transversal effect is applied, the result are flat transducers which shorten proportionally to the applied voltage, as d31 is negative"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002570_016918610x552187-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002570_016918610x552187-Figure2-1.png",
        "caption": "Figure 2. Relationship between object and finger coordinate frames.",
        "texts": [
            " \u03b5 = [xT, \u03beT]T \u2208 R6: Configuration (position and orientation) displacement of the object, where x = [x, y, z]T \u2208 R3 is translation and \u03be = [\u03be, \u03b7, \u03b6 ]T \u2208 R3 is rotation. \u03b5f \u2208 R6: Configuration displacement of the finger. \u03b1o \u2208 R2: Displacement parameter of the contact point on the object surface. \u03b1f \u2208 R2: Displacement parameter of the contact point on the finger surface. \u03c6 \u2208 R: Relative orientation between two objects, where the relationship between their two coordinate systems Co and Cf is defined as in Fig. 2. aTb \u2208 R4\u00d74: Homogeneous transformation matrix representing the configuration of b in a, as defined by: aTb := [ aRb apb 01\u00d73 1 ] , (1) where the symbols apb and aRb are a position vector and a rotation matrix, respectively. Other homogeneous matrices are similarly defined. To derive the contact frame velocity on a rigid body surface, we consider the contact surface geometry (metric tensor MC, curvature KC and torsion TC) [22, 23]. The configuration of the contact coordinate frame C in the local coordinate frame L is represented by a contact displacement parameter \u03b1"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002342_j.mechmachtheory.2011.06.006-Figure7-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002342_j.mechmachtheory.2011.06.006-Figure7-1.png",
        "caption": "Fig. 7. An equivalent multibody model of a gymnast on a trampoline.",
        "texts": [
            " Examples 6.1. A gymnast on a trampoline Let us consider a gymnast on a trampoline as a first example of the application of the algorithm. This example has been considered in [17,39]. In [17,39], the gymnast was modelled as a planar multibody system depicted in Fig. 6. See [39] for the assumptions on which this multibody model is based. In order to illustrate the proposed method for the determination of joint reaction forces, the multibody system shown in Fig. 6 is replaced by a multibody system shown in Fig. 7. In Fig. 7, body (V0) is a fixed base body and bodies (V1) and (V2) represent fictitious bodies. Joints 1, 2, 3, and 6 are unactuated, while the driving torques in the remaining joints are given by \u2212\u03c42 e\u21924, \u03c43 e\u21925, \u03c46 e\u21927, \u03c44 e\u21928, and \u03c45 e\u21929, where \u03c4i(i=1, \u2026, 6) represent the torques of muscle forces in the gymnast's joints. Note that bodies (V6) and (V3) act to each other by the couples with torques M \u2192 3 = \u03c41 i \u2192 and M \u2192 2 = \u2212\u03c41 i \u2192 , respectively, and that the axes of joints 3 and 6 coincide. It is taken that the trampoline acts to body (V5) by a force system which is equivalent to the force 5 N \u2192 = 0; N\u03b75 ; N\u03b65 T applied at the mass centre C5 and a couple with torque 5M \u2192 = M\u03be5 ;0;0 T equals to the resultant moment vector of the trampoline force system about point C5"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000410_0094-114x(80)90020-8-Figure3-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000410_0094-114x(80)90020-8-Figure3-1.png",
        "caption": "Figure 3.",
        "texts": [
            " The purpose of what follows is to demonstrate an application of the theory to very important and barely investigated phenomena. No attempt is made to solve in detail specific, dimensioned cases. Such is here regarded as belonging more to the realm of numerical analysis than algebraic kinematics. The 4-bar linkage which we have selected is conveniently replaced by a 6-bar C-R-RR-R-R- chain which has the following (actual or optional) dimensional constraints. a 1 2 = R 2 = ~ / 2 ~ : R 3 = f l s 4 : R 4 : 0 ~ 4 5 : 0 0L23 : t1'34 = ~- 2 It is illustrated schematically in Fig. 3, and particular attention is drawn to the location of the fundamental frame of reference. The present technique takes great advantage from a judicious choice of the fundamental frame. We may write down the various screw motor elements as follows. to] = 0 ) j k /~t = a120)~j oJv = 0 ~ur = ~vk to2 = 0)2k P2 = 0 = 0) /so= \\ = o / / 0) 5 ---- 0) 5 (co,o3) , c 2co3co,+so2so,) sO2s03 ~ s02c03c04- cOzs04 - c03 \\ s03c04 - s02c04 + c02c03s04 '\\ ~'\u00a35 = 0450)5 C02C04\"Jr\" SO2cO3sO 4 ) sO3s04 / o)6 = 0)6 / sOisa61 \\ 1~6 = \u00b076 i ~ CO1Sa61 ) \\ C~61 d61sOlcct61 + rtcOlsa61 ) - r~sO~s~6~ + a6~cO~ca6~ + a l 2 C O t 6 1 - d 6 1 S O t 6 1 - - dl2COlsa61 Stationary positions of the two revolutes are obviously of significance, since one of them would normally be chosen as the input pair and the other as the output pair",
            " L'une d'elles [1,2] est basle sur l'alg~bre des syst~mes-vis [4-8], et on l'a utilis~e [3] aussi pour d~terminer les extrema pour le mouvement absolu d'un point arbitraire dans une boucle spatiale. Dans cet article, on utilise la th~orie pour d~terminer une configuration d'incertitude [10,11] d'un ra~canisme spatial. Un caract~re particulier de l'alg~bre est l'usage qu'on fair des ~quations de clSture [2,7,9] de la cha~ne. L'exemple choisi pour d@montrer 1'application de l'alg~bre des syst~mes-vis est le m~canisme R-S-C-R-, qu'on peut remplacer corm~od~ment par une boucle C-R-R-R-R-R-, illustr~e dans la Figure 3. Nous consid~rons trois configurations sp~ciales de ce m~canisme, indiquant par i~, leurs connexit~s. Ce sont les positions d'arr~t de chacun des deux couples tournants et la configuration d'incertitude de la boucle. On peut comparer directement les r~sultats avec le travail de Hunt [10,113, qui utilise une approche purement g~om~trique. La condition d'arr@t de 1'articulation 5, consid~r~e dans la subdivision 2a, demande la specification des variables des articulations. Un r~sultat semblable s'applique aussi pour une position stationnaire du couple 6, trait~ en 2b"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000178_isic.2007.4450908-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000178_isic.2007.4450908-Figure2-1.png",
        "caption": "Fig. 2. Left: COG motion in the sagittal plane during single support phase, Right: COG motion in the sagittal plane during double support phase. During single support phase, the COG travels from (\u2212Lxs, H) to (Lxs, H), ZMP travels from (Lxs/axs,0) to (Lxs/axs, 0). The zero point of the coordinate is located at the center of the support. During double support phase, the COG travels from (\u2212Lyd,H) to (Lyd, H), the ZMP travels from (\u2212Lxd/axd, 0) to (Lxd/axd, 0). The zero point of the coordinate is located at the center of the two feet when the humanoid robot stands in double support phase.",
        "texts": [
            " To use IPM as a model of humanoid robot, the following ideal assumptions are made: 1) A concentrated mass of robot resides at hip 2) Robot motions in sagittal and frontal planes are inde- pendent 3) The COG Position of IPM is constant in z-axis 4) The position of ZMP is linear to COG The relationship between ZMP, COG and reaction force is depicted in Fig.1. The position of COG is (X, H), the position of ZMP is (aX+b), and the normal vector of ground force is parallel to the vector (cX + d,H). The dynamics of the IPM can be given by Fx : Fz = x\u0308 : (z\u0308 + g) = (cx + d) : H (1) Because z = H , we can get x\u0308 = g(cx + d)/H (2) A. IPM for Humanoid Motion in the Sagittal Plane One step of humanoid walking consists of two phases, single support and double support phase. The COG motion during these two phases in the sagittal plane is shown in Fig.2. The IPM motion in the frontal plane can be obtained in a similar way. The dynamics equation is given by      x\u0308gs : ( \u00a8zgs + g) = xgs(cxs \u2212 1 axs ) : H x\u0308gd : ( \u00a8zgd + g) = xgd(cxd \u2212 1 axd ) : H (3) where xgs is the COG position of IPM in the sagittal plane during the single support phase, xgd is the COG position of IPM motion in the sagittal plane during the double support phase. Solving the above equations, we get the COG motion as follows      xgs(t) = Cxs1cosh t sxs + Cxs2sinh t sxs xgd(t) = Cxd1cos t sxd + Cxd2sin t sxd (4) With boundary conditions xgd(0) = \u2212Lxs, xgs(Ts) = Lxs, xgd(0) = \u2212Lxd, xgd(Td) = Lxd"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001704_acc.2009.5160384-Figure5-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001704_acc.2009.5160384-Figure5-1.png",
        "caption": "Fig. 5. An illustration for the proof of Theorem 4.",
        "texts": [
            " In which case there does not exist any common \u03b8 such that the X or Y axis distance will be reduced. They will remain at the same position after executing the LP. Theorem 4: If the solution of the LP problem yields a square formation then the number of agents on the boundary of the square is more than two. Proof: We will prove this by contradiction.Let, the number of agent on the square are two and they are located at diagonally opposite corners. The other agents are the interior of the square. For the sake of simplicity we will consider only three agents. This is given in Fig. 5,where MN is the Voronoi edge between agents located at A and B. According to Theorem 1, there always exists (\u03b8, d) such that they can meet at an unique point on the Voronoi edge (MN ). Let the unique meeting point be F . Let us define a very small positive quantity \u2206d such that \u2206d < min 1 4 {\u03b71, \u03b72, \u03b73, \u03b74}. After broadcasting (\u03b8, \u2206d) the agents will move from their positions. The new position of agents A and B are C and D, respectively. The interior agent will remain in the interior of the square"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003761_9781118818428-Figure15.1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003761_9781118818428-Figure15.1-1.png",
        "caption": "Figure 15.1 represents one of a total of nTx ducts in parallel, each of length Lx [m], of internal diameter dx [m] and at constant, uniform temperature Tw. Gas enters from the left at x = 0 with temperature Ti. Computation proceeds in terms of temperature difference, or \u2018error\u2019, \u03b5T, defined as T \u2212 Tw, of which the value at inlet, \u03b5Ti, is Ti \u2212 Tw.",
        "texts": [
            " For expansion and compression exchangers respectively: Q\u2032 xe = Pwr\u2215\u03b7bth [W] (15.3a) Q\u2032 xc = Pwr (1 \u2212 \u03b7bth)\u2215\u03b7bth [W] (15.3b) Less steam, more traction \u2013 Stirling engine design without the hot air 151 The expansion exchanger provides the example: Equation 15.3a is substituted into Equa- tion 15.2, giving: \u03b5Ti = (Pwr\u2215\u03b7bth)(1 \u2212 e\u2212NTU)\u22121\u2215m\u2032cp [K] (15.4) Temperature \u2018error\u2019 \u03b5T may now be expressed as a function of x\u2215Lx [i.e., as \u03b5T(x\u2215Lx)] on substituting Equation 15.4 into Equation 15.1. Over the elemental length of duct dx (Figure 15.1), rate of loss of available work, T0ds \u2032 gen may be evaluated (Bejan 1994), q\u2032T0\u03b5T\u2215T2: T0ds \u2032 gen = q\u2032T0\u03b5T\u2215T2 [W] (15.5) In Equation 15.5 q\u2032 [W] is heat rate by convection over channel length dx, T is local temperature and T0 is ambient temperature \u2013 the lowest temperature at which q\u2032 could realistically be rejected. For present purposes T0 can be replaced by TC. In terms of \u03b5T (now known), q\u2032 = \u2212h \u22c5 \u03c0dx\u03b5Tdx, in which h [W/m2K] is coefficient of convective heat transfer. Dividing the right-hand side by \u03c1|u|cp (and multiplying by m\u2032cp\u2215Aff to compensate) exposes Stanton number, St = h\u2215\u03c1|u|cp: q\u2032 = \u2212\u03b5T m\u2032cp Stdx\u2215rh (15"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001903_physreve.79.011705-Figure15-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001903_physreve.79.011705-Figure15-1.png",
        "caption": "FIG. 15. Schematic illustration to explain why 0 0 max 90\u00b0.",
        "texts": [
            " From the discussion above, spontaneous twist deformations in the bulk exist when KA2q3 K2 /Lz. Since we have chosen Lz=2 /q and K2 K as noted above, this condition leads to 2 qA 2 1. B1 Equation B1 is satisfied when qA=3 /16 0.6, while it is not when qA= /16 0.2 or qA= /160 0.02. This argument is consistent with the presence of twist deformation in the bulk for qA=3 /16 and its absence for qA= /16 or /160 in the case of =90\u00b0. Next we show why 0 cannot exceed some value 0 max 90\u00b0 in the case of qA=3 /16 as found in Fig. 6. We again rely on a similar schematic illustration in Fig. 15, showing how 0 is determined with the variation of . As in the above argument, 0 is determined by the intersection of the curve representing f\u0303a ( 0 ) with a straight line with a slope \u2212K2 /Lz. When is larger than some value max indicated in Fig. 15, the line does not intersect with the curve f\u0303a ( 0 ), indicating that no solution exists, or that 0 cannot be determined. When = max, 0 is evidently smaller than 90\u00b0. We note that this behavior is observed only when f\u0303a ( 0 =90\u00b0 ) K2 /Lz, which is the case for qA=3 /16 as mentioned above. Otherwise, the line with a slope \u2212K2 /Lz always intersects the curve f\u0303a ( 0 ) and therefore 0 can reach 90\u00b0 when =90\u00b0. APPENDIX C: EXPLICIT FORM OF ny max(z , (0)) In this appendix we will give the explicit form of the maximum azimuthal distortions ny max(z , 0 ) necessary for plotting the solid lines in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003680_physreve.88.012701-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003680_physreve.88.012701-Figure1-1.png",
        "caption": "FIG. 1. The present numerical models: (a) equivalent 2D elasticfoundation model; (b) randomly distributed 3D cross linkers model.",
        "texts": [
            " Based on our numerical simulations, two empirical relations are proposed for the critical force and wavelength of the localized buckling of microtubules. As will be shown in the paper, the localized buckling of microtubules, predicted by the present model, is in reasonable agreement with some recent experimental data. 012701-11539-3755/2013/88(1)/012701(8) \u00a92013 American Physical Society The microtubule cross linker system, studied in the present paper, is composed of a microtubule surrounded by cross linkers randomly distributed along the longitudinal direction, see Fig. 1. The microtubule is modeled as an elastic hollow cylinder column as presented in Sec. II A, whereas, the modeling of cross linkers is illustrated in Secs. II B1 and II B2. The two key parameters of the cross linkers are the spring constant k and the spacing of cross linkers Ld , whose values will be given based on available experimental data. Buckling of microtubules is stimulated by imposing compressive axial displacement at one end of the microtubule with the other end fixed. The microtubule is modeled as a hollow cylinder column with an outer diameter of 25 nm and a thickness of 1",
            " For the 2D in-plane buckling of a microtubule supported by a continuum elastic foundation, the wavelength and critical buckling force are given by [4] \u03bb = 2\u03c0 ( EI Ec )1/4 , (1) and Fcr = 2 \u221a EcEI = 8\u03c02 EI \u03bb2 , (2) where EI is the bending rigidity of the microtubule and Ec is the elastic-foundation modulus. The elastic-foundation modulus Ec in the 2D condition is directly proportional to the spring constant k and is inversely proportional to the spacing Ld of the uniformly distributed linear springs [5] as Ec = k Ld . (3) This formula (3) is commonly used provided that Ld is much smaller than the wavelength. To compare the present numerical model with the elasticfoundation model, a 2D numerical model is shown in Fig. 1(a) where all cross linkers are aligned on the same plane and are perpendicular to the microtubule. All out-of-plane displacements and rotations are not allowed. In addition, as assumed in the elastic-foundation model, all cross linkers in the equivalent elastic-foundation model shown in Fig. 1(a) are capable of bearing both compressive and tensile forces [12,13]. 2. Randomly distributed 3D cross linker model The major goal of the present study is buckling behaviors of a microtubule surrounded by 3D randomly distributed discrete cross linkers. The morphological details of the cross linkers are modeled based on experimental observations. All cross linkers are attached to the microtubule in random directions assigned by the uniform distribution rule, see Fig. 1(b). The cross linkers are modeled as linear springs with negligible bending rigidity [9,21], thus, they can bear axial tension but are vulnerable to any axial compression. To realize this, the load increments are sufficiently small, and every spring element is verified after each load increment and will be permanently removed if the axial force becomes compressive (actually, if a small nonzero threshold value of compressive force, such as 1 to 2 pN, is 012701-2 used, our results remain essentially unchanged"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002548_j.cnsns.2010.06.004-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002548_j.cnsns.2010.06.004-Figure1-1.png",
        "caption": "Fig. 1a. Rotor supported on fluid-film bearings.",
        "texts": [
            " The governing equation for the bearing is modified Reynolds equation without considering the slip and Darcy\u2019s equation. Bifurcation diagrams are presented to observe the essential features of nonlinear dynamics of the rotor-bearing system. Poincar\u00e9 map, journal trajectory, time response, FFT-spectrum, etc. are also employed to study the complex dynamics of the system as rotor spin speed is increased. A uniform shaft of length L mounted with a disk of mass md, rotating at constant speed X about the Z-axis and simply supported on two bearings is shown in Fig. 1. The shaft is discretized into a finite number of elements. The element is considered to be initially straight and is modeled as eight-degrees-of-freedom element, two translations and two rotations at each node. Axial motion is assumed to be negligible. The cross-section of the element is circular and is considered to be uniform. A typical cross-section of the shaft, in a deformed state, located at a distance z from the left end can be described by translations in X- and Y-directions as well as the small rotations about X- and Y-directions",
            " (4), the right-hand side is the force vector consisting of hydrodynamic forces at the bearing ends and unbalance forces due to the disk eccentricity. The non-dimensional forces at the bearing locations can be expressed as, FY \u00bc FB /pS sin / FB r pS cos / 1 FX \u00bc FB /pS cos / FB r pS sin / \u00f05\u00de where S is the Sommerfeld number. The non-dimensional unbalance forces at the disk location are given by, Fd Y \u00bc 2 CX2 g ! ed C md 1\u00fe md sin t Fd X \u00bc 2 CX2 g ! ed C md 1\u00fe md cos t \u00f06\u00de where md is the ratio of disk mass to shaft mass. Fig. 1 shows the schematic diagram of the porous journal bearing. The lubricant flow through the porous bush and the clearance space is viscous and laminar. The flow through the porous bush is governed by Darcy\u2019s law and that through the clearance space is governed by modified Reynolds equation. The differential equations for the porous hydrodynamic journal bearing [17] are given below For the porous media Kx o2 p0 oh2 \u00fe R H 2 o2 p0 o y2 \u00fe Kz D LB 2 o2 p0 o z2 \u00bc 0 \u00f07\u00de For the clearance space o oh h3 o p oh \u00fe D LB 2 o o z h3 o p o z \u00bc 6 1 2 _/ \u00f0 e sin h\u00de \u00fe 12 _e cos h\u00fe b o p0 o y y\u00bc0 \u00f08\u00de The non-dimensional bearing forces along the radial and the circumferential directions are given by FB r \u00bc Z 1 0 Z h2 0 p cos hdhd z FB / \u00bc Z 1 0 Z h2 0 p sin hdhd z \u00f09\u00de The cavitation model adopted in the present analysis is based on Reynolds boundary condition"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002930_978-3-642-31988-4_26-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002930_978-3-642-31988-4_26-Figure2-1.png",
        "caption": "Fig. 2 Precise catenary modeling of a cable",
        "texts": [
            " For building the FAST, a six-cable driven parallel manipulator is adopted as the firstlevel adjustable feed support system, and a similarity model of the six-cable driven parallel manipulator is set up in Beijing. This section will discuss the modeling method of the six-cable driven parallel manipulator. Considering cable mass and elastic deformation, precise catenary equation and simplified catenary equations are set up in this section. For setting up a cabl e model with precise catenary equation, the symbols used in Fig. 2 are defined as: l0 is the unstrained length of the cables; \u0394l the strain of the cable; T the tension applied to the fixed end of the cable; \u03c1 the unstrained linear density; E the elastic modulus; A0 the unstrained cross-sectional area; H the horizontal component of the cable tension vector and V its vertical component. Using the variables and coordinate system above, we will briefly reproduce Irvines derivation [8] in this paper. As shown in Fig. 2 a point along the length of the strained cable can be denoted by Cartesian coordinate and. To begin with, the cable must satisfy the geometric constraint: \u2211 x = 0 H + d H \u2212 H = 0 (1) \u2211 z = 0 H dz dx + d ( H dz dx ) + \u03c1gdl0 \u2212 H dz dx = 0 (2) where, dl dl0 = T E A0 + 1 (3) T = H \u221a 1 + ( dz dx )2 (4) From dl = dx \u221a 1 + ( dz dx )2 , Eq. 2 can be expressed as: d ( H dz dx ) + \u03c1g E A0 T + E A0 dl = 0 (5) Assuming dz dx = p, Eq. 5 can be written as: dp dx + \u03c1gE A0 H \u221a 1 + p2 H \u221a 1 + p2 + E A0 = 0 (6) Therefore, x = \u2212 H \u03c1g sh\u22121 ( dz dx ) \u2212 H2 \u03c1gE A0 dz dx + c (7) Where, sh\u22121(x) = ln ( x + \u221a 1 + x2 ) , x \u2208 (\u2212\u221e,+\u221e) (8) x = \u2212 H \u03c1g ln \u239b \u239d dz dx + \u221a 1 + ( dz dx )2 \u239e \u23a0 \u2212 H2 \u03c1gE A0 dz dx + c (9) Integrating and applying the boundary conditions as follows: x = 0, z = 0 x = L , z = h (10) The length of cable is l, the unstrained length of the cable is l0, and \u0394l represents the strain of the cable"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001972_s11071-010-9801-8-Figure3-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001972_s11071-010-9801-8-Figure3-1.png",
        "caption": "Fig. 3 Characteristic of the logical gain \u03ba(zn\u22121) with respect to |zn\u22121|",
        "texts": [
            " Suppose that the initial values of the auxiliary system are all zero and then it can be reformulated into a standard 2nd order system with a feedback controller depicted in Fig. 2. Regarding this standard control architecture, the controller \u03ba(zn\u22121), which is a function of zn\u22121, represents the action of \u03be sgn(\u2022). Note that \u03ba(zn\u22121) can be treated as a selfmanipulated logical gain. It possesses a nature that the magnitude varies with respect to the size of zn\u22121 automatically. In more explicit words, the state dependent gain \u03ba(zn\u22121) has the feature that \u03ba(zn\u22121) \u00b7zn\u22121 = \u03be sgn(zn\u22121). The property of \u03ba(zn\u22121) is depicted in Fig. 3, where the rectangular areas are always equivalent, i.e., Aabcd = Aa\u2032b\u2032c\u2032d = \u03ba \u00b7 |zn\u22121| = \u03be . The representation of the nonlinear component is analogous to the analysis technique adopted to forecast whether a limit cycle occurs when a system is subjected to input nonlinearities [15]. In the following, the stability and convergence feature of (13) is addressed in the sense of high gain control. First, according to Fig. 2, it is confident that the system is stable by choosing kn > 0. Once kn has been determined, the corresponding roots loci is drawn Fig. 4. In addition refer to (16), owing to \u03bc > 0, the control force always counteracts the direction of system output zn\u22121. In other words, the state zn\u22121 will be forced toward zn\u22121 = 0. Meanwhile, consider Fig. 3, owing to the nature of the logical control gain \u03ba(zn\u22121), it reveals that the smaller the |zn\u22121|, the larger the \u03ba(zn\u22121) such that the undesirable effects caused by external disturbance can be attenuated as small as possible by means of high gain control. Another alternative to analyze the type1 system, i.e., (16), subject to a relay control input in frequency domain is the use of describing function method [16]. It can be shown that the intersection of the nonlinear control component, i.e., the relay control, and the type-1 system is at origin"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002448_j.proeng.2012.07.040-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002448_j.proeng.2012.07.040-Figure2-1.png",
        "caption": "Fig. 2. Schematic diagram of the heat transfer",
        "texts": [
            " 1), in which the surfaces of the flexibly mounted stator and the surfaces of the rotor The fourth kind boundary condition is satisfied on the end faces of the rings - the rotor and the stator - being in contact with the fluid in the radial clearance gap. This is the case of heat transfer by conduction, where the heat flux at the interface between the rings and fluid film has the same value as that flowing into the seal elements. for 1 rs 0f f r f r z h z T T k k z z= = \u2202 \u2202= \u2202 \u2202 and f rT T= , (4) for 1 ss 0f f s f s z h z T T k k z z= = \u2202 \u2202= \u2202 \u2202 and f sT T= . The heat transfer between the surfaces of the rings and (Fig.1) and the surrounding fluid occurs through natural convection (Fig. 2), both for the stator and the rotor and can generally be written as: ( ) o o w r r r r T k H T T r \u221e= = \u2202\u2212 = \u2212 \u2202 , (5) where: , ,s f rk k k \u2013 heat conduction coefficients for the stator, the fluid film and the rotor, respectively, Hw \u2013 convective heat transfer coefficient, \u2013 surrounding fluid temperature, Tr \u2013 rotor temperature. For the energy equation, the boundary condition was the nonlinear temperature distribution at the inlet to the radial clearance gap approximated with a parabolic function [14, 16]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000115_j.jmatprotec.2007.09.010-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000115_j.jmatprotec.2007.09.010-Figure1-1.png",
        "caption": "Fig. 1 \u2013 Schematic diagram of cold rolling.",
        "texts": [
            " Based on the results of this study, empirical relations for prediction of dimensionless minimum film thickness and dimensionless maximum film temperature rise in inlet zone have been developed. Significant reduction in minimum film thickness (thermal) has been observed with the increase in the cold rolling speed and slip. 2. Governing equations The coupled solution of governing equations (Reynolds and energy equations) has been achieved for the thermohydrodynamic lubrication (in the inlet zone) of high-speed cold strip rolling. Inlet zone has been indicated in a schematic diagram of the strip rolling shown in Fig. 1. Present analysis treats the bounding solids of inlet zone as rigid. The numerical solution of energy equation requires very fine meshing of computational domain for acceptable accuracy of results. On the contrary, fine meshing consumes more CPU time and precious memory space due to involved computation. In order to enhance computational simplicity as well as savings in computational costs, in this work temperature profile across the film thickness has been approximated by Legendre polynomial of order 2",
            " (1) and (18), the following expression has been evolved: m\u0307 = (ur + us) h 2 \u2212 h 3 \u03041A \u2212 h 3 ( \u03040 + 2 5 \u03042 ) B (19) e c h n o l o g y 2 0 0 ( 2 0 0 8 ) 238\u2013249 where A and B have been defined earlier and given below to Eq. (1). Divergence of mass flux (from Eq. (19)) leads to the generalized Reynolds equation, which is as follows (Elrod and Brewe, 1986): \u2207 \u00b7 m\u0307 = 0 or \u2207 \u00b7 \u0304ph3 \u2207p = 6 ur \u2207h \u2212 2 \u2207 \u00b7 \u03041 \u03040 h (ur \u2212 us) (20) where \u0304p = \u03040 + 0.4 \u03042 \u2212 \u03042 1 3\u03040 3. Computational procedure The solution of the present thermo-hydrodynamic line contact problem (inlet zone), as shown in Fig. 1, starts with the known isothermal pressure distribution within the inlet zone as obtained from the converged solution of the Reynolds equation given as \u2202 \u2202x [ h3 \u2202p \u2202x ] = 6 0(ur + us) \u2202h \u2202x (21) where h is the lubricant film thickness in the inlet zone and is given by: h = h0 + (R2 \u2212 b2) 1/2 + (R2 \u2212 x2) 1/2 (22) A von-misses yield criterion is applied with the assumption that forward and backward tension is absent. Therefore Eq. (21) is subjected to boundary conditions as shown below (i) p = y at x = b (ii) p = 0 at x = \u221e While solving Eq"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001370_icsens.2009.5398292-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001370_icsens.2009.5398292-Figure2-1.png",
        "caption": "Fig. 2. Flat flexible displacement sensor",
        "texts": [
            " (4) It is easy to mount multiple sensors in the longitude direction of the pneumatic actuator to obtain the average value for reducing the measuring error. The flexible displacement sensor we developed can be installed in MPA directly without any rigid frames to mount conventional displacement sensors. The sensor is made from electro-conductive rubber, which is flexible and lightweight just like the MPA. Therefore, using this flexibile sensor, the flexibility and lightweight properties of MPA will not lost. 978-1-4244-5335-1/09/$26.00 \u00a92009 IEEE 520 IEEE SENSORS 2009 Conference Figure 2 shows a flat rubber sensor. Two conductive wires are set in each end to measure the resistance of electroconductive rubber. The deformation between these two contact points can be calculated by measuring the change of resistance. This sensor can be winded around the MPA (Fig. 1) to measure the circumference displacement. Using a model of the geometric structure of the MPA, the axial displacement can be estimated from the circumference displacement measured from this flexibile sensor. The hardness of flexible electro-conductive rubber is about 15 (ASKER C in SRIS0101 1), which is more flexible than the rubber used in MPA"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003713_978-1-4614-2419-2-Figure49.6-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003713_978-1-4614-2419-2-Figure49.6-1.png",
        "caption": "Fig. 49.6 Trajectories of points on cross-country pole",
        "texts": [
            " In both diagrams it can be clearly seen that the shooter is aiming the target which leads to lateral motion in x- and y-direction as well as rotation around the z-axis. At the point where the shooter pulls the trigger this motion is reduced to a minimum. Nowadays cross-country poles have a thin-walled design and are manufactured from carbon to reach a minimum weight. New developments show that the tip of the pole slightly shifted to the front leads to a defined buckling of the pole to the back under load, which supports the motion of the athlete. In the experiment carbon cross-country poles were used equipped with 484 H. Berger et al. Figure 49.6 illustrates the motion of the pole relatively to the stationary coordinate system. In addition the trajectories of five randomly chosen points are displayed as well. The PONTOS software is capable of subtracting the rigid body motion of any test object under load. Compensating the rigid body motion of the cross-country pole allows the evaluation of the buckling of the pole (see Fig. 49.7). In the beginning after 0.18 s it can be clearly seen that the pole buckles to the back. Just a little later after 0"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003887_j.procir.2015.06.103-Figure5-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003887_j.procir.2015.06.103-Figure5-1.png",
        "caption": "Fig. 5. The meshing coordinate system between the crown and internal spiral bevel gears",
        "texts": [
            " The meshing between the external and internal spiral bevel gears in the nutation drive can be considered as the external and internal spiral bevel gear meshing with crown gear (an imaginary gear), which has a pitch cone angle of 90 , and the pitch cone is at right angle to its axis. The pitch cone angle of the internal bevel gear is larger than 90 , and the pitch cone angle of the external bevel gear is less than 90 . The motion between the crown gear and the external or internal bevel gear can be considered as the pure rolling of two pitch cones. To establish the meshing function of spiral bevel gears, five coordinate systems are established in Fig. 4 and Fig. 5 to describe the meshing between the crown and external bevel gear and the meshing between the crown and internal bevel gear. This paper takes the meshing between crown gear and external spiral bevel gear as an example. The screw 1$ and 2$ represent instantaneous screw of spiral bevel gear and crown gear that relative to the frame, respectively, and screw 3$ presents relative spiral of 1$ to 2$ . The fixed coordinate system ),,( 0000 kjiS represents the original location of the external and internal spiral bevel gears attached coordinate system ),,( kkkkS kji )2,1(k in Fig. 4 and Fig. 5. The coordinate system ),,( 0000 kjiS fixed to the crown gear, represents the original location of the crown gear rotatable system. The coordinate system ),,( 1111 kjiS is attached to the external spiral bevel gear and rotates about the axis 1k of the centre of the external spiral bevel gear by angular speed 1\u03c9 in the Fig. 5, with 1 represents the rotational angle of the external spiral bevel gear. The coordinate system ),,( 2222 kjiS is attached to the internal spiral bevel gear and rotates about the axis 2k of the centre of the external spiral bevel gear by angular speed 2\u03c9 in the Fig. 5, with 3 represents the rotational angle of the external spiral bevel gear. The coordinate system ),,( ccccS kji is attached to the crown gear and rotates about the axis ck of the centre of the crown gear by angular speed c\u03c9 , with 2 represents the rotational angle of the crown gear. The direction of screw 1$ is along the axis )2,1(kkk of the centre of spiral bevel gear, and the direction of screw 2$ is along the axis ck the centre of the crown gear. For a pair of spatial meshing gears, not only rotating around the gear axis, but also moving along the axis",
            " The transformation matrix from cS to 0S is 1000 0100 00cossin 00sincos 22 22 0 cM (17) The transformation matrix from 0S to 0S is 1000 0sincos0 0cossin0 0001 11 110 0M (18) The transformation matrix from 0S to 1S is 1000 0100 00cossin 00sincos 11 11 1 0M (19) The transformation matrix from cS to 1S is 1000 0 0 0 333231 232221 131211 00 0 1 0 1 aaa aaa aaa cc MMMM (20) Where 2112111 sinsinsincoscosa 2112112 cossinsinsincosa 1113 cossina 2112121 sinsincoscossina 2112122 cossincossinsina 1123 coscosa 2131 sincosa 2132 coscosa 133 sina (21) The parameter 1 represents the pitch cone angle of external spiral bevel gear. Similarly, the transformation matrix from cS to 2S in the Fig. 5 can be obtained as 1000 0 0 0 333231 232221 131211 00 0 2 0 2 bbb bbb bbb cc MMMM (22) Where 2232311 sinsinsincoscosb 2232312 cossinsinsincosb 2313 cossinb 2232321 sinsincoscossinb 2232322 cossincossinsinb 2323 coscosb 2231 sincosb 2232 coscosb 233 sinb (23) 2.2. Tooth Equation of the Double Circular-Arc Spiral Bevel Gear As shown in Fig. 6, the basic tooth profile of the spiral bevel gears in the normal section is a double circular-arc profile, and adopts the profile of the model GB 12759-91 as the basic tooth profile, which consists of eight sections"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003702_cjme.2013.04.801-Figure13-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003702_cjme.2013.04.801-Figure13-1.png",
        "caption": "Fig. 13. Support phase of heavy-duty six-legged robot under ant-type tripod gait",
        "texts": [
            "5 N \u2022 m, respectively, when the value of rotation angle \u03b8  is zero, and the values of included angles \u03b22 and 2\u03b2  are 9\u00b0. with rotation angle \u03b8 of 0\u00b0 The axes of the abductor joints are perpendicular to the bearing platform. According to the structure of the robot, it is realized that the turning track of the abductor joint and the plane of the bearing platform are parallel to each other. Therefore, a maximum static torque exists for the abductor joint when the heavy-duty six-legged robot passes over a slope of 34.8\u00b0 using the ant-type tripod gait. A model of the support phase of the robot is shown in Fig. 13. Some parameters are defined in Fig. 13. The legs of the transfer phase are ignored, but their weights are added to the center of gravity of the bearing platform. Plane C intersects plane A at straight line l1. Planes B, C, and D are parallel to each other. The included angle is 90\u00b0 between plane C and plane A. Axes A1 and A6 are located in plane B. Plane C crosses the axes of A2 and A5. Plane D crosses the axes of A3 and A4. The rotation angle \u03b82 is between leg 2 and plane C. \u03b86 is a rotation angle between leg 6 and plane B. \u03b84 is a rotation angle between leg 4 and plane D\u2032",
            " To stop the bearing platform from rolling and pitching, the following conditions should be met: \u03b22\u03b24\u03b26 and 2\u03b2   4\u03b2   6\u03b2  . Based on the machine characteristics of the robot and the need to prevent the bearing platform from rolling, yawing, and pitching, a range of \u201330\u00b0 to 30\u00b0 is confirmed for \u03b82, \u03b84, and \u03b86. The foothold of leg 2 is defined as point 2o . Point 4o represents the foothold of leg 4. Point 6o is the foothold of leg 6. f2, f4, and f6 are respectively considered to be the frictions on footholds 2o , 4o , and 6o . Based on Fig. 13, some relations can be obtained: f2  2xF , f4  4xF , and f6  6xF . The analysis model for the maximum static torque of the ZHUANG Hongchao, et al: Method for Analyzing Articulated Torques of Heavy-duty Six-legged Robot \u00b7808\u00b7 abductor joint is shown in Fig. 14, where d2 is the length between foothold 2o and the vector of force 6xF . The distance from foothold 2o to the vector of force 4xF is expressed as d3. do represents the distance from foothold 2o to the vector of the G component in the direction of x"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001405_techpos.2009.5412069-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001405_techpos.2009.5412069-Figure1-1.png",
        "caption": "Fig. 1. Inverted pendulum, \u03b1 and \u03b8",
        "texts": [
            " Section 4 details the application of MPC to the inverted pendulum, and Section 5 provides experimental results. The rotary pendulum module consists of a flat arm which is instrumented with a sensor at one end such that the sensor shaft is aligned with the longitudinal axis of the arm. A fixture is supplied to attach the pendulum to the sensor shaft. The opposite end of the arm is designed to be mounted on a rotary servo plant resulting in a horizontally rotating arm with a pendulum at the end. The inverted pendulum made by QUANSER Company is shown in Fig. 1. The following equations describe the complete non-linear system; uKKKKK 43 2 221 )sin()cos( =++\u2212 \u03b8\u03b1\u03b1\u03b1\u03b1\u03b8 &&&&&& (1) 0)sin()cos( 625 =\u2212\u2212 \u03b1\u03b8\u03b1\u03b1 KKK &&&& (2) where 5,,1K=iK are constants, u is the signal applied to the DC motor, and \u03b1 and \u03b8 are shown in Fig 1. The linearized model of the system is derived by linearizing equations (1) and (2) about 0=\u03b1 . 1 (3) (4) where , and 1 0.001876 0.009308 6.3 10 0 1 0.00066 0.010 0.3665 0.8647 0.0018760 0.7927 0.1303 1 # 0.0012180.0011730.23790.2292 # $1 0 0 00 1 0 0% A sampling period, &', of 10ms is used. The open loop system is unstable. Consider the problem of regulating to the origin the discrete-time linear time invariant system described by Eqs. (3) and (4) while fulfilling the constraints ()* + + (,- , (5) ()* + + (,- (6) at all instants "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002135_13506501jet651-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002135_13506501jet651-Figure1-1.png",
        "caption": "Fig. 1 System of coordinates",
        "texts": [
            " Cases at low speed (500 r/min), for which the hydrodynamic effect remains weak (the minimum film thickness is reduced). Moreover, the heating effects are not significant (little variation in temperature in the film). 2. Cases at high speed (2600 r/min), for which both the hydrodynamic effects and the heating effects are significant. The basic equations of the THD and TEHD problems are written in cylindrical coordinates. The associated boundary conditions are also specified. In the chosen system of coordinates (Fig. 1), for a Newtonian fluid and under a laminar flow regime, the Proc. IMechE Vol. 224 Part J: J. Engineering Tribology JET651 at Harvard Libraries on July 2, 2015pij.sagepub.comDownloaded from generalized Reynolds equation (1) can be written as \u2202 \u2202r ( G1r \u2202p \u2202r ) + 1 r \u2202 \u2202\u03b8 ( G1 \u2202p \u2202\u03b8 ) = r\u03c9 \u2202G2 \u2202\u03b8 (1) The (O, r , \u03b8) plane is located on the active surface of the runner, rotating at an angular speed \u03c9 around the z axis. The G1 and G2 functions are defined by G1 = \u222bh 0 h \u03bc [ z \u2212 I2 J2 ] dz and G2 = 1 J2 \u222bh 0 z \u03bc dz with I2 = \u222bh 0 z \u03bc dz and J2 = \u222bh 0 1 \u03bc dz The energy equation (2) allows the calculation of the temperature field in the oil film"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000390_an9800500154-Figure5-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000390_an9800500154-Figure5-1.png",
        "caption": "Fig. 5. Calibration graph for the determination of glycerol in aqueous buffer solutions using immobilised GDH on silanised glass beads by covalent bonding.",
        "texts": [
            " The slope is reduced in the presence of serum, and pooled serum or serum control samples must be used for calibration. Alternatively, a standard-additions procedure can be used. The results were linear up to 4 x M in the diluted serum. 0.2 % 0.1 2 0 2 4 6 8 1 0 Lactate concentrationh x lo4 Fig. 4. Calibration graphs for the determination of lactate using LDH immobilised on silanised glass beads by covalent bonding. A, Aqueous buffer solutions; and B, serum control samples diluted 1 + 20 with buffer. The calibration graph for glycerol in aqueous buffer solution is shown in Fig. 5. It is non-linear in the range 104-10-3 M. Similar calibration graphs were obtained using standards of tristearin (after hydrolysis) in aqueous solution, and standards of glycerol in serum control samples, with a non-zero intercept for the latter due to the glycerol in the sample. Pu bl is he d on 0 1 Ja nu ar y 19 80 . D ow nl oa de d by S ta te U ni ve rs ity o f N ew Y or k at S to ny B ro ok o n 31 /1 0/ 20 14 0 2: 17 :2 3. pH EBect Immobilisation of enzymes frequently causes a displacement of the optimum P H "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000597_978-3-540-73129-0-Figure6.122-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000597_978-3-540-73129-0-Figure6.122-1.png",
        "caption": "Fig. 6.122. Microdrop injector. After [325]",
        "texts": [
            "121), a microbuffer (volume< 5 mm3) and an oil sensor offers a viable solution [324]. The oil sensor measures the level in the buffer and the pump provides the lubrication as needed. Dosing of the smallest quantities of liquid on the order of nanoliters and microliters was the goal in the cooperation between the Research Centre of Rossendorf and the GeSiM company of Dresden in the development of a microdrop injector [325]. The unit consists of a micro-injection pump (MEP) and a microsieve functioning as a diode for liquids (Fig. 6.122). The piezoelectrically driven injection pump functions similar to an ink-jet printer head, applying microdrops to the sieve. These droplets mix themselves with the liquid located below through surface tension. The microsieve makes use of surface tension effects to prevent the carrier liquid from soaking through into the injection chamber containing air. The disadvantage of half-opened systems is offset by the advantageous ideal liquid separation between the injection and carrier liquids by the air/sieve interface"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003713_978-1-4614-2419-2-Figure12.1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003713_978-1-4614-2419-2-Figure12.1-1.png",
        "caption": "Fig. 12.1 View of HEM",
        "texts": [
            " An engine mount must be able to provide large damping in the low-frequency range to limit the engine motion typically caused by road input forces transmitted through the vehicle suspension. However, the engine mount must also provide low stiffness with low damping for good vibration isolation in the high-frequency range [1, 2]. supports the static engine weight and provides the primary dynamic stiffness of the HEM and only a small amount of damping [3]. The mount is filled with a glycol fluid mixture of antifreeze and distilled water [7]. This was done through the fluid refill screw, while the lower chamber is sealed off with a rubber bellow, as shown in Fig. 12.1. This rubber bellow is extremely soft and makes almost no contribution to the dynamic properties of the HEM. Movement of the top rubber mount relative to the base causes a pumping action where fluid is pumped from the top to the bottom chamber and back. Upper and lower chambers are divided by an aluminium part that contains the decoupler and inertia track. The decoupler allows low resistance flow from the top to the bottom chamber if the amplitude were small. The inertia track is a spiral channel that also allows flow from the upper to the lower chamber"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000027_j.optlastec.2007.10.010-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000027_j.optlastec.2007.10.010-Figure1-1.png",
        "caption": "Fig. 1. Schematic arrangement of the ALFa setup.",
        "texts": [
            " It was observed that WC\u20136wt% Co and WC\u201312wt% Co were good for cladding. But as the thickness was increased, the process became unstable, leading to bulk defect in the material. Therefore, WC\u201317wt% Co was selected for the laser fabrication in the present study. Automated Laser Fabrication setup at the University of Waterloo, Canada, consists of a fiber-coupled 1 kW average power pulsed Nd:YAG laser system (LASAG FLS-1042N) integrated with powder feeder (Sulzer Metco: 9MP-CL) and laser workstation (Fadal: VMC 3016 with rotary and tilting table). Fig. 1 presents the schematic arrangement of the setup. In this setup, the laser pulses is generated at preset energy, pulse duration and pulse frequency using pulsed Nd:YAG laser. The beam is coupled to a 0.6mm core diameter optical fiber and transferred for the laser processing. The laser beam is collimated and focused using 115:115mm optics. The focal spot is kept above the fabrication point, allowing a defocused beam spot of 1.5mm for the material deposition. WC\u2013Co powder (size range 10\u201345 mm) is fed into the molten pool using a closed loop controlled powder feeder through a powder-feeding nozzle of exit diameter of 1"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002129_s0263574709005426-Figure7-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002129_s0263574709005426-Figure7-1.png",
        "caption": "Fig. 7. 3P3R Assur virtual chain.",
        "texts": [
            " Thus, the normalized screws corresponding to the virtual joints represented in the C-system are C $\u0302rz = \u23a1 \u23a31 0 0 \u23a4 \u23a6 ; C $\u0302px = \u23a1 \u23a30 1 0 \u23a4 \u23a6 ; C $\u0302py = \u23a1 \u23a30 0 1 \u23a4 \u23a6 . (17) It can be observed that the orthogonal PPR Assur virtual chain represents the movements in a planar Cartesian system. Other Assur virtual chains can be found in refs. [1\u20133]. represent the movements in the tridimensional Cartesian system the 3P3R Assur virtual chain is used. The 3P3R virtual chain is composed of three orthogonal prismatic joints (in the x, y and z directions), and a spherical wrist, composed of three rotative joints1 (in the x, y and z directions). Fig. 7 shows the 3P3R Assur virtual chain with the virtual links Ci labelled. Consider a kinematic pair composed of two links Ei and Ei+1. This kinematic pair has the relative velocity defined for a screw R$j (joint j) in relation to a reference frame R. The joint j represents the relative movement of the link Ei with respect to the link Ei+1. This relation can be represented as a graph,15 as shown Fig. 8. Where vertices represent links and arcs represent joints. The relative movement is also indicated by the arcs directions"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000597_978-3-540-73129-0-Figure6.62-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000597_978-3-540-73129-0-Figure6.62-1.png",
        "caption": "Fig. 6.62. Compliant spatial robot with 3 DOF and optimized design of a flexure hinge with 2 DOF [91]",
        "texts": [
            " The advantages of these hinges are the easy miniaturization ability and the natural lack of backlash, friction, and stick-slip effects. Since flexure hinges gain their mobility exclusively from a deformation of matter their attainable angle of rotation is strongly limited and the achievable movements and the workspace of these positioning devices are notably small. By using pseudo-elastic shape memory alloys as flexure hinges larger movements are possible. Due to the large reversible strains of SMA, deflections of the hinges of \u00b130 \u25e6 are achievable. Figure 6.62 shows a spatial compliant robot with 3 DOF (degrees of freedom) and six integrated combined flexure hinges. These combined hinges with 2 DOF and intersecting axes have replaced the conventional universal joints. The structure of the robot was developed for 3D assembly tasks with movements in x-, y- and z-directions. The robot is driven by three linear direct drives. Each drive is connected with the working platform by two links forming a parallelogram, allowing only translational movements of the platform and keeping the platform parallel to the base plane"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000538_jsen.2008.928923-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000538_jsen.2008.928923-Figure2-1.png",
        "caption": "Fig. 2. Structure and design of the ruthenium oxide sensor array. (a) Schematic diagram of the sensor array. (b) Fabrication process of screen-printed electrode.",
        "texts": [
            " Then, the graphite paste was screen printed onto the end of silver lines as the working area and backed in 100 C over for 10 min. Finally, the protective dielectric film of the biosensor was deposited, and then the sensor was exposed in the UV radiation on a UV-curing machine for 40 s. Also, the main emission line of the high pressure mercury lamp (500 W) is 365 nm and the distance between sensor and lamp is 15 cm. Specially, a black filter was used to protect the flexible substrate from heat damage. The schematic diagram of the sensor array was shown in Fig. 2(a). Fabrication of the electrode consists of five steps [shown in Fig. 2(b)], includes: 1) consecutive printing of silver track; 2) printing of carbon ink; 3) depositing ruthenium oxide thin film; 4) printing of insulating UV-gel; and 5) growing the planer reference electrode on the silver layer. Generally, ruthenium oxide thin films are prepared using several technologies. In this paper, it was deposited onto the working area by RF sputtering system, and the sputtering conditions of ruthenium oxide film were as follows: the deposition chamber was operated at 5 10 torr and the working pressure was maintained at 10 torr"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001555_09544062jmes1979-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001555_09544062jmes1979-Figure2-1.png",
        "caption": "Fig. 2 BFM: (a) before folding and (b) during the folding [8]",
        "texts": [
            " Then, some experimental static tests are applied on hexagonal aluminum columns and the graph of instantaneous folding force versus the axial displacement was sketched. Finally, the experimental results are compared with the theoretical predictions of the calculated relationship. Wierzbicki and Abramowicz [8] introduced a basic folding mechanism (BFM), calculated the dissipated energy rate, and the average folding force of a hexagonal column. In this study, the instantaneous folding force is calculated instead of an average value using the BFM. Figure 2 shows the BFM before and after the folding initiation. At the start time, \u03b1 = 0 and \u03b3 = 90\u25e6. At the commencement of the folding, \u03b1 increases and \u03b3 decreases continuously. When the folding is initiated, \u03b3 and \u03b2 varies with respect to \u03b1 and \u03c80 according to the following formulas [8] tg \u03b3 = tg\u03c80 sin \u03b1 , tg \u03b2 = tg\u03b3 sin \u03c80 (1) The instantaneous folding distance (axial displacement) designated by \u03b4 indicates the decrease of the axial distance between the upper and the lower edges of the BFM. As shown in Fig",
            " In equation (4), the first integral calculates the dissipated energy of extensional deformation on the small area, called toroidal surface, and the second integral the dissipated energy of the inextensional deformation, or, in other words, the dissipated energy of bending around the hinge lines. Performing the first integration results in [8] E\u03071 = 4N0bH\u03c0 (\u03c0 \u2212 2\u03c80) tg\u03c80 cos \u03b1 { sin \u03c80 sin ( \u03c0 \u2212 2\u03c80 \u03c0 \u03b2 ) + cos \u03c80 [ 1 \u2212 cos ( \u03c0 \u2212 2\u03c80 \u03c0 \u03b2 )]} \u03b1\u0307 (5) where b is the small radius of toroidal surface and angles \u03b1, \u03b2, and \u03c80 are shown in Fig. 2. The dissipated energy rate in equation (5) refers to the extensional deformation on toroidal surface which is the value of the first integral in equation (4). The second integral in equation (4) was separately calculated for the fixed horizontal hinge lines and then for the inclined hinge lines. In other words, the second integral in this equation shows that the inextensional deformations involves bending around the fixed horizontal hinge lines AB and BC and bending around the inclined hinge lines UB and BL designated by E\u03072 and E\u03073, respectively",
            " The length of each edge of the hexagon is equal to 2C . Thus, the summation of the internal dissipated energy rate in the hexagonal column is calculated as follows E\u0307int = 6E\u03071 + 6E\u03072 + 6E\u03073 (8) where E\u03071, E\u03072, and E\u03073 are results of equations (5), (6), and (7), respectively. The term E\u03072 from equation (6) refers to the bending around the horizontal hinge lines AB and BC during the first fold creation in a BFM, but actually, in a hexagonal column, the bending process happens around the horizontal hinge lines AB and BC, as well as DL and LG (Fig. 2). Thus, equation (8) is rewritten as E\u0307int = 6E\u03071 + 12E\u03072 + 6E\u03073 (9) The external work rate in a compressing process of a hexagonal column is calculated as E\u0307ext = P\u03b4\u0307 = P2H sin \u03b1\u03b1\u0307 (10) where P is the external force on the hexagonal column under axial loading.Thus, using the following equation E\u0307ext = E\u0307int (11) Proc. IMechE Vol. 224 Part C: J. Mechanical Engineering Science JMES1979 at University of British Columbia Library on June 22, 2015pic.sagepub.comDownloaded from which shows that the external work rate, which is required for compressing in a column, is equated with the internal dissipated energy rate, the following relation is reached P2H sin \u03b1\u03b1\u0307 = 6E\u03071 + 12E\u03072 + 6E\u03073 (12) Considering the hexagonal column geometry, it is shown that the external angle of the square column is 2\u03c80 = 60\u25e6 and so \u03c80 = 30\u25e6"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002949_etep.1642-Figure5-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002949_etep.1642-Figure5-1.png",
        "caption": "Figure 5. The magnetic field distribution.",
        "texts": [
            " equals the supply frequency so that rotor-related skin effects and the saturation of the upper rotor tooth can be more severe than the rated load. Furthermore, the rotor magnetic field distribution is no longer symmetrical upon a broken bar fault, as illustrated in Figures 5 and 6. In addition to changes in the broken bar regions, the field and flux density distributions at other positions in the stator and rotor core are also distorted and increased to some extent. It can be seen that the presence of a broken bar in Figure 5(b) or two broken bars in Figure 5(c) results in highly saturated regions in the neighboring core iron that may cause a progression of the fault. If observed around the air gap, the radial and tangential components of flux density waveforms at one time instant can be obtained, as presented in Figure 7. In order to study the variation of the air-gap flux density at standstill, the frequency spectrum of the air-gap flux density was analyzed by the harmonic frequency analysis. Figure 8 shows that the air-gap flux density contains high spatial harmonics components and varies significantly with broken bar faults"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002790_mesa.2012.6275561-Figure6-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002790_mesa.2012.6275561-Figure6-1.png",
        "caption": "Fig. 6. Schematic representation of the haptic sensors. The kinesthetic sensor is at the shaft of the end-effector, the tactile sensor can either be mounted at the gripping surface (position 1) or at the tip of the end-effectors (position 2).",
        "texts": [
            " For our kinesthetic sensor, the approach of [6] is used as groundwork and adapted to our special needs. This means further reducing the size of the sensor. This is done by reducing the number of sensing points to three, one for the force in each spatial dimension. Torques acting on the instruments tip only occur very rarely, so they can be neglected. There are two possibilities where the tactile sensors can be integrated in the end-effectors: at the gripping surface or at the front of the instrument, cf. Fig. 6. While the first possibility would enable slip detection of e.g. needles, the second possibility is more reasonable for examining tissue: the surgeon can slide over tissue surfaces and the danger of harming the tissue is lower. Unlike the approaches using FBGs for tactile sensors presented in Chapter II, in our approach the FBGs are inscribed at the tip of a fiber. Forces are then applied axial to the fiber, not lateral. This implicates that the grating is not strained uniformly, which leads to a chirped signal, see Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002717_0954406212454390-Figure6-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002717_0954406212454390-Figure6-1.png",
        "caption": "Figure 6. Geometry when gears meshing at point Mc.",
        "texts": [
            " dV \u00bc dVi dVo \u00bc \u00f0dVi1 \u00fe dVi2\u00de \u00f0dVo1 \u00fe dVo2\u00de \u00f021\u00de where dVi1 \u00bc 1 2 r2a1d 1, dVo1 \u00bc 1 2 r2f1d 1, dVi2 \u00bc 1 2 r2a2d 2, dVo2 \u00bc 1 2 r2f2d 2: \u00f022\u00de Here, the axial thickness of the gear is a unit thickness. Given the fundamental law of gearing, it is known d 2 \u00bc r1 r2 d 1, !1 \u00bc d 1 dt \u00f023\u00de Equations (21) to (23) yield the instantaneous flowrate formula Q \u00bc dV dt \u00bc 1 2 \u00f0r2a1 r2f1\u00de r1 r2 \u00f0r2a2 r2f2\u00de !1 \u00f024\u00de at SIMON FRASER LIBRARY on June 16, 2015pic.sagepub.comDownloaded from During the delivery process, rf1 and rf2 can be represented by equations (25) and (26) (Figure 6). r2f1 \u00bc f 2inv \u00fe r21 2finvr1 cos \u00f025\u00de r2f2 \u00bc f 2inv \u00fe r22 2finvr2 cos \u00f026\u00de where finv is the distance from the contact point Mc to the pitch point P. Equations (24) to (26) yield Qinv\u00f0finv\u00de\u00bc 1 2 r2a1 r1 r2 r2a2 r1\u00f0r1 r2\u00de 1 r1 r2 f2inv !1 \u00f027\u00de It is clear finv \u00bc ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi \u00f0xMc f xPf \u00de 2 \u00fe \u00f0 yMc f yPf \u00de 2 q , \u00f0xMc f , yMc f , 1\u00de 2 RAB f \u00f0\u20191, xt\u00de \u00f028\u00de where (xMc f , yMc f ) is the coordinate of point Mc and RAB f is the position vector of line AB which belongs to Rl f"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003726_20130828-3-uk-2039.00050-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003726_20130828-3-uk-2039.00050-Figure2-1.png",
        "caption": "Fig. 2. Coordinate system for the 3-DOF helicopter. The elevation, pitch and travel angles, E(t), \u03a8(t), \u0398(t), respectively, are shown in their positive sense. Also shown are the fan thrust vectors from the advance and back fans, Fa(t) and Fb(t), respectively.",
        "texts": [
            " The helicopter body consists of two independently controlled fans connected together by a short arm. This arm is free to pitch around a midpoint that is pivoted at the end of a second bar, termed the elevation arm. As the fans do not produce enough thrust to overcome their own weight, mechanical assistance is provided by an adjustable counterweight at the opposite end of the elevation arm. This arm is free to rotate about the horizontal plane, and is mounted atop a vertical shaft that is free to spin freely within a cylindrical housing. The coordinate system is shown in Figure 2. In this way the helicopter can \u2018fly\u2019 across a significant portion of the surface of a sphere with a radius determined by the length of the elevation arm to the helicopter body. The elevation angle was measured with an analogue servo potentiometer, whilst the travel angle was measured with a digital rotary encoder. The fan control signals and elevation angle measurement signals were interfaced to the power and signal conditioning electronics via a miniature slip-ring assembly housed within the cylindrical base of the helicopter"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000795_s0006-3495(77)85536-7-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000795_s0006-3495(77)85536-7-Figure1-1.png",
        "caption": "FIGURE 1 (a) Bead model for a helical flagellum. Note the three rotations: w (flagellum), Ql (head), and fl (flagellum). Symbols are defined in the text. (b) Detail of the arrangement of beads. The projections of the beads on the x-axis are equally spaced.",
        "texts": [
            " (2c) The position of the head is given by XN = Lx+ a + Vt, (3a) YN = ZN = 0. (3b) By differentiation of Eqs. 2 and 3, one obtains vxi = V, (4a) i = O, 1 ... N 1 ivyi -bwsin(kxi wt), (4b) ivlN = bwcos(kxi + wt), (4c) VXN = V, (5a) VYN = VZN = 0. (5b) for the Cartesian velocities of the beads. In addition to these velocities, the head must rotate around the x-axis and the flagellum around the tangent of the center line, each with angular velocity Q in the opposite direction to W, as indicated in Fig. 1. This fact, discovered by Chwang and Wu (4), is necessary to cancel the total torque, as will be shown later. Hydrodynamic Interaction In the preceding section the geometry and kinematics of the helical bead model were presented. Now we will show how the dynamics of the particle can be deduced by using an interaction tensor formalism, as is common in polymer hydrodynamics. The force exerted on the solvent by the ith bead of the model, Fi, is proportional to the relative velocity v, of the bead with respect to the fluid Fi= ,v (6) where Di is the Stokes' law frictional coefficient of the spherical bead Di = 6irno r, i = 0, 1, "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000558_j.1467-9450.1977.tb00279.x-Figure6-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000558_j.1467-9450.1977.tb00279.x-Figure6-1.png",
        "caption": "Fig. 6. The response apparatus (the letters defined in the text).",
        "texts": [
            " Thus the angular size of the visual field was about 53\u201d~37\u201d and the visual angle of the proximal stimulus in turning position 1 was 20\u201d. thus giving a correct projection of the \u201cdistal\u201d stimulus. The subject looked binocularly at a specially built re- Scand. J . Psvchol. 18 sponse apparatus (E) and responded by setting slant and curvature of the two perceived turning positions of each stimulus on this apparatus. The lower part of it was covered by white cardboard (0. The construction of the response apparatus is shown in Fig. 6. The central portion of it (A) was a piano string (0.5 mm thick) covered with a black tube shrunk onto it (2.3 mm outer diameter when shrunk). This bendable string was given different curvatures by the subject\u2019s pulling or pushing at B and fastening the string in the desired position with a stopping screw (C). The distance between C and E was 13.5 cm. The slant of the string was adjustable with the aid of a wheel (D). The curvature was recorded with a potentiometer ( E ) which position was read by a digital ohm-meter"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001003_j.engfailanal.2006.11.052-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001003_j.engfailanal.2006.11.052-Figure2-1.png",
        "caption": "Fig. 2. Identification of the strain gauges on the housing gearbox (motor side).",
        "texts": [],
        "surrounding_texts": [
            "In this work nine rosettes were bonded (45 rosettes \u2013 3 linear strain gauges) in the following positions: rosette 1 in the central zone of the cover of the housing to measure the stresses in the cover; rosette 2 in a zone away from weld areas. In the box body, to measure the nominal stresses; rosette 3 in the weld toe of weld zone on the lower reinforcement box, place where fatigue cracks were started; rosettes 4\u20139 were bonded in the six supports of the housing to obtained the reaction forces (Figs. 2 and 3). The same procedure of surface preparation has been conducted by the authors in the study of cast steel railway coupling used for coal transportation [7]. The data was collected with a portable PC and the HPVEE processing system was used for treatment and analysis of the signals. The data was obtained in service, in the Lisbon\u2013Porto Intercity passengers train, with a maximum speed of 160 km/h and in the Entroncamento\u2013Guarda freight train, with a maximum speed of 120 km/h (Table 3). The program of the rosettes readings were made in order to get the best possible comparative information. Those stages are the critical section of the journey, in what concerns speed, power and track oscillations. For the Lisbon\u2013Porto journey in the Intercity, the following 6 stages were selected: Entroncamento\u2013Fa\u0301tima: Acquisition time \u2013 894 s; Number of km \u2013 23.3 km; Fa\u0301tima\u2013Pombal: Acquisition time \u2013 1180 s; Number of km \u2013 40 km; Pombal\u2013Alfarelos: Acquisition time \u2013 1037 s; Number of km \u2013 29 km; Alfarelos\u2013CoimbraB: Acquisition time \u2013 900 s; Number of km \u2013 18.7 km;"
        ]
    },
    {
        "image_filename": "designv11_12_0000177_tro.2006.882921-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000177_tro.2006.882921-Figure2-1.png",
        "caption": "Fig. 2. Guaranteed region p p q q near a convex obtuse vertex.",
        "texts": [
            " To make an argument for a region, we need to look at the four different possible cases separately. 1) Convex Obtuse Angle Corners: When one sink is moved from a vertex to another point along one of the two sides, the potential on one side monotonically increases, while the other side monotonically decreases. The only difference of the obtuse corner from the straight line is that equivalent potential changes have to be considered on each side. So, we find the length of a segment on side one for a given segment such that (see Fig. 2). Then, is the worst change of the potential, Now, we extend the idea to a region where we can guarantee no pivot point exists. We make an argument similar to the one for moving from an obtuse vertex along the side, looking at each side separately. For side one, since for the motion from to , some points on side one increase in potential while some decrease, we want to prove that the length of a segment on side one with no net change in potential (segment ) is bigger than the length corresponding to the movement of the sink from to for a given finite ",
            " 4) Concave Acute Angle Corners: The bound when a sink is moved along a side of a concave acute corner is analyzed in the same way as the convex acute case. The bound for the region around a concave acute corner is analyzed in the same way as the concave obtuse case. Fig. 6 shows the geometry, although we omit the details for brevity. These bounds provide the means for an algorithm to peel off the boundary, as given in Algorithm 2. Algorithm 2: Guaranteeing no pivot points near the boundary Require Ensure perimeter {if not, reduce } if CORNER then if OBTUSE then % Step: where is found by (Fig. 2). % Region: if Convex then Side 1: where satisfies (Fig. 2) Side 2: (Fig. 2) else {Concave} Side 1: (Fig. 3) Side 2: (Fig. 3) end if else {ACUTE} % Step: (Fig. 4) % Region: if Convex then Side 1: (Fig. 5) Side 2: (Fig. 5) else {Concave} Side 1: (Fig. 6) Side 2: (Fig. 6) end if end if else {SIDE} (Fig. 1) end if For a given object, we have checked each vertex and the boundary, and peeled off the boundary all around the object with a finite thickness everywhere, except points on the boundary where the force is zero. The analysis on the existence of a finite resolution (away from critical points) is guaranteed, and suggests a constructive algorithm"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003348_pime_proc_1970_185_115_02-Figure11-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003348_pime_proc_1970_185_115_02-Figure11-1.png",
        "caption": "Fig. 11. A back-to-back layout of the rolling contact drive",
        "texts": [],
        "surrounding_texts": [
            "D380 DISCUSSION ON S. Y. POON\ncomments, whilst acknowledging the qualitative conclusions of the strip approximation method, are intended to convey some doubts about its validity as an accurate design tool.\nConsider two friction drive point contacts, operating under e.h.1. conditions, with spin, as described in the author's paper. Assume they are elliptical with semi-major axes bl and b, respectively and of the same eccentricity. According to equation (1 l), the relative velocity distributions are given by:\nneglecting c for the purpose of this argument. effective traction coefficients are given by equation (6): Taking the same number of strips in each contact the\nIt can be seen that, for a given Hertzian maximum pressure, provided that U, = U,, G, = Gz, and 0, = SZ,, and that the same disc machine data is used in each case, the strip approximation suggests that the traction coefficient-sliding speed characteristic is identical for each contact. Figs 6 and 7 appear to substantiate this conclusion in that at given P,,,, U, sliding speed, and temperature, the effective coefficient of traction depends only on the spin parameter.\nExperimental evidence obtained in our laboratories, using friction drive components kinematically identical to those illustrated by the author, is not consistent with this interpretation. Typical values are given below :\nRolling speed, 1400 in/s Contact surface temperature at inlet, 80\u00b0C Hertzian maximum pressure, 316 000 lb/in2\nG Q a b Pmax 0.002 0.051 0.0190 0.0397 0.041 0.003 0.049 0.0381 0.0807 0.030\nwhere pmax is the limiting coefficient of traction. The lubricant in the tests was a synthetic fluid developed for use in rolling contact friction drives.\nThe effect of the differences in G and l2 on the relative velocity distribution is insignificant, leading to the conclusion that a scale effect exists which causes a reduction in traction coefficient as the contact size is raised. Further evidence of a scale effect is contained in the contact radius term of Hint's (12) equation for the coefficient of traction in a circular contact at high sliding speed.\nIncidentally, measurements of the traction over a wide range of sliding speeds have given cutves typified by Fig. 9. It shows an almost linear rise of traction up to a maximum of about 5 per cent sliding, compared with the 05-1 per cent mentioned for non-spinning contacts. It also\n007 \\\n0.04\n0.03\n0.01 ' 2 I 0-\n0 0.1 0.2 0.3 0.4 SLIDE/ROLL RATIO 6.0.033 in, u;0.019 in N=3501b,U=200 i n h S2= 0035 ,G=00013 Roller surface temperature = I00\"C\nratio for a spinning joint contact\nshows a steady decay as slide ratio increases beyond 5 per cent. It should be emphasized, however, that this behaviour was recorded with constant roller surface temperature at inlet to the contact, and is probably not representative of conditions in a real transmission if the maximum traction coefficient is exceeded.\nFig. 9. Variation of traction coefficient with slide/roll\nR E F E R E N C E\n(12) PLINT, M. A. 'Traction in elastohydrodynamic contacts', Proc. lnstn rnech. Engrs 1967-68 182 (Pt 11, 300.\nS. Y. Poon Graduate (Author) M. J. French is quite correct that the parameter G related to the geometry of the contact should include a factor\nR R+R, sin 6\nif the second-order term in y is taken into account in the first of equations (2) : the term in question is\nOn the question of the relative effect of temperature and pressure on the traction, the answer is that it not only depends on the type of lubricant, but for the same lubricant, it is affected by the temperature range. For instance, the effect is not nearly as pronounced between 25\" and 50\u00b0C as between 50\" and 75\u00b0C for the mineral oil considered in the paper. The reason is that, at a lower temperature range, the maximum traction coefficients are closer to the upper limit or 'ceiling' of the lubricant (10) for pressure in excess of 176 000 lb/in2. A number of synthetic fluids for rolling contact drive tested seem to indicate that the coefficient of traction is less sensitive to temperature variation than the mineral oils under similar conditions, but at higher operating temperatures or lower contact pressure the behaviour of these fluids exhibits a pattern not unlike the mineral oil. The nature of relative significance of temperature and pressure on the traction has been explored from the viewpoint of physics of fluid by the\nProc lnstn Mech Engrs 1970-71 Vol 185 76/71\nat University of Leeds on June 5, 2016pme.sagepub.comDownloaded from",
            "THE EFFECT OF SPIN ON THE TRACTIVE CAPACITY OF ROLLING CONTACT DRIVES D381\nauthor in his contribution to the discussion of a paper by Plint (12).\nH. Gaggermeier has found that the traction results from a laboratory disc machine are sensitive to even a very small amount of shaft misalignment. Shaft misalignment can be resolved into a spin component and a transverse sliding component (Fig. lo). Even for a small skew-angle, /3, it can be seen that the transverse sliding component, in terms of slide ratio, can be of the same order of magnitude as the forward slide ratio in the region we are concerned with here. One degree of skew, for example, corresponds to a transverse slide ratio of 1 7 . 5 ~ It is, therefore, not surprising that the traction that can be supported in the direction of rolling is greatly reduced, The conclusion, however, in no way invalidates the strip approximation in which the transverse slide ratio is small in comparison with that of the rolling direction,\nS . Lingard has made an interesting observation on the \u2018scale effect\u2019 which causes a reduction in traction coefficient as the contact size is raised. In analysis, it is perhaps more profitable to consider it in terms of the parameters used in the modelling of the physical system or the assumptions underlining the analysis which are likely to be affected by the physical scale of the problem. It would appear that on close examination the thermal effect is the likely cause as the contact dimension is increased. Judging from the data given by S . Lingard, the contact load on the second test, at 2000 lbf per roller, is about four times the load on the first test if the same Hertzian pressure is maintained, and there will be a corresponding increase in viscous heating at the contact.\nAssuming the drive consists of two sets of rollers arranged back-to-back (Fig. 1 l), the rate of heat generating at all the contacts in the second test is in excess of 4 kW. In view of the compactness of the design, high local thermal gradients must exist, and these will be particularly severe on the central disc. Experiences have shown that even with the much simpler configuration of a disc machine, the surface temperature of the discs cannot be effectively controlled by the inlet oil jet temperature when a high ratio of heat is generated. This is due to the presence of a stable, laminar thermal boundary layer (1.1) and (13). It would be too much to expect that uniform temperature can be achieved with the test rig. S. Lingard\u2019s contribution highlights the need for further work on the heat transfer aspect of the rolling contact drive.\nREFERENCE\n(13) O\u2019DONOGHUE, J. P. and CAMERON, A. \u2018Friction and temperature in rolling sliding contacts\u2019, ASLE Trans. 1966 9, 186.\nProc lnstn Mech Engrs 1970-71 Vol185 76/71\nat University of Leeds on June 5, 2016pme.sagepub.comDownloaded from"
        ]
    },
    {
        "image_filename": "designv11_12_0003159_978-1-4419-9985-6-Figure1.6-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003159_978-1-4419-9985-6-Figure1.6-1.png",
        "caption": "Fig. 1.6 Illustration of the spring system of the transducer. The ring above the transducer (to the left) is mounted on the spring supported arms. These arms are connected to the rigid cross via three springs as shown in the right figure",
        "texts": [
            "5 illustrates the mounting of the MEMS structure on the transducer. The central grey region of the sensor element (see Fig. 1.5 to the right) is glued on the rigid central cross of the transducer, which acts as a base structure. The four outer grey regions (stud bumps) are glued on four arms of the transducer, which contain the spring system. Together with the sensor element these arms form a second cross. Beside the spring close to the sensor element, each of the four arms also has springs in its other end as illustrated in Fig. 1.6. According to the right of Fig. 1.6 there are three springs for each arm. The two springs at the sides of the arm end connect to the rigid cross, making it possible to fabricate the transducer in a single sheet of metal. The constellation of the three springs increases the isotropy of the sensor by to some extent compensate for the earlier mentioned lack of isotropy of the sensor elements. The ring above the transducer in Fig. 1.6 is used to simulate the rigid mounting of the spring arms on the sensor flange. The rigid cross is mounted in the sensor housing. The diameter of the transducer was selected to be 100 mm, steel thickness 1 mm and the smallest spring width was 0.2 mm. With these parameters simulations of a transducer with mounted sensor element gave results as shown in Table 1.2. With a maximum allowed Silicon stress level of 300MPa, this transducer could be possible to use for forces up to 40 N and torques to 3"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001038_j.finel.2008.01.006-Figure6-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001038_j.finel.2008.01.006-Figure6-1.png",
        "caption": "Fig. 6. Adding balance masses and distribution of continuous eccentricity masses.",
        "texts": [
            " As explained earlier, the silicon steel core consists of a series of stacked steel laminations that swirls around the central shaft with a designated angle. It is straightforward to assume that each of the steel laminations are all identical and with similar mass eccentricity. Also, the whole stack is rotated with a skewed angle of 24.5\u25e6 counting from the first lamination up to the last lamination relatively. It is then reasonable to assume that there exists an eccentricity line that links the eccentricity of each lamination on the silicon steel core which swirls up along the shaft. Fig. 6 shows the overall side view and the lateral view of the whole silicon steel core assembly. The mass eccentricity line is divided into five equal elements as shown at right in the lateral view of the whole assembly. Besides, the mass eccentricity of the silicon steel core was lumped into nodal masses named as m23 through m28 corresponding to each of the element nodes 23 through 28. The radius of eccentricity, e, is defined as the distance of the located lumped mass eccentricity that is relative to the geometric center on each of the silicon steel laminations",
            " Therefore, the bearing damping is recommended to be introduced into the system to reduce the vibration amplitude no matter at the resonance point or elsewhere. Although the vibration amplitude of the induction motor has been investigated thoroughly as addressed in the foregoing descriptions, the most important and meaningful work is to suppress these vibrations ultimately. Therefore, considerable adding balance masses are attached on the balance planes so as to suppress the vibrations. The relationship between the adding balance masses and eccentricity masses can be interpreted through what has been shown in Fig. 6. The adding balance masses at the two balance planes are represented as m1 and m2. The phase angle 1 between m23 and m1, and the phase angle 2 between m1 and m2 are as defined correspondingly in the figure. The eccentric masses at the nodes of silicon steel core are noted to be from nodes 23 to 28 and are designated as m23 to m28 with a total skew angle of 24.5\u25e6. The circular mounting locations of the adding balance masses have already been set as em =32.15 mm because the physical protrusions (as indicated in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003348_pime_proc_1970_185_115_02-Figure8-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003348_pime_proc_1970_185_115_02-Figure8-1.png",
        "caption": "Fig. 8. Velocities in the contact area of a friction drive",
        "texts": [
            " The experiments with two-disc machines have shown that the coefficient of traction differs greatly from the slide ratio. The author shows how the coefficient of traction at rolling-spinning motion can be calculated from a series of experiments on a two-disc machine, using a summation resulting from the strip approximation. There he used the simplification that the slide ratio is the same along the whole strip as on the centre line of the contact ellipse, where the velocities u1 and have the same direction. Fig. 8 shows the contact area of the drive in Fig. 1. At the centre line there is shown a small skew-angle /3 between the velocities u1 and u2. In my own experiments on the two-disc machine of the Laboratory for Machine Elements at the Munich Technical University, a great influence on the coefficient of traction was found already at skew-angle /3, which is much smaller than one degree. If this sliding motion resulting from the skew-angle /3 were considered in the calculation, then the coefficient of traction would rise more slowly over the slide ratio"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003713_978-1-4614-2419-2-Figure14.18-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003713_978-1-4614-2419-2-Figure14.18-1.png",
        "caption": "Fig. 14.18 Schematic of 2 D. O.F model used to design isolation system",
        "texts": [
            " Future work will include similar torsional vibration measurements on the shaft as the test frame accelerates linearly on a shaker table, in order to study the cross-axis sensitivity of each of the sensors. Acknowledgements The authors would like to thank Cummins Power Generation and those involved with the testing. Those individuals include Phil DeBerry from Anger Associates Inc., Laurent Britte and Bill Flynn from LMS International, and Doug Hanson, Joe Hauser, and Jason Cheah from Cummins Power Generation. Appendix Fig. 14.18, which is appropriate if the mass of the belt is much less than the mass of the flywheel and motor and if the modes of vibration of the motor and flywheel are much greater than the frequency band of interest. The parameters of the system are: I1 and I2, the inertia of the flywheel and rotor mass respectively, r1 and r2 the effective radii of the flywheel and motor pulleys, y1 and y2 their respective angles and K the stiffness of the belt. The disturbance torques t1and t2 are also shown. Dissipation in the belt can also be included with a viscous damping parameter, C"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000388_14644193jmbd97-Figure7-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000388_14644193jmbd97-Figure7-1.png",
        "caption": "Fig. 7 Displacement of the bearing centre due to rocking motion",
        "texts": [],
        "surrounding_texts": [
            "The curvature difference for a Hertzian contact is found to be a function of the elliptical parameters as follows (for more information see [34]) F\u00f0r\u00de \u00bc \u00f0k 2 \u00fe 1\u00de= 2@ \u00f0k2 1\u00de= \u00f01\u00de where k is the elliptical eccentricity parameters and @ and = are complete elliptic integrals of the first and second kind, respectively. By assuming values of the elliptical eccentricity parameter, a table of k versus F(r) can be obtained and then it can be shown that the local force and deflection relationship of two bodies may be written as follows [34] d \u00bc d 3 2 WP r 1 n21 E1 \u00fe 1 n22 E2 2=3P r 2 \u00f02\u00de where d* is dimensionless deflection and is expressed as d \u00bc 2@ p p 2k2= 1=3 \u00f03\u00de d* is given as a function of F(r) [15]."
        ]
    },
    {
        "image_filename": "designv11_12_0001847_1.4000269-Figure4-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001847_1.4000269-Figure4-1.png",
        "caption": "Fig. 4 Geometry of journal bearing under starved lubrication",
        "texts": [
            " 3 ii , the journal bearing is in an unstable state because cavitation generation is not observed. On the other hand, at the restriction of the supply flow rate to Qin =1.4 10\u22126 m3 /s and Qin=0.5 10\u22126 m3 /s, the amplitude is small and generation of cavitation is observed in a wide range of numbers of revolutions. It is concluded that stabilization is attainable using the starved lubrication effect. This phenomenon is explained by generating cavitation in a wider range of the bearing clearance and forming an oil film focused on a region adjacent to the loading line. Figure 4 is a schematic diagram of a cylindrical journal bearing. Under starved lubrication conditions due to the restriction of the oil supply, an amount of lubricating oil in a wedge region is insufficient, and a region of formation of an oil film varied with the amount of oil supplied. Therefore, in this study, an analysis is under flooded and starved lubrication ts Transactions of the ASME of Use: http://www.asme.org/about-asme/terms-of-use p o 5 l w d e e T t fl e u a c o w s t Q m i J Downloaded Fr erformed using a model for calculating a change in a region of il film formation by changing the amount of oil supplied",
            "org/ on 09/01/2017 Terms QV = z1 z2 1 2 h t Rd dz 4c To perform an analysis, taking into consideration a change in the flow in the oil supply, repeated calculations are carried out in the process of analysis by correcting film formation regions at the inlet and the outlet applying boundary conditions considered to provide the oil film flow balance. The flow rate at arbitrarily chosen grid points i , j is calculated using the following equation: qi,j = \u2212 hi,j 3 12 R p i,j + Uhi,j 2 5 If a flow of the oil supplied to the test bearing is considered as an average flow relative to the axial direction see Fig. 4 , the flow of the oil supplied per unit width can be expressed by the following equation: qin = Qin/L 6 Assuming that the oil flow qe at the oil film outlet and the flow qin, of the supplied oil both reach the oil film inlet, the oil flow qs at the oil film inlet can be expressed as a sum of both flows as follows: qs = qe + qin 7 Therefore, applying the Reynolds condition, at which a pressure gradient at the outlet equals to zero, the oil flow at the oil film outlet can be expressed as follows: qe = Uhe/2 8 In the analysis, calculations start under flooded lubrication conditions considered to be the initial state"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003274_ssp.210.26-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003274_ssp.210.26-Figure1-1.png",
        "caption": "Fig. 1 Conventional independent suspension - McPherson strut",
        "texts": [
            " Because of the constant characteristics of spring and damping components they are not able to cope with the increasingly sophisticated demands of automotive suspensions. The use of non-conventional suspension such as hydropneumatic suspension can eliminate some of the disadvantages of conventional suspensions [5,6]. The role of spring components in conventional suspensions fulfils generally coil springs (less often torsion bars and leaf springs) [1]. This element has constant spring characteristics. Independent kind of suspension \u2013McPherson strut is showed on fig.1. 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: 130.88.90.140, The University of Manchester, Manchester, United Kingdom-05/05/15,19:17:18) Increasing static load in mechanical suspension increases the static deflection of spring , and for full load the available wheel displacement is significantly reduced (this is one of the dominant disadvantages of this type of suspension)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000894_iros.2008.4651029-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000894_iros.2008.4651029-Figure1-1.png",
        "caption": "Fig. 1. The parallel and flexible meso-manipulator structure.",
        "texts": [
            " The presented studies do not anyway allow the analyses application to patients with even a partial lack of the fingers functionality, for instance due to a stroke or to a surgical finger tendon reconstruction [23]: however such techniques can contribute relevantly for an objective monitoring assistance to the rehabilitation activity of the patient, who needs to avoid the muscular enfeeblement or shortening [23]. The currently presented system takes place in this field. The parallel meso-manipulator herein concerned is shown in figure 1: in this paper the problems of kinematics and dynamics are addressed with a particular attention to the robotic compliance. Two different aspects of this property can be underlined: the flexure, in the hinges chosen as joints, assure a functional compliance, while a collateral compliance, that allows the compensation of high and unexpected forces and helps the motion, can be identified from the global architecture flexibility Strong points of this architecture are: the possibility to realize also wide motions, the positioning accuracy obtained through more close kinematic chains, and the structure capacity of fronting the arise of resistant and also finite forces while the platform remains always parallel to itself"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001864_jp102207w-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001864_jp102207w-Figure1-1.png",
        "caption": "Figure 1. Upper view of dialysis system and notation in the theory.",
        "texts": [
            "17 If we dissolve chitosan molecules in acetic acid aqueous solution and put it into a dialysis tube and then dialyze it into sodium hydroxide solution, we have a hydrogel, because the charges are lost and lots of hydrogen-bonding sites in chitosan interact to make a cross-linking network. Therefore, the inflow of sodium hydroxide and the outflow of acetic acid play the role of changing the charges of chitosan molecules to meet the gelation condition, but those molecules do not take part in the cross-linking points. We first develop a theory for this case and then compare it with experimental results. The illustration of the experimental system and the notation is given in Figure 1. The chitosan solution is sandwiched between a set of cover glass with the radius R and immersed into an extradialytic sodium hydroxide solution. F0\u2032, Fs\u2032, F0, and Fs denote the concentrations of sodium hydroxide near the core and near the extradialytic solution in the gel layer, in the core solution, and in the extradialytic solution, respectively. C0\u2032, Cs\u2032, C, and Cs are the concentrations of acetic acid near the core and near the extradialytic solution in the gel layer, in the core solution, and in the extradialytic solution, respectively"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001884_s11431-010-3100-y-Figure5-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001884_s11431-010-3100-y-Figure5-1.png",
        "caption": "Figure 5 The 7-bar mechanism.",
        "texts": [
            " When the link group GHK is acted upon the moving platform, it exerts a constraint on loop I, so the mobility of the mechanism is 3, but the moving platform has six motions, d3,6 I (x y z), which is similar to the planar 4-bar linkage. In the planar 4-bar linkage, the mobility of the 4-bar linkage is 1, but the connecting link has three motions including two translations and one rotation. In this example, dX I = dI, d X II = dII, there is no virtual constraint, so the mobility can be also obtained from eq. (1) or the \u201cK-G\u201d criterion. Example 2. Figure 5 is a 7-bar mechanism analyzed by Bagci [16]. In loop I, the four axes of the link group RCCR are parallel. Loop II is formed by adding the link group SRP at points E, F and G to links 2 and 7 of loop I. Loop III is formed by adding the link group to links 5 and 7 of loop II. All links in loop I cannot rotate about x-, y-axes, so dX I = dX I ( 0 0, x y z) = 4, FI = PI i=1fi dX I =64=2, d2,7 I ( 0 0, x y z)=4. In loop II, on account of the displacements caused by link 4, the rank of the link group EFG is dgz II (,x y z)=6, dX II =d2,7 I ( 0 0, x y z)+dgz II (, x y z)= dX II (, x y z)=6, FII = PII i=1 fi dX II =56=1"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003325_00207179.2012.658868-Figure4-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003325_00207179.2012.658868-Figure4-1.png",
        "caption": "Figure 4. Double pendulum with torque input.",
        "texts": [
            " Last, when using SeDuMi to solve the LMI problems obtained from this approach, additional constraints may be required to ensure solutions with satisfactory condition numbers. The authors behind this technique explore other convexification methods yielding standard LMI problems in Tuan and Apkarian (1999); in Tuan and Apkarian (2002), a relaxation method using monotonicity concepts is provided, where in this case the resulting finite family of LMI constraints is of polynomial order with respect to the number of parameters. We now consider the problem of controlling the double pendulum, shown in Figure 4, about an eventually periodic trajectory. In our setup, a vessel containing fluid is rigidly attached to the end of the second link of the pendulum. We will assume the vessel is designed in such a way that the oscillations of the fluid inside it are not significant and hence the fluid vessel system may be modelled as a point mass. The mass of the fluid, denoted m, is time-varying and it is assumed that this mass and its rate of variation, _m, are available for measurement at each discrete instant k",
            " With this said, we will assume that it is possible to estimate accurately the fluid mass acceleration \u20acm; for instance, it may be possible to measure _m at a faster rate than the sampling frequency, and then, based on a number of values of _m over the sampling interval, we may use some numerical approximation to calculate \u20acm\u00f0k\u00de rather accurately. This section is divided into three subsections: the first presents the nonlinear model of the double pendulum along with the reference trajectory; the second gives the corresponding NSLPV model; and the last formulates the synthesis problems and provides the simulation results. 4.1 Model and reference trajectory The pendulum, shown in Figure 4, consists of two rigid links connected by a hinge joint so that they can rotate with respect to each other. The first link is directly coupled to the shaft of a DC motor mounted to the end of a table. The resultant torque u applied to the first link is the only control input to this rigid body system. The states of the system are the angular displacements q1, q2 and angular velocities q3 \u00bc _q1, q4 \u00bc _q2. We define the state vector q\u00bc (q1, q2, q3, q4). The double pendulum can be regarded as a two link robot with a single actuator"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001823_j.rcim.2010.06.010-Figure3-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001823_j.rcim.2010.06.010-Figure3-1.png",
        "caption": "Fig. 3. 8-Axis finishing cell and the corresponding kinematic scheme.",
        "texts": [
            " This operation is followed by sanding or a shot-blasting, which guarantees the mechanical characteristics by surface treatment prestressing. These operations, currently performed manually (sawing, knock out and polishing) are long and difficult. Their automation, mainly for economic and safety reasons, required the definition of a robotized finishing cell doing machining and polishing. This cell is also referred to as \u2018\u2018machine\u2019\u2019 in this paper. This cell integrates an ABB IRB 940 Tricept robot with 6 axes (q1\u2013q6) moved by a linear axis (track q7) and associated with a rotary axis (retournor q8) carrying the part (Fig. 3). This architecture is kinematically redundant and contains 8 controlled axes [15], thus allowing the finishing of large-sized parts (1.2 3 m2). Because of the variability of the parts inherent in the foundry process (torsion, net curtain, etc), the cell is equipped with an optical 3D measurement system enabling the integrity of the parts (thickness) to be preserved by adjusting the process trajectory on the real surface profile. The constraints in the machining and polishing processes require on the one hand, a sufficient machine capability related to the ratio between part dimensions (surfaces to be treated of approximately 1 m2) and geometrical specifications (form speci- fication about 0",
            " With respect to the results obtained by the SOM method, we retained a quadratic interpolation function of the form: Kxx\u00f0y,z\u00de \u00bc ay2\u00febz2\u00fecyz\u00fedy\u00feez\u00fe f The six coefficients (a, b, c, d, e and f) are determined by a least squares method (Fig. 8). The measured rigidity values are approximately twice as low as the values obtained by modeling. These variations can be explained by the difficulty of modeling the complex bodies and the joints. 6.1.1.2. Wrist. Deformation measurements were performed for each wrist axis (q4, q5 and q6) (Fig. 3) for several loading cases. The experimental results reveal relatively linear behavior for stiffness in torsion. It is observed that rigidity in torsion of the last axis, associated with q6, is low compared with the two others. We thus seek to minimize the torque exerted on this axis by the cutting pressures. 6.1.2.1. Weight of the load. The structure is mainly subjected to efforts of inertia, the weight of the embarked load and cutting pressures on the tool. Since advance speed must remain constant, we make the assumption here of neglecting the effects of inertia on the structure"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002084_ac102113v-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002084_ac102113v-Figure1-1.png",
        "caption": "Figure 1. Diagram of the film electrode transmission thin-layer spectroelectrochemical cell. The thin layer (TL) is formed between the working electrode (WE) and the optical transmission window. The thickness of TL is determined by the spacer (SP). BDD, metal minigrid, or other thin film electrodes are supported by optically transparent substrate (Q). The thin layer is connected to the auxiliary electrode (AE) via electrolyte capillary (EC) between the optical window and outer cell body. Electric contact with WE is provided by a circular wire or foil lead outside of the cell cavity. Reference measurement is provided by a miniature Ag/AgCl electrode (RE) via cracked glass junction (GJ). Reference electrode capillary is also used as a sample loading port (LP), which is exhausted via a similar discharge port (DP) on the opposite side. Part of electrolyte capillary up to the diffusion limit (DL) from working electrode can be electrochemically active. The bottom diagram shows the schematic view from the side of the working electrode (cutout) along the probe beam path. The width of the EC is exaggerated for clarity.",
        "texts": [
            "38 Herein, we report on the extensive characterization of the spectral and electrochemical response of ferricyanide, cytochrome c (Cc), and its cyanide derivative (Cc-CN) in aqueous media using the new cell design along with several electrode materials. Using a potential step/spectrophotometric approach, we report redox- and liganddependent changes in the rates of electrochemical reactions and assess anaerobic performance. Thin-Layer Spectroelectrochemical Cell. The transmission TLE cell design (Figure 1) builds upon an original design by Mantele et al.7 but with significant modifications which greatly increased the responsiveness, predictability, and reliability of the cell. The optically transparent working electrode (20 mm \u00d7 20 mm overall, 10 mm diameter active area) is pressed against the front surface of the cell body (PEEK) though a circular Kapton polyimide film spacer (Figure 1). The inner diameter of the spacer (10 mm) is larger than that of the central through bore of the cell (9 mm) except for three small tabs that extended into the bore by \u223c1 mm. A transparent back window (9 mm diameter \u00d7 5 mm)sCaF2 (used here), BaF2, Si, and fused silica, etc.swas pressed against the electrode through the spacer tabs by a hollow nut on the back surface of the window. Thus, the electrochemical volume of the cell is defined by the narrow space between two flat surfaces separated by the spacer of the desired thickness (7.5-75 \u00b5m) and the i.d. of the spacer. This volume is sealed with one O-ring encircling the spacer (Figure 1, bottom) and another between the perimeter of the nonelectrochemical surface of the back window and the hollow nut (not shown). A circular contact of the smallest possible diameter (17 mm) is pressed against the working electrode outside of the front O-ring (WE lead in Figure 1). Approximately \u223c10 cm of fine Pt wire serves as the counter electrode. It is either located in a 1 mm \u00d7 1 mm groove in the cell body encircling the back window (shown in Figure 1, top) or coiled in a separate compartment that is connected to that groove through a PEEK frit (1.5 mm diameter). The latter configuration provides better separation between the sample and the counter electrode medium at the expense of a small increase in ohmic (25) Bernad, S.; Mantele, W. Anal. Biochem. 2006, 351, 214\u2013218. (26) Yildiz, A.; Kissinger, P. T.; Reilley, C. N. Anal. Chem. 1968, 40, 1018\u2013 1024. (27) Benken, W. v.; Kuwana, T. Anal. Chem. 1970, 42, 1114\u20131116. (28) DeAngelis, T. P.; Hurst, R",
            " Langmuir 2002, 18, 450\u2013457. (36) Haacke, G. Annu. Rev. Mater. Sci. 1977, 7, 73\u201393. (37) Haymond, S.; Zak, J. K.; Show, Y.; Butler, J. E.; Babcock, G. T.; Swain, G. M. Anal. Chim. Acta 2003, 500, 137\u2013144. (38) Dai, Y.; Proshlyakov, D. A.; Zak, J. K.; Swain, G. M. Anal. Chem. 2007, 79, 7526\u20137533. 543Analytical Chemistry, Vol. 83, No. 2, January 15, 2011 resistance. Continuity in the electric circuit between the electrochemical cavity and the counter electrode is established via a cylindrical capillary volume (EC in Figure 1; \u223c50 \u00b5m wide) filled with supporting electrolyte. The capillary is formed by the outer cylindrical surface of the back window and the inner surface of central bore. For aqueous measurements, the cell is equipped with a miniature standard Ag/AgCl reference electrode. The reference electrode is connected to the side of the cylindrical capillary between the working and counter electrodes via a cracked glass junction. Utilization of the capillary electrolyte space surrounding the optical transmission window for circuit continuity allowed one to minimize the dimensions of the thin layer in the plane of the electrode",
            " All potentials are reported vs a Ag/AgCl reference electrode. Background voltammetric scans obtained using a blank sample immediately before or after the measurements were subtracted from those of the sample. For chronoamperometric measurements, no background correction was performed. UV-vis spectra were recorded continuously at 0.3 s intervals using Hewlett-Packard 8453 diode array spectrophotometer (Agilent Technologies, Santa Clara, CA). Optical spectra were probed in the central area of the electrode (\u223c6 mm diameter) as shown in Figure 1B. The start of the optical and amperometric measurements was electronically synchronized. The apparent electron-transfer rates upon application of the potential step were determined by a nonlinear regression analysis of optical transient at characteristic wavelengths, fitting with mono- or biexponential functions using Igor Pro software (Wavemetrics, Inc., Lake Oswego, OR, USA). For aerobic measurements, the analyte solution is loaded by placing a small volume of sample (5-10 \u00b5L) in the center of the working electrode immediately prior to the assembly of the cell otherwise filled with the supporting electrolyte",
            " Thus, the A-t curves were converted into derivative form for direct comparison with simultaneously acquired amperometric data. In the 0.1 M supporting electrolyte, the maximal apparent rate of electrolysis of Fe(CN)6 3-/4- was observed at 1 and 4 s in the \u2206A/\u2206t curves for the 1 and 5 mM concentrations, respectively. With 1 M supporting electrolyte most of the electrolysis occurred within instrument sampling time (0.3 s). In both cases, the maximal rate of optical changes probed in the central part of the electrode (as shown in Figure 1) correlated with the slowest amperometric phase. The initial, optically silent ampero- metric phase is attributed to the charging of the double layer and electrolysis in the peripheral, dark areas of the cavity. Optical changes associated with the reduction of Cc during the double potential step are shown in Figure 4A. Since the cell configuration and conditions for the protein measurement were the same as in Figure 3, the ohmic resistance was similar in both cases. The apparent rate of reduction and oxidation of Cc was determined from the temporal changes between peak and trough in the difference absorption spectrum",
            " The maximal apparent rate of a redox reaction can be limited by three factors: analyte diffusion, the overall conductivity of the cell, and the intrinsic heterogeneous electron-transfer rate constant. The effective ohmic resistance at the optically active center of the layer is determined by the geometry of the electrode and resistivities of both the electrode film and the solution above. The relative contribution of the film electrode to the overall ohmic resistance increases with increasing thickness of the sample layer (TL, Figure 1) as the amount of analyte increases and the resistance of the thin layer decreases. On the other hand, the thickness of the analyte layer determines the maximal diffusion-limited reaction rate and the spectral path length (water absorption in IR, minimal detectable analyte concentration, etc.). Our results show that the rate of optical changes for Fe(CN)6 4- on Au-MG tracks the final amperometric phase (Figure 3) and scales with the electrolyte concentration. Since ohmic resistance of Au-MG is negligible and the intrinsic rate of electron transfer on the electrode is constant between experiments, changes in the apparent rates seen in Figure 3 represent the maximal attainable redox rates as limited by the ohmic resistance of the electrolyte"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002811_1754337112441112-Figure5-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002811_1754337112441112-Figure5-1.png",
        "caption": "Figure 5. FEA model of an aluminum striker bar (white) colliding into a foam sample (black) using quarter symmetry.",
        "texts": [
            " The following describes an FEA foam model with an experimentally derived material loading response and a phenomenologically developed unloading response. Models developed in this way have the potential of describing and comparing structural response prior to product fabrication. A collision of the aluminum striker bar with a PU foam sample was modeled using the dynamic finite element code LS-DYNA (Version 971, LSTC, Livermore, CA). Because the foam cell size (;0.2mm) was significantly smaller than the specimen diameter, the foam was assumed to be isotropic and homogeneous. The foam sample, shaded black in Figure 5, was modeled with 2772 linear, solid elements with two symmetry planes. Large strain magnitudes in the model necessitated the use of fully integrated elements to minimize hourglassing. The elastic aluminum striker rod, shaded white in at Virginia Tech on November 11, 2014pip.sagepub.comDownloaded from Figure 5, was similarly modeled with 2709 elements. The rod length was shortened to 2.5mm to reduce the number of elements, while density was correspondingly increased to achieve the correct mass. The striker bar velocity was directed normal to the impact surface and the specimen face opposite of the impact was constrained in-plane. The contact type was \u2018\u2018surface to surface\u2019\u2019 where friction was neglected, as the impact had no obliquity. The simulation duration was 2ms. A standard foam material model was selected to characterize the PU foam (Mat #57 low density foam) as used by Sambamoorthy and Halder"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002057_20090819-3-pl-3002.00080-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002057_20090819-3-pl-3002.00080-Figure2-1.png",
        "caption": "Fig. 2. Hydraulic pump.",
        "texts": [],
        "surrounding_texts": [
            "As for modelling, the considered mechatronic system is divided into a hydraulic subsystem and a mechanical subsystem, which are coupled by the torque generated by the hydraulic motor (Schulte [2007]). Here, the actuator dynamics of the displacements units is taken into account by a first order lag element, i.e., a PT1-system. Consequently, the dynamics of displacement unit of the pump, which is realised as a tiltable swashplate, is governed by TuP \u02d9\u0303\u03b1P + \u03b1\u0303P = kP uP, (1) whereas the dynamics of the motor according to a bend axis design is described by TuM \u02d9\u0303\u03b1M + \u03b1\u0303M = kM uM. (2) Here, a normalised swashplate angle \u03b1\u0303P = \u03b1P/\u03b1maxP and a normalised bend axis angle \u03b1\u0303M = \u03b1M/\u03b1maxM have been introduced. The volume flow rate of the pump can be stated as a nonlinear function of the swashplate angle qP = VP(\u03b1P)nP = V\u0303P(\u03b1P)\u03c9P, (3) with V\u0303P = VP/2\u03c0 . Nevertheless, the reasonable assumption of a small swashplate angle |\u03b1P| \u2264 18\u25e6 , c.f. Kugi et al. [2000], allows for the approximation qP = V\u0303P\u03b1\u0303P\u03c9P, (4) with \u03b1\u0303P \u2208 {\u22121,1}. Accordingly, the volume flow rate into the hydraulic motor can be formulated as a nonlinear function of the bend axis angle qM = VM(\u03b1M)nM = V\u0303M(\u03b1M)\u03c9M, (5) with V\u0303M = VM/2\u03c0 . As before, the assumption of small angles |\u03b1P| \u2264 20\u25e6 leads to a simplified relationship qM = V\u0303M\u03b1\u0303M\u03c9M, (6) with \u03b1\u0303M \u2208 {0,1}. The pressure dynamics of the hydrostatic transmission in closed-circuit configuration can be modelled by two capacitances (Jelali and Kroll [2004], Kugi et al. [2000]), which account for both the fluid compressibility and the elasticity of the connecting hoses qCA = CA p\u0307A, qCB = CB p\u0307B. (7) The internal leakage oil flow is modelled as a laminar flow resistance (Jelali and Kroll [2004]), which depends linearly on the difference pressure qleak = kleak(pA\u2212 pB) (8) characterized by the leakage coefficient kleak. The external leakage could be introduced as well but shall be neglected in the control-oriented design model. A volume flow balance leads directly to the resultant volume flows in the two capacitances qP\u2212qM\u2212qleak = qCA , (9) \u2212qP +qM +qleak = qCB . (10) A model order reduction becomes possible, if some reasonable symmetry assumptions are made. Considering identical capacitances CA = CB =: CH leads to an order reduction by one and results in a differential equation for the difference pressure \u2206p\u0307 = 2 CH ( V\u0303P \u03b1\u0303P\u03c9P\u2212V\u0303M \u03b1\u0303M\u03c9M\u2212 kleak\u2206p ) . (11) The longitudinal dynamics of the working machine is governed by the equation of motion. The vehicle with the drive chain system (vehicle mass mv, wheel radius rw, gear box transmission ratio ig, rear axle transmission ratio ia, damping coefficient dvc at the drive shaft, see also Fig. 4) can be described by the following second order differential equation mv r2 w i2a i2g\ufe38 \ufe37\ufe37 \ufe38 JV \u03c9\u0307M + dg i2g\ufe38\ufe37\ufe37\ufe38 dvc \u03c9M = V\u0303M\u2206p \u03b7M\u03b1\u0303M\ufe38 \ufe37\ufe37 \ufe38 \u03c4M \u2212 \u03b7g\u03c4L iaig\ufe38 \ufe37\ufe37 \ufe38 \u03c4U . (12) Here, the torque \u03c4M = V\u0303M\u2206p\u03b7M\u03b1\u0303M of the hydraulic motor depends on its mechanical-hydraulic efficiency \u03b7M . The overall system model involves four first order differential equations. Introducing the normalized angles of the displacement units \u03b1\u0303i, i \u2208 {P,M}, the difference pressure \u2206p, and the motor angular velocity \u03c9M as state variables, the corresponding state-space representation becomes  \u02d9\u0303\u03b1P \u02d9\u0303\u03b1M \u2206p\u0307 \u03c9\u0307M =  \u2212 1 TuP \u03b1\u0303P + kP TuP uP \u2212 1 TuM \u03b1\u0303M + kM TuM uM \u2212 2kleak CH \u2206p+ 2V\u0303P CH \u03c9P\u03b1\u0303P\u2212 2V\u0303M CH \u03c9M\u03b1\u0303M \u2212 dVC JV \u03c9M + V\u0303M\u03b7M JV \u2206p\u03b1\u0303M\u2212 \u03c4U JV  . (13) The input voltages ui, i \u2208 {P,M}, of the corresponding proportional valves for the displacement units serve as physical control inputs. Feedfoward compensation of the actuator dynamics using its approximated inverse in form of a proper transfer function, as seen in Fig. 5, leads to a simplified control design with u\u0304i \u2248 \u03b1\u0303i, where u\u0304i represents the inverse input. Therefore, only the differential equations for the difference pressure and the motor angular velocity are used for feedback control design. The remaining model uncertainties are taken into account by the disturbance torque \u03c4U . On the one hand, these uncertainties stem from external load forces and drive resistances acting on the vehicle. On the other hand, a varying vehicle mass mv contributes to parameter uncertainty."
        ]
    },
    {
        "image_filename": "designv11_12_0001960_s12239-009-0053-x-Figure3-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001960_s12239-009-0053-x-Figure3-1.png",
        "caption": "Figure 3. Multi-body analysis model of a manual transmission.",
        "texts": [
            " A realistic multibody dynamic model for a transmission system, which may reflect real working conditions, should be constructed for an accurate loading analysis. An FF manual transmission consists of a clutch, input and main shafts, mating helical gears, final-drive gears in the differential section, and housing. The 3-D six-speed manual transaxle model, which combines the manual transmission, final drive gearing, and differential into a single unit, is shown in Figure 2. A multibody dynamic analysis model for Figure 2 was constructed using MSC/ADAMS and is represented in Figure 3. This model is based on the following three assumptions: (1) shafts and gear teeth are flexible, and bearings are considered to be bushings with 6 degrees of freedom; (2) gear meshing stiffness due to bending varies along a moving contact point between two helical teeth; and (3) fluctuating torque or acceleration is transmitted through the clutch input to the transmission input shaft. 2.1. Bending Stiffness of a Gear Tooth Figure 4 shows a schematic of three components of forces acting against a helical gear tooth"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002420_s10237-009-0187-9-Figure3-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002420_s10237-009-0187-9-Figure3-1.png",
        "caption": "Fig. 3 Maps between the reference, intermediate and deformed configurations for each bar element",
        "texts": [
            "1 Bar kinematics The bar elements employed in our model are defined by the following kinematic assumptions: \u2022 Each bar is a body much longer in one direction that in the other 2 perpendicular directions. \u2022 The bar remains as a straight body in all the configurations, with a constant area. \u2022 The cross-section of the bar remains perpendicular to the centroid axis and has a constant area. \u2022 The bar in the reference configuration is oriented in such a way that its long axis is parallel to the vector E1 of a reference triad (see Fig. 3). \u2022 The active deformations deform the bar only in the longitudinal direction, that is in the direction of E1. Also, and in agreement with our hypothesis in the threedimensional case, no density changes occur during this active deformation. \u2022 The elastic deformations correspond to a change in the longitudinal along E1 axis plus a rotation R, constant for each bar. Such a motion of the bar is visualised with the maps indicated in Fig. 3. The positions of the material point in each configuration are accordingly given by, X0 = X0 i Ei X = Xi Ei x = xi ei = xi REi where R is a rotation matrix that transforms the vector Ei into ei , i.e. R = ei \u2297 Ei and ei = REi . We remark that in the previous equations we do not constrain the material points to be only along the axis E1 (the bar is assumed slender, but the area of the cross-section is not zero). From the kinematic assumptions, the three configurations are related through the following relations: X = X0 + ua(X 0 1)E1 = ua(X 0 1)E1 + X0 i Ei x = R(X + ue(X1)E1) = ue(X1)e1 + Xi ei (27) = (ua(X 0 1)+ ue(X1))e1 + X0 i ei , where summation on the repeated subscript i = 1, 2, 3 must be understood"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002522_mesa.2012.6275544-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002522_mesa.2012.6275544-Figure1-1.png",
        "caption": "Figure 1. The model parameters for a general planar robot.",
        "texts": [
            " Since these parameters directly affect the stability of the system, a modified FFSM (MFFSM) is developed to make the margin sensitive to these parameters as follows: = \u2219 ( ) \u210e \u2219  \u0305 (2) where = 1 if the projection of the CG is inside or on the support polygon and = \u22121 if it is outside of the support polygon. For analysis of (2) and finding out all the parameters defined in the equation, a planar robot is considered first and then is generalized to the spatial case as well as to the mobile wheeled robots. Assume that there is a planar robot over an uneven terrain as shown in Fig. 1. In this figure, is the position vector of th ground contact point, is the position vector of CG, = \u2212 is ith tipover vector and = \u2016 \u2016 is its length, \u210e is height of CG with respect to th ground contact point in the gravity direction, is the net moment vector acting at the center of gravity, and ! is the net force vector acting at the CG which can be derived from ! = \u2212 \u2211 ! where ! is th foot force vector whose normal component is \" with magnitude of = \u2016\" \u2016. Hence, \u0305 = \u2211 , known as normal foot force magnitude, directly correlates with the amount of the net force and could be multiplied by FFSM for top-heaviness sensitivity [6]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003880_apcase.2015.44-Figure8-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003880_apcase.2015.44-Figure8-1.png",
        "caption": "Fig. 8. Overview of each sensor",
        "texts": [
            " Also, when decrease of the angle of the caregiver\u2019s upper half body and caregiver\u2019s contact to the carereceiver are detected simultaneously, air release from the artificial muscle is executed automatically. Thus, the caregiver obtains the assist force from the assist suit during actions 4) to 6). Therefore, the burden to the caregiver\u2019s waist during actions 4) to 6) can be reduced. The mounting position of the tilt sensor and the pressure sensor is shown in Fig. 7. The tilt sensor is mounted on the caregiver\u2019s back, and the pressure sensor is mounted on the back of the caregiver\u2019s right hand. Overview of each sensor is shown in Fig. 8. When the caregiver performs \u201c4) Lifting\u201d, he/she holds the care-receiver by both hands with putting the palm of the left hand on the back of the right hand. At this moment, the pressure sensor reacts. Therefore, the caregiver\u2019s contact to the care-receiver can be detected by the pressure sensor. In this study, RAS-2C produced by Kondo Kagaku co.,ltd. is used as a tilt sensor. The principle of the tilt sensor is illustrated in Fig. 9. As shown in Fig. 9, the tilt sensor is always outputting the reference voltage to the direction of gravitational acceleration"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001196_0094-114x(72)90004-3-Figure5-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001196_0094-114x(72)90004-3-Figure5-1.png",
        "caption": "Figure 5. Internal angles of a spherical four-bar linkage.",
        "texts": [
            " A mechanism which requires a revolute joint on axis 3 for two positions of the input and allows the same revolute joint to exist in all other positions is represented as RCP,*C. 5. Conditions for Overconstraint in the RCCC Chain A direct position analysis of the RCCC linkage can be obtained by using spherical t r igonometry to find the rotations and by using the projection equations of the matrix analysis to find the translations. The internal joint angles, ~,, are related to the rotations, 0,, as $ i = ~ - - 1 8 0 \u00b0. (3) Figure 5 shows the two phases of a spherical four-bar linkage, with the internal joint angles and some auxilliary quantities which decompose the four-sided figures into sets of triangles. The angle/3 is the diagonal of the linkage. Each angle X~ is the internal angle of a triangle which is opposi te to the side a, of the same triangle. Finally, the subscripts 5 -7 for the internal joint angles refer to the second phase of the linkage. The analysis is then sin a4 sin ~bt/sin/3 tan X4 = (cos a~ sin m - sin a4 cos a~ cos 6~)/sin/3 (4) with X4 ~ 180\u00b0 cos/3 = cos a4 cos cq + sin a4 sin og I COS ~)1 (5) sin/3 = cos oq sin oq - sin c~4 cos cq cos qb~ c o s X4 (6) COS/3 - - COS O~ 2 COS ~ 3 cos ~b3 = sin a2 sin a3 (7) with ~b3 ~ 180 \u00b0 ~b6 = 360 \u00b0 - ~b3"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002335_detc2010-28926-Figure17-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002335_detc2010-28926-Figure17-1.png",
        "caption": "Figure 17. Example of a fundamental extension in 3D. a) The 3D dyad. b) The 3D triad.",
        "texts": [],
        "surrounding_texts": [
            "Once we have all the AGs in 2D it is possible to derive all the topological data of all the 2D mechanisms, by composing different AGs under the condition that the following composition rule is satisfied: The composition rule of AGs: Let G and T be two AGs. G can be composed on T if: a. Any one of the ground vertices of G is connected to an inner vertex of T or to the ground. b. The number of vertices that G is connected to is greater than or equal to two. Figure 15 depicts several examples of determinate trusses that are compositions of one triad and one dyad. Once we have all the compositions of Assur Graphs, i.e., various determinate trusses, the process of obtaining various linkages from them is done easily by connecting a driving link to one of the ground vertices. For example, in Figure 16 for a determinate truss (16a), which consists of a composition of a tetrad on a triad, there are four corresponding linkages since there are four ground vertices. B A C 1 2 4 5 6 3 B A C 1 2 4 5 6 3 (a) (b) 1 2 3 4 6 5 8 9 10 7 11 12 1 3 4 6 5 8 9 10 11 12 (a) (b) 9 Copyright \u00a9 2010 by ASME Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 08/12/2018 Terms of Use: http://www.asme.org/about-asme/terms-of-use 10 Copyright \u00a9 2010 by ASME The idea of the classes, and that each fundamental Assur Graph is the representative of a class of Assur Graphs is the same as in 2D, only this time the operation of transforming one Assur Graph into the successor is done through 2-extension. In the 2- extension the vertex that is added is connected this time to two other vertices and not to one as is done in the 1-extension. Example of deriving an Assur Graph from 3D triad (Figure 18a) appears in Figure 18b,c where the edges being split are indicated by the bold edges. a) The fundamental Assur Graphs spatial triad. b,c) The resulting Assur Graph after applying the 2-extension. The structure of the 3D map is the same as for the 2D Assur Graph. The first column consists of the spatial dyad, which also does not have a correspondence in the 3D Assur Graph, thus there are no derivations in that column. The first row consists of the fundamental Assur Graphs, and each column contains all the derivations from that representative using 2- extension operation. In contrast to 2D, there is no mathematical proof for the completeness of the 3D Assur Graphs. The main reason for that is that there are still mathematical problems that have not yet been resolved by the mathematicians in the rigidity theory community. Among these is the Assur Graph, appearing in Figure 19, for which there is no derivation from any fundamental Assur Graphs. The main problem is that the degree of each vertex in this graph is at least five. It is expected that some of the problems that engineers and mathematicians encounter in 3D are to (b) (c) (d) (e) 10 Copyright \u00a9 2010 by ASME Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 08/12/2018 Terms of Use: http://www.asme.org/about-asme/terms-of-use 11 Copyright \u00a9 2010 by ASME be solved by using Assur Graphs. For example, there are known structures and graphs for which all the formulas for calculating the degrees of freedoms, such as: Grubler equation or Laman's theorems provide wrong answers. However, when we decompose the graphs into Assur Graphs we reveal the reason for the problem and might be able to overcome it. For example, in Figure 20a appears a floating structure, a structure with no grounding, and although its DOF is equal to six, i.e., should be a rigid body, it has a finite motion since the upper and the lower parts can rotate related to each other along the virtual axis \u2013 FD. But, when we first ground the structure by pinning joints E,H and G (Figure 20b) the resultant structure is decomposed into a triad and two dyads (Figures 20c,d). Now that we have the decomposition and the building blocks, the AGs, the problem is revealed. The triad is connected to the other structure, the two dyads, by only two joints. It is easy to prove (Shai, 2008) that in any dimension d, the number of joints that any AG should be connected to the remaining graph should be at least d."
        ]
    },
    {
        "image_filename": "designv11_12_0001540_s11431-010-4176-0-Figure5-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001540_s11431-010-4176-0-Figure5-1.png",
        "caption": "Figure 5 The structure of the pump. 1, The vibrator; 2, the cover of the pump; 3, the chamber of the pump; 4, the clamp; 5, the base of the pump.",
        "texts": [],
        "surrounding_texts": [
            "+x direction with a speed of U. The resultant force of thrust T is along the +x direction, generated by the vibrator, and the interacting force between the vibrator and the fluid is F(x, t).\nTo simplify the problem, the assumptions are made as follows.\n1) The fluid is inviscid, incompressible and in steady state. Only the average values of the flow are taken into account.\n2) When the vibrator is working, the surrounding fluid is continuous with no gap between the vibrator and the fluid.\n3) Heat transfer between the vibrator and the fluid is neglected.\n4) The coordination system is assumed to be rest with the fluid in order to simplify the deduction of the output power of the vibrator (eqs. (1) to (5)). If the vibrator is free in the head, the thrust power will drive it along the \u2212x direction. However, the head of the vibrator is clamped. Therefore, the thrust power will force the fluid to flow along +x direction (eq. (6)).\nThe velocity of the vibrator is expressed as\n( , ) ( , )( , ) ,h x t h x tv x t U t x \u2202 \u2202 = + \u2202 \u2202 (1)\nwhere h(x, t) is the amplitude for the vibrator, its general solution is\n( , ) ( )sin(2 ) sin(2 ) / ), h x t a x Ut f t x Ut \u03bb = + \u03c0 \u22c5 \u00d7 \u03c0( +\n(2)\nwhere a is the prefactor of the first order polynomial, f is the vibration frequency and \u03bb is the equivalent wave length. The first vibration mode is a special solution when \u03bb\u03c0 + =sin[2 ( ) / ] 1.x Ut According to the Newton\u2019s second law, the interactive force F(x, t) between the fluid and the vibrator is\nd( , ) { ( , ) ( )} d\n{ ( , ) ( )},\nF x t V x t m x t\nU V x t m x t x\n=\n\u2202 \u2202\u239b \u239e= +\u239c \u239f\u2202 \u2202\u239d \u23a0\n(3)\nwhere m(x) is virtual mass of the fluid that interacts with the vibrator at position x.\nThe average power of the vibrator can be derived by utilizing the interaction force between the fluid and the vibrator:\n0\n( , ) ( , )d\n( ) | ,\n( , ) ( , ) ( , ) .\nl\nx l\nx l\nh x tP F x t x t\nUm x B\nh x t h x t h x tB U t t x\n=\n=\n\u23a7 \u2202\u23aa =\u23aa \u2202 \u23aa\u23aa =\u23a8 \u23aa \u23a7 \u23ab\u23aa \u2202 \u2202 \u2202\u23a1 \u23a4= +\u23a8 \u23ac\u23aa \u23a2 \u23a5\u2202 \u2202 \u2202\u23a3 \u23a6\u23a9 \u23ad\u23aa\u23a9 \u222b (4)\nBy energy conservation, the power that drives the fluid can be obtained by subtracting the energy consumed by fluid flow from the power generated by the vibrator. For the purpose of simplification, the energy consumed by fluid motion is assumed to occur only at the end of the vibrator (x=l). Thus, the average power of creating flow for the fluid is\n2\n2 2\n2\n1 ( ) ( , ) 2\n1 ( , ) ( , )( ) . 2\nx l\nTU P Um l V l t\nh x t h x tUm l U t x\n=\n= \u2212\n\u23a7 \u23ab\u2202 \u2202\u23aa \u23aa\u239b \u239e \u239b \u239e= \u2212\u23a8 \u23ac\u239c \u239f \u239c \u239f\u2202 \u2202\u239d \u23a0 \u239d \u23a0\u23aa \u23aa\u23a9 \u23ad\n(5)\nIn eq. (5), m(l) is the virtual mass of fluid, which interacts with the vibrator at the position x=l. Again, by energy conservation, we can get\n21 ( ) . 2 TU Um l U= (6)\nSubstituting eq. (6) into eq. (5), we obtained\n22\n2 ( , )( , ) .1 x l x l h x th x t U tt = = \u23a1 \u23a4\u2202\u239b \u239e\u2202 = \u23a2 \u23a5+ \u239c \u239f\u23a2 \u23a5\u2202\u2202 \u239d \u23a0\u23a3 \u23a6 (7)\nAccording to the vibration theory, the relationship between the parameter a and the driving frequency f in eq. (2) can be expressed as\n2 2\n0 0\n,\n1 2 ba f fc f f = \u239b \u239e \u239b \u239e \u2212 +\u239c \u239f \u239c \u239f\n\u239d \u23a0 \u239d \u23a0\n(8)",
            "where f0 is the resonant frequency, c is the damping ratio and b is a constant. Parameters b and c are determined by the vibration experiment.\nFor the first vibration mode: c=0.060; b=1/500. For the second vibration mode: c=0.015; b=1/7000. From eqs. (7) and (8), the fluid velocity in the interacting region around the vibrator can be solved. Then according to the mass conservation principle, the average velocity in pump chamber can be obtained. Finally, the flow rates can be calculated by multiplying the velocity with the area of the stream tube. The relation between the flow rates and the driving frequencies under the second vibration mode is too complicated to get the exact solution. However, by referring to the analysis for the first vibration mode, the part of the vibrator from the clamped end to the second pitch line can be ignored. Thus, the second vibration mode can be treated as the first vibration mode so that the relationship between the flow rate and the driving frequencies can be solved approximately. Both the relations for the first and the second vibration modes are given as follows.\nQi=q( f ), (9)\nwhere i=1,2 represent the first and the second vibration modes, respectively. The dynamic results between the flow rates and the driving frequencies for the first and the second vibration modes are shown in Figure 3.\nFigures 4 and 5 show the structures for the piezoelectric vibrator and the fishtailing type of piezoelectric pump, respectively. Figure 6 is the photograph of the vibrator and the pump. Tables 1 and 2 show some critical electronic and geometric parameters, respectively.\nV1 is the driving voltage (all the voltages are the peak values) in the vibration experiments, V2 is the driving voltage in the flow rates experiments, D33 is the piezoelectric constant and C is the capacitance. Parameters shown in Table 2 were also given in Figures 4 and 5.",
            "Figures 7(a)\u2013(d) show the FEM results of the four phases in a working period for the first vibration mode when it is immersed in water with an AC voltage of 100 V. Point A is the pitch point of the first vibration mode. During phase (a), the vibrator begins to move upward and leaves its equilibrium position. The vibrator reaches its positive maximum amplitude of 0.44 mm during phase (b). In phase (c), the vibrator moves downward to its equilibrium position. In phase (d), the vibrator reaches it negative maximum amplitude of \u22120.44 mm.\nFigure 8 shows the first vibration mode attained in the experiment, in which the vibrator is driven by the same voltage in water. Similarly, the whole process involves four parts: (a) the vibrator is leaving its equilibrium position; (b) it reaches its positive maximum amplitude of 0.5 mm; (c) it moves downward to its equilibrium position; (d) it moves downward continuously to the negative maximum amplitude of \u22120.5 mm. Finally, it moves back to its initial position (a).\nBy comparing the results of simulations and experiments,\nit is not hard to find that the vibrator behaves similar in both cases, and their difference in the values of the maximum amplitude is about 12%.\nFigure 9 shows the second vibration mode in FEM simulation. The vibrator was modeled as being immersed in water and driven by a 100 V\u2019s alternating voltage. Point A and point C are pitch points of the vibration. In (a), point B moves downward while point D moves upward. In (b), the point B reaches its negative maximum amplitude while point D reaches its positive maximum amplitude of about 0.05 mm. In the next phase, all points move toward their equilibrium positions as shown in part (c). In the final phase (d), point B is moving upward while point D is moving downward, then point B reaches its positive maximum amplitude and point D reaches its negative maximum of about \u22120.05 mm.\nFigure 10 presents the second vibration mode results, which were obtained from experiments under the same physical conditions as for FEM simulation. Points A and C are the pitch points for the second vibration mode. In (a), all points are leaving their initial positions. In (b), point B reaches the negative maximum position while point D"
        ]
    },
    {
        "image_filename": "designv11_12_0000193_icma.2005.1626722-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000193_icma.2005.1626722-Figure1-1.png",
        "caption": "Fig. 1 Laser cladding process",
        "texts": [
            "-7803-9044-X/05/$20.00 \u00a9 2005 IEEE Index Terms \u2013 Laser cladding, Trinocular CCD-based optical detector. I. INTRODUCTION Laser cladding is an advanced laser materials processing technique, which has been used in manufacturing, part repair, metallic rapid prototyping, and coating for a decade [1]. In this process, a laser beam melts powder particles and a thin layer of the moving substrate to form a layer (clad) on the substrate with a thickness ranging from 0.1 to 2 mm as shown in Figure 1. For part fabrication or prototyping using this technique, similar to other layered fabrication techniques, a threedimensional CAD solid model is used to produce a part without intermediate steps. This approach to produce a mechanical component in a layer-by-layer fashion allows industries to fabricate a part with features that may be unique to laser powder deposition technique. These features include a homogeneous structure, enhanced mechanical properties, and production of complex geometries. In general, an automated laser cladding technique, as a complex mechatronics system, is the blending of five common technologies: laser, computer-aided design (CAD), robotics, sensors/control, and powder metallurgy as shown in Figure 2"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000353_038-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000353_038-Figure1-1.png",
        "caption": "Figure 1. Geometry of the problem.",
        "texts": [
            " The last boundary condition for the director field , that is, the vanishing of the 2 implies the vanishing of the shear strain component \u03b513 from equation (14) and which in turn from equation (9) implies the vanishing of the shear stress component \u03c313, which is one of the boundary conditions associated with displacement field. Therefore, the appropriate boundary conditions at the free surface x3 = 0 will be the vanishing of the normal and shear components of stress tensor in the x-z plane; that is, \u03c333 = \u03c313 = 0, at x3 = 0, (20) where \u03c333 = (1 + \u03b9\u03c9\u03c4R)(c13u1,1 + c33u3,3), (21) \u03c313 = (1 + \u03b9\u03c9\u03c4R)cR 44(\u03c9)(u1,3 + u3,1). (22) For an incident (qP or qSV) wave at free surface x3 = 0, the qP and qSV waves will get reflected in the x1\u2013x3 plane as shown in figure 1. Accordingly, if the wave normal of the incident wave makes an angle \u03b80 in the negative direction of the x3-axis, then those of the reflected qP and qSV waves make angles \u03b81 and \u03b82 in the same direction. To satisfy the boundary conditions (20) , the following components of the displacement vector u are considered: u1 (\u03b2)(x1, x3, t) = U(\u03b2)d1 (\u03b2)e\u03b9(\u03c9t\u2212k1 (\u03b2)x1\u2212k3 (\u03b2)x3), (23) u3 (\u03b2)(x1, x3, t) = U(\u03b2)d3 (\u03b2)e\u03b9(\u03c9t\u2212k1 (\u03b2)x1\u2212k3 (\u03b2)x3), (24) where \u03b2 = 0 corresponds to incident waves and \u03b2 = 1, 2 correspond to reflected waves"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001835_s12239-010-0044-y-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001835_s12239-010-0044-y-Figure2-1.png",
        "caption": "Figure 2. Schematic diagram of a constant velocity joint.",
        "texts": [
            " However, these joints are designed to first address the shudder of the vehicle; thus, the emerging issue of idle vibration characteristics was not carefully studied in previous designs. CV joints have to control the compensation of height, when there is a difference in the center position between the power-train and the front wheels. They also have to manage the situation where the working length of the drive shaft changes in the axial direction in response to the combination of steering-angle changes and road conditions. The components of the CV joint are shown in Figure 2. It consists of three spiders or trunnions that spread out at intervals of 120o and are perpendicular to the intermediate shaft. The rollers are attached to the end of each spider. Depending on the joint angles, the rollers at the end of the spiders move to the radial direction of the tulip and they also move to the axial direction with simultaneous rotation in the inside of the tulip when there is a length change in the driveshaft. The CV joint connected to the knuckle mainly compensates the rotation angles, and it is called a Birfield joint (BJ) or Rzeppa joint"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002472_icelmach.2012.6350085-Figure9-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002472_icelmach.2012.6350085-Figure9-1.png",
        "caption": "Fig. 9. Permeability distributions (fc = 2 kHz).",
        "texts": [
            " Variation in Harmonic Fields with Load To understand the above-mentioned observations, the variation in the harmonic magnetic field with switching frequency and load conditions are investigated. Fig. 8 shows the distribution of harmonic flux densities decomposed by applying Fourier transformation to the result of 2-D FEM under the no-load condition. The flux in the rotor of IM can be separated into the following [8]: a) Fundamental rotational field b) Stator slot harmonics c) Inverter carrier harmonics. The figure indicates that the harmonic fields are concentrated at the rotor surface, particularly near the top of the bars, whereas the fundamental field deeply enters into the core. Fig. 9 shows the variation in the permeability of the core with load. It is observed that the saturated area at the top of the bars increases with the load. This is attributed to the increase in the fundamental secondary currents in the bars with the load. This result implies that the rotor slots approaches to open slots under load conditions and that the harmonic fluxes enter into the rotor bars. This must be the reason why the harmonic secondary cage losses increase with load in Fig. 7. Note that the harmonic secondary cage losses rapidly decrease with the switching frequency because the skin effect is greater than that in the electrical steel sheets due to the dimension of rotor bars",
            " 11 shows the loss-density distributions and fluxdensity waveforms near top of a bar. The variation in the waveforms along the thickness direction of the electrical steel sheet is estimated by the 1-D FEM. These waveforms include both the stator slot harmonics and the inverter carrier harmonics. It is observed that the harmonic ripples are concentrated at the surface of the electrical steel sheet because of the skin effect. However, in the case of the load condition, the waveform becomes nearly rectangular because of the magnetic saturation, as shown in Fig. 9. Fig. 12 shows the distributions of harmonic flux densities along the thickness direction of the electrical steel sheet. These results are obtained by applying Fourier transformation to the waveforms in Fig. 11. It is observed that both the slot harmonic and inverter carrier harmonic are concentrated at the surface of the electrical steel sheet, particularly the inverter carrier harmonic. On the other hand, the distribution of the slot harmonic is nearly independent on the switching frequency of the inverter, whereas the skin effect of the inverter carrier harmonic increases with an increase in the switching frequency. The figure also indicates that the skin effect weakens when the motor is loaded. This is attributed to a decrease in the permeability near the top of the bars, as shown in Fig. 9. From the above-mentioned results, it is concluded that the increase in the saturated area of the rotor core with load causes considerable effect on the variation in the additional motor loss components, as follows: a) The ratio of harmonic primary copper and secondary cage losses increases, whereas the ratio of harmonic core loss decreases. b) The dependence of harmonic core losses on the switching frequency decreases because the skin effect in the electrical steel sheet used for the core weakens"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002133_016173461003200205-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002133_016173461003200205-Figure2-1.png",
        "caption": "FIG. 2 Ro bot, with sche matic (a) and ex per i men tal setup (b).",
        "texts": [
            " The ro bot arm is pri - mar ily in tended for mount ing on a wheel chair or bed to be con trolled by phys i callyhand i capped per sons us ing a key pad or joy stick, but can also ac cept in put com mands and co - or di nates from a com puter in a sep a rate soft ware mode. Ro bot co or di nates are read and writ - ten in ar rays of six val ues, three of which de scribe Car te sian co or di nates and three de scribe the ori en ta tion of the ro bot grip per hand in quan ti ta tive mea sure ments of \u2018yaw\u2019, \u2018pitch\u2019 and \u2018roll.\u2019 Figure 2a il lus trates the six joints of the ro bot num bered for con ve nience. The sev enth de gree of free dom cor re sponds to the grip ping abil ity of the ro bot. The ro bot arm may be folded into a com pact form or un folded as re quired, has a reach of 80 cm and each of the six joints may ro tate a full 360\u00b0, ex cept joint 5 which has a 120\u00b0 arc of move ment. The ro bot can be in structed to move ei ther in di vid ual or mul ti ple joints at spec i fied an gu lar ve loc i ties or it at UCSF LIBRARY & CKM on March 11, 2015uix",
            " The trans ducer was then held in the ro bot grip per and bound tightly with ca ble ties to en sure the trans ducer and its nee dle did not shift dur ing the bi opsy, sim u lat ing the firm grip of a hu man op er a tor. The cen ter axis of the trans ducer was aligned to run di rectly through the ro ta tional axis of the grip per such that ro ta tion of that joint would re sult in the nee dle tip trac ing out a cir cle of di - am e ter 14 mm cen tered on that same axis. It was also im por tant to ini tially align the trans - ducer face par al lel to the x-y plane of the ta ble. A stan dard ized un folded po si tion and ori en ta tion of the grip per, as shown in pho to graph of figure 2b, was used as the ini tial prebi - opsy con fig u ra tion of the ro bot through out this study. The tis sue phan tom used was a spec i men of bone less tur key breast ob tained from a butcher. The phan tom was 50 mm long, 40 mm wide and 20 mm thick, sim u lat ing the flat outer hemi sphere of a tar get or gan. The phan tom was se cured to an acous ti cally-ab sorb ing rub ber base in a wa ter bath with mul ti ple pins to pre vent it from mov ing dur ing the bi opsy. Once the ro bot had as sumed its ini tial un folded po si tion dur ing 3D scan ning of the phan - tom, at a user-con trolled trig ger, a com plete 3D im age vol ume of echo data in depth r, az i - muth an gle q and el e va tion an gle f was trans mit ted from the 3D scan ner, with no user se lec tion of im age slices, to a com puter run ning our MATLAB (Mathworks, Natick, MA) im age seg men ta tion al go rithm"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003246_1.4004116-Figure12-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003246_1.4004116-Figure12-1.png",
        "caption": "Fig. 12 Movement sketch of mechanism combined with deformed limacon gear and eccentric shaft",
        "texts": [
            "url=/data/journals/jmdedb/27948/ on 03/26/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use When u1 \u00bc p=2 or u1 \u00bc 3p=2, the maximal pressure angles for the three kinds of noncircular gears with different proportion coefficients are derived and shown in Fig. 11. With the same scale coefficients, the maximal pressure angles of elliptical and limacon gears are equal and smaller than eccentric gear. 3.3 Motion Formula of Serial Mechanism. The oscillation equipment for continuous casting machine is simplified as shown in Fig. 12, which comprised an electromotor, a deformed limacon gear pair, an eccentric shaft, and a linkage mechanism. The length of link O2A fixed on the driven gear represents the eccentricity of the eccentric shaft. Because the length of O2A is much less than the length of BC, it can be considered that point B performs a rectilineal motion along y-axis. The crystallizer to solidify liquid steel is mounted on the oscillation table, which is connected to the four-bar linkage as shown in Fig. 12. Since the parameter ra \u00bc O2A is much smaller than rb \u00bc jABj, the displacement formula of point B can be simplified as follows yB \u00bc ra cos u\u00fe rb (28) The velocity of point B can be derived by differentiating yB as follows B \u00bc x2ra sin u (29) where x2 is the angle velocity of the driven gear. If the input angle velocity x of the driving gear is uniform, Eq. (29) could be represented by B \u00bc xi21ra sin u (30) where i21 is the transmission proportion of the deformed limacon gear pair. The format of nonsinusoidal oscillation for molds in continuous casting machine is the best way at present to improve the surface quality of casting blanks, by which the speed of molds remains in uniform rise and quick drop",
            " The approaches for designing noncircular gears to satisfy the above requirement can be categorized to traditional and optimized design method generally. In traditional methods, the effect of each design parameter must be analyzed first to determine their values for the requirement, but using the optimization method, the proper parameter values can be calculated by computer program one time. So, an optimization method based on the genetic algorithm is employed below. 4.1 Mathematical Model of Serial Mechanism for Optimization. The design variables in Fig. 12 are represented by X \u00bc \u00bdx1; x2; x3; x4 T \u00bc \u00bde; b;m1;x T . Parameter b is the angle of linkage O2A at the beginning position and it yields the equation u \u00bc u2 \u00fe b, where u2 is the polar angle of the driven gear and u is the polar angle of the linkage O2A in the movement process of the mechanism. To obtain uniform velocity in the rising process of molds, the objective function for optimizing the mold velocity is expressed as follows f \u00f0X\u00de \u00bc 1 n Xn i\u00bc1 ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi \u00f0 i 0\u00de2 q (31) where 0 is the expected velocity in the rising region of the mold, i obtained by Eq",
            " Considering the little effect of the length of linkage O2A and AB on the optimization result, they are defined as constant values in the calculation. The constraints include the following: (1) The range of transmission ratio is determined by parameter e \u00bc e=R, which is defined as 0 < x1 < 1. (2) To prevent the undercutting phenomenon, the module m and the minimal curvature radius qmin of deformed limacon gears must satisfy the following equation m2h a\u00f0z0 h a\u00de qmin\u00f0qmin \u00fe mz0\u00desin2a where h a, a, and z0 are the tooth addendum coefficient, profile angle, and number of teeth of a shaper cutter, respectively. (3) In Fig. 12, the linkage O2A can be fixed with any angle at the beginning position, so 0 < x2 < 2p. (4) Based on the equation m1 \u00fe m2 \u00bc 2, the deformed coefficient m1 is defined as 0 < x3 < 2. (5) According to the practical experience, the angle velocity x of the driving gear can be expressed by 0 < x4 < 5. 061004-6 / Vol. 133, JUNE 2011 Transactions of the ASME Downloaded From: http://mechanicaldesign.asmedigitalcollection.asme.org/pdfaccess.ashx?url=/data/journals/jmdedb/27948/ on 03/26/2017 Terms of Use: http://www"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000470_j.ijfatigue.2007.12.003-Figure3-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000470_j.ijfatigue.2007.12.003-Figure3-1.png",
        "caption": "Fig. 3. The configuration of contact fatigue testing machine.",
        "texts": [
            " The wheel specimen has 90 mm in diameter and 15 mm in thickness, and the rail specimen has 110 mm in diameter and 15 mm in thickness. The contact surface thickness of rail specimen was set to 5 mm to maintain a constant contact stress regardless of wear. In order to keep the constant hardness of specimen on contact surface, the heat treatment process was applied. The contact surface was grinded to simulate the surface roughness of railway wheel. The contact fatigue test was carried out on the specially designed fatigue testing machine as shown in Fig. 3. To detect cracks and damage, the vibration signal and cycles were saved continuously by using a computer based acquisition system. The testing machine was stopped in every 50,000 cycles to observe cracks on the surface. One test specimen was forced to revolve with the contact stress ranging from 1000 to 1500 MPa. In order to prevent excessive heating, water was continuously dropped on the contact surface. In the pre-rolling process, 2 105 cycles were applied prior to the main test. After the preliminary process (pre-rolling), the surface was removed by grinding for various depths"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002129_s0263574709005426-Figure4-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002129_s0263574709005426-Figure4-1.png",
        "caption": "Fig. 4. Two-link chain and its associated screw displacements.",
        "texts": [
            " (6)) can be represented by a homogeneous transformation given by p\u03022 = A(\u03b8, t)p\u03021, (7) where A(\u03b8, t) = [ R(\u03b8) d(t) 0 1 ] , (8) and the elements of R(\u03b8) and of d(t), according to Tsai,8 are given by R(\u03b8) =\u23a1 \u23a3 cos \u03b8 + s2 x (1 \u2212 cos \u03b8) sysx (1 \u2212 cos \u03b8) \u2212 sz sin \u03b8 szsx (1 \u2212 cos \u03b8) + sy sin \u03b8 sxsy (1 \u2212 cos \u03b8) + sz sin \u03b8 cos \u03b8 + s2 y (1 \u2212 cos \u03b8) szsy (1 \u2212 cos \u03b8) \u2212 sx sin \u03b8 sxsz(1 \u2212 cos \u03b8) \u2212 sy sin \u03b8 sysz(1 \u2212 cos \u03b8) + sx sin \u03b8 cos \u03b8 + s2 z (1 \u2212 cos \u03b8) \u23a4 \u23a6, and d(t) = ts + [I \u2212 R(\u03b8)]s0. 3.1.2. Successive screw displacements. We now use the homogeneous transformation screw representation to express the composition of two or more screw displacements applied successively to a rigid body.8 Figure 4 shows a rigid body \u03c3 which is guided to a fixed base by a dyad made up of two kinematic pairs, denoted by $1(q1) and $2(q2), respectively. The first kinematic pair connects the first moving link to the base, and the second kinematic pair connects the second link (\u03c3 ) to the first. We call the axis of the first kinematic pair the fixed joint axis and the axis of the second kinematic pair the moving joint axis. As the rigid body is rotated about and/or translated along these two joint axes, the best way to obtain its resultant displacement is to displace the rigid body \u03c3 about/along the fixed axis and, subsequently, displace the body about/along the moving axis"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000867_pesc.2008.4592719-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000867_pesc.2008.4592719-Figure1-1.png",
        "caption": "Fig. 1. Vector diagram of PMSM in stator coordinates",
        "texts": [
            " 2 2 3 *2* M M Me s p LT (16) The foundation formula for implementing the flux estimator (17) may be implemented directly, or approximated by various methods to avoid integrator drift. Stator currents are measured from the motor, which in fact leads to D-Q variables, and voltages vD, vQ are calculated from dc-link voltage since voltage vectors applied are known. D sQsQssQsQ sDsDssDsD tirv tirv 0 0 \u00b7 \u00b7 (17) Initial conditions have to be known, and they can be obtained from the initial position of the rotor with the machine at standstill in Fig. 1 (18). 00 00 0 0 0 00 \u00b7sin \u00b7cos ; rMMQsQ rMMDsD MDs rsrsors (18) Then, stator flux position can be obtained as sD sQ s arctan (19) And finally, torque is estimated from (3). It should be noted that for practical implementation may be more convenient to use current model (5) for flux estimator instead of voltage model (6), if real voltage applied to the motor cannot be measured and is not exactly known. Simulation results for the DTC of PMSM shown in Fig. 4 are presented for various operating conditions"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003348_pime_proc_1970_185_115_02-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003348_pime_proc_1970_185_115_02-Figure1-1.png",
        "caption": "Fig. 1. A typical kinematics of a rolling contact drive (a), and the angular velocity of disc 1 with respect to disc 2 ( b )",
        "texts": [
            " Principal radii of the roller in x and y directions respectively. Principal radii of the toroidal disc track in x and y directions, respectively. Radius of the pitch circle of the track. Radius of curvature of the contact area, 1. ('+'). 2 R, r Y Velocity differences in the contact area in the rolling direction; overall velocity difference of the rolling bodies in the rolling direction. Mean rolling velocity of the contacting bodies. Right-hand Cartesian co-ordinate axes with reference to the centre of the control area, x being the rolling direction (Fig. 1). Coefficient of traction referring to the line contact or strips, the effective coefficient of traction, p = TIP, p = FIN. Angular velocity of the roller and the toroidal disc, respectively. Angular velocity of the spin component normal to the tangent plane of the constant area, i.e. about the z axis. Spin parameter (= w,b/ U). Local slide ratio, 5 = Su/U. KINEMATICS Consider a toroidal disc with a 'grooved' track, rotating about a fixed axis AA, driving a roller follower with an axis pivoted about a point 0 and inclined at an angle with AA [Fig",
            " The experiments with two-disc machines have shown that the coefficient of traction differs greatly from the slide ratio. The author shows how the coefficient of traction at rolling-spinning motion can be calculated from a series of experiments on a two-disc machine, using a summation resulting from the strip approximation. There he used the simplification that the slide ratio is the same along the whole strip as on the centre line of the contact ellipse, where the velocities u1 and have the same direction. Fig. 8 shows the contact area of the drive in Fig. 1. At the centre line there is shown a small skew-angle /3 between the velocities u1 and u2. In my own experiments on the two-disc machine of the Laboratory for Machine Elements at the Munich Technical University, a great influence on the coefficient of traction was found already at skew-angle /3, which is much smaller than one degree. If this sliding motion resulting from the skew-angle /3 were considered in the calculation, then the coefficient of traction would rise more slowly over the slide ratio",
            " In analysis, it is perhaps more profitable to consider it in terms of the parameters used in the modelling of the physical system or the assumptions underlining the analysis which are likely to be affected by the physical scale of the problem. It would appear that on close examination the thermal effect is the likely cause as the contact dimension is increased. Judging from the data given by S . Lingard, the contact load on the second test, at 2000 lbf per roller, is about four times the load on the first test if the same Hertzian pressure is maintained, and there will be a corresponding increase in viscous heating at the contact. Assuming the drive consists of two sets of rollers arranged back-to-back (Fig. 1 l), the rate of heat generating at all the contacts in the second test is in excess of 4 kW. In view of the compactness of the design, high local thermal gradients must exist, and these will be particularly severe on the central disc. Experiences have shown that even with the much simpler configuration of a disc machine, the surface temperature of the discs cannot be effectively controlled by the inlet oil jet temperature when a high ratio of heat is generated. This is due to the presence of a stable, laminar thermal boundary layer (1"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003136_1.3645806-Figure5-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003136_1.3645806-Figure5-1.png",
        "caption": "Fig. 5 Wear t rack for a r a d i a l l y l oaded ba l l bea r i ng toge ther w i t h pred ic ted m ic ros l i p pa t te rn for th is cond i t i on",
        "texts": [
            " 3 and 4 where it is seen tha t the value of y and the locus of c is unaffected although the values of X do increase with increasing load. I t is also seen tha t at light loads the second value of y may not occur within the limits of the contact zone. The effect of increasing R/r is seen in Fig. 3 to result in a positive trend for the values of 7 . T h e effect of decreasing values of a is seen to result in a negative tendency for the values of 7 and for the two values of 7 to approach each other, Fig. 4. Some confirmation of the foregoing arguments is provided by the wear tracks shown in Figs. 5 and 6. Fig. 5 shows the wear tracks obtained for a radially loaded ball bearing and is in agreement with the predictions of a previous analysis [3] as shown in Fig. 5(b). The wear tracks shown in Fig. 6 were obtained from a single ball thrust bearing. The single region of sticking with a slight tendency in the direction of negative y is in agreement with pat terns of the type shown in Fig. 4. Journal of Basic Engineering m a r c h 1 9 6 6 / 215 Downloaded From: http://fluidsengineering.asmedigitalcollection.asme.org/pdfaccess.ashx?url=/data/journals/jfega4/27271/ on 03/03/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use Nm MA SO* tOO*. ZOO*. \u2022 f \u2022 ( [ i w i n r \"yf"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000285_0094-114x(78)90011-3-Figure4-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000285_0094-114x(78)90011-3-Figure4-1.png",
        "caption": "Figure 4. Minimal configuration of manipulator.",
        "texts": [
            " In this way and as the result of performing the algorithm, the position and velocity of the ith member (Aj, qjj, co~, vs) and the vector coefficients (a, ~) in the expressions for accelerations are determined. For the sake of operating store economy, all information on the ith member are put in the place where earlier information on ( i - l)th member were kept. The algorithm results completely from the formulae (2)-(22), and the block-scheme in Fig. 3. Forming of the Algorithm in Detail A mechanical configuration of the kinematic chain with three members was adopted (Fig. 4). The first member of the chain is connected to the absolute system and has 1 degree-of-freedom (O - rotation round the vertical axis of the absolute system). The members are interconnected by joints, forming simple kinematic pairs of the Vth class. Between the first and second member a relative rotation is performed, designated by the angle 6, between the second and third by 0. Accordin~y, the active mechanism has in all 3 degrees-of-freedom. These angles were also adopted to be the generalized coordinates of the system. The geometrical parameters are shown in Fig. 4. The lengths I~ (i = I, 2, 3) are the distances between the center of the ith joint to the center of the gravity of the ith member, and the lengths It (i = I, 2, 3) are the distances from the ith joint center to the i + Ith joint center, i.e. the member lengths. The member masses are m,(i = I, 2, 3). It is supposed that the second and third member are in the form of canes, i.e. homogenous circular cylinders with diameters, nq~v'ble in comparison to their length. Hence these members have two moments of inertia round their center of gravity jm, JN3, Js~, Js,~, for the axes, passing through the center of wavity a"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002945_s10846-011-9649-2-Figure7-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002945_s10846-011-9649-2-Figure7-1.png",
        "caption": "Fig. 7 Output distribution of each state in Markov chain for Rcom = 10 and a N = 125 and b N = 150",
        "texts": [
            " The remaining states that are not explicitly labeled with values ranging from 1% to 8% represent states where the node is moving and has an ideal number of neighbors. When Rcom = 10, 12, and 15, the probabilities of being in the desired state are 32% 22% 16%, respectively. As seen from these results, to reach and stay at the desired state for a mobile node is less probable when communication range increases. It is an expected result since larger communication range means more local neighborhood information and more neighboring nodes that results in a less stable position (aggregated force on a mobile node is not zero). Figure 7 shows the possible outcome percentages of each state in the Markov chain for the experiment comparing varying numbers of nodes with Rcom = 10 and N = 125 and 150, respectively. Referring back to Fig. 6b, Rcom = 10 and N = 100, the probability of being in the ideal state for a mobile node is 32%, the probability of being in the stopped, non-ideal state is only 20% and the sum of all remaining moving states is 48%. When there are a total of 125 mobile nodes in a manet, the probability of being in the ideal state is 27% as seen in Fig. 7a. When N = 150 it can be seen that the probability of being in the ideal state is 18%, the probability of being in the stopped, non-ideal state is now 30% and the sum of all remaining moving states is now 42%. This data reveals that as the network area is overcrowded with mobile agents, increased energy is consumed in the search for optimal spatial orientation. As crowding increases, fewer nodes will be able to find an optimal orientation and more nodes will continue to search for better spatial configuration. Figure 5 demonstrates that these will eventually converge to a stationary distribution, but as seen in Fig. 7, when the number of mobile nodes increases beyond the minimal number necessary to completely cover the network area, more nodes will be in a stopped state with a non-ideal number of neighbors (due to overcrowding) than the stopped state with an optimal number of neighbors and high fitness. Effectiveness in area coverage (Anac) is an important performance metric of our fga approach (Section 3.3). Figure 8 shows the nac for N = 100 and Rcom = 5, 10, 12, and 15. Comparison with the convergence rate of the same experiments discussed in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003348_pime_proc_1970_185_115_02-Figure10-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003348_pime_proc_1970_185_115_02-Figure10-1.png",
        "caption": "Fig. 10. The relative velocities induced by the shaft misalignment on a disc machine",
        "texts": [],
        "surrounding_texts": [
            "D380 DISCUSSION ON S. Y. POON\ncomments, whilst acknowledging the qualitative conclusions of the strip approximation method, are intended to convey some doubts about its validity as an accurate design tool.\nConsider two friction drive point contacts, operating under e.h.1. conditions, with spin, as described in the author's paper. Assume they are elliptical with semi-major axes bl and b, respectively and of the same eccentricity. According to equation (1 l), the relative velocity distributions are given by:\nneglecting c for the purpose of this argument. effective traction coefficients are given by equation (6): Taking the same number of strips in each contact the\nIt can be seen that, for a given Hertzian maximum pressure, provided that U, = U,, G, = Gz, and 0, = SZ,, and that the same disc machine data is used in each case, the strip approximation suggests that the traction coefficient-sliding speed characteristic is identical for each contact. Figs 6 and 7 appear to substantiate this conclusion in that at given P,,,, U, sliding speed, and temperature, the effective coefficient of traction depends only on the spin parameter.\nExperimental evidence obtained in our laboratories, using friction drive components kinematically identical to those illustrated by the author, is not consistent with this interpretation. Typical values are given below :\nRolling speed, 1400 in/s Contact surface temperature at inlet, 80\u00b0C Hertzian maximum pressure, 316 000 lb/in2\nG Q a b Pmax 0.002 0.051 0.0190 0.0397 0.041 0.003 0.049 0.0381 0.0807 0.030\nwhere pmax is the limiting coefficient of traction. The lubricant in the tests was a synthetic fluid developed for use in rolling contact friction drives.\nThe effect of the differences in G and l2 on the relative velocity distribution is insignificant, leading to the conclusion that a scale effect exists which causes a reduction in traction coefficient as the contact size is raised. Further evidence of a scale effect is contained in the contact radius term of Hint's (12) equation for the coefficient of traction in a circular contact at high sliding speed.\nIncidentally, measurements of the traction over a wide range of sliding speeds have given cutves typified by Fig. 9. It shows an almost linear rise of traction up to a maximum of about 5 per cent sliding, compared with the 05-1 per cent mentioned for non-spinning contacts. It also\n007 \\\n0.04\n0.03\n0.01 ' 2 I 0-\n0 0.1 0.2 0.3 0.4 SLIDE/ROLL RATIO 6.0.033 in, u;0.019 in N=3501b,U=200 i n h S2= 0035 ,G=00013 Roller surface temperature = I00\"C\nratio for a spinning joint contact\nshows a steady decay as slide ratio increases beyond 5 per cent. It should be emphasized, however, that this behaviour was recorded with constant roller surface temperature at inlet to the contact, and is probably not representative of conditions in a real transmission if the maximum traction coefficient is exceeded.\nFig. 9. Variation of traction coefficient with slide/roll\nR E F E R E N C E\n(12) PLINT, M. A. 'Traction in elastohydrodynamic contacts', Proc. lnstn rnech. Engrs 1967-68 182 (Pt 11, 300.\nS. Y. Poon Graduate (Author) M. J. French is quite correct that the parameter G related to the geometry of the contact should include a factor\nR R+R, sin 6\nif the second-order term in y is taken into account in the first of equations (2) : the term in question is\nOn the question of the relative effect of temperature and pressure on the traction, the answer is that it not only depends on the type of lubricant, but for the same lubricant, it is affected by the temperature range. For instance, the effect is not nearly as pronounced between 25\" and 50\u00b0C as between 50\" and 75\u00b0C for the mineral oil considered in the paper. The reason is that, at a lower temperature range, the maximum traction coefficients are closer to the upper limit or 'ceiling' of the lubricant (10) for pressure in excess of 176 000 lb/in2. A number of synthetic fluids for rolling contact drive tested seem to indicate that the coefficient of traction is less sensitive to temperature variation than the mineral oils under similar conditions, but at higher operating temperatures or lower contact pressure the behaviour of these fluids exhibits a pattern not unlike the mineral oil. The nature of relative significance of temperature and pressure on the traction has been explored from the viewpoint of physics of fluid by the\nProc lnstn Mech Engrs 1970-71 Vol 185 76/71\nat University of Leeds on June 5, 2016pme.sagepub.comDownloaded from",
            "THE EFFECT OF SPIN ON THE TRACTIVE CAPACITY OF ROLLING CONTACT DRIVES D381\nauthor in his contribution to the discussion of a paper by Plint (12).\nH. Gaggermeier has found that the traction results from a laboratory disc machine are sensitive to even a very small amount of shaft misalignment. Shaft misalignment can be resolved into a spin component and a transverse sliding component (Fig. lo). Even for a small skew-angle, /3, it can be seen that the transverse sliding component, in terms of slide ratio, can be of the same order of magnitude as the forward slide ratio in the region we are concerned with here. One degree of skew, for example, corresponds to a transverse slide ratio of 1 7 . 5 ~ It is, therefore, not surprising that the traction that can be supported in the direction of rolling is greatly reduced, The conclusion, however, in no way invalidates the strip approximation in which the transverse slide ratio is small in comparison with that of the rolling direction,\nS . Lingard has made an interesting observation on the \u2018scale effect\u2019 which causes a reduction in traction coefficient as the contact size is raised. In analysis, it is perhaps more profitable to consider it in terms of the parameters used in the modelling of the physical system or the assumptions underlining the analysis which are likely to be affected by the physical scale of the problem. It would appear that on close examination the thermal effect is the likely cause as the contact dimension is increased. Judging from the data given by S . Lingard, the contact load on the second test, at 2000 lbf per roller, is about four times the load on the first test if the same Hertzian pressure is maintained, and there will be a corresponding increase in viscous heating at the contact.\nAssuming the drive consists of two sets of rollers arranged back-to-back (Fig. 1 l), the rate of heat generating at all the contacts in the second test is in excess of 4 kW. In view of the compactness of the design, high local thermal gradients must exist, and these will be particularly severe on the central disc. Experiences have shown that even with the much simpler configuration of a disc machine, the surface temperature of the discs cannot be effectively controlled by the inlet oil jet temperature when a high ratio of heat is generated. This is due to the presence of a stable, laminar thermal boundary layer (1.1) and (13). It would be too much to expect that uniform temperature can be achieved with the test rig. S. Lingard\u2019s contribution highlights the need for further work on the heat transfer aspect of the rolling contact drive.\nREFERENCE\n(13) O\u2019DONOGHUE, J. P. and CAMERON, A. \u2018Friction and temperature in rolling sliding contacts\u2019, ASLE Trans. 1966 9, 186.\nProc lnstn Mech Engrs 1970-71 Vol185 76/71\nat University of Leeds on June 5, 2016pme.sagepub.comDownloaded from"
        ]
    },
    {
        "image_filename": "designv11_12_0003169_s13198-013-0164-7-Figure8-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003169_s13198-013-0164-7-Figure8-1.png",
        "caption": "Fig. 8 Flux lines distribution: a starting operation of healthy motor; b starting operation of faulty motor with 10 % SE; c steady state operation of healthy motor; d steady state operation of faulty motor with 10 % SE",
        "texts": [
            " Unless of some troubles in previously existent harmonics such as the small increasing in PSH1 amplitude from -34.44 to -34.14 dB, none of SE index signatures are observed, unfortunately the MCSA fails to detect SE fault. It was shown in previously simulation results (Figs. 6, 7) that MCSA failed to detect purely SE in induction motor (Li et al. 2007; Concari et al. 2008; Vitek et al. 2010; Nandi et al. 2011;). In this paragraph the air gap magnetic flux density analysis signature is used to detect SE fault. Figure 8a\u2013c presents the magnetic flux distribution, the flux line maps are taken from healthy and faulty saturated induction motor under 10 % of SE at starting operation and steady state operation it can be observed that when SE occurs, the magnetic flux distribution is deformed. 4.1 Spectrum analysis of air gap magnetic flux density of healthy and faulty motor Figure 9a, b presents the air gap flux density spectrums for full-load healthy induction motor under linear and nonlinear permeability. Comparison of Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002984_00368791311303474-Figure1-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002984_00368791311303474-Figure1-1.png",
        "caption": "Figure 1 Non-recessed symmetric hole-entry journal bearing system",
        "texts": [
            " They found that for a given eccentricity, adjustment configuration and degree of misalignment with unidirectional load increases the friction variable and shaft attitude angle. To the best knowledge of authors no investigation is yet available in literature that address to the performance of a compensated hole-entry hybrid misaligned worn journal bearing operating in turbulent regime. The present study is therefore aimed to study analytically the performance of symmetric hole-entry hybrid journal bearing system as shown in Figure 1, operating in turbulent regime by considering the combined influence of wear and misalignment in the analysis. The Reynolds equation based on Constantinescu lubrication theory has been solved by using FEM together with orifice and capillary restrictors flow equation as a constraint. The variation in the bearing static and dynamic characteristics parameters have been presented for the various values of bearing operating and geometric parameters as a function of wear depth parameter and misaligned parameters for the various values of Reynolds numbers"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002241_978-1-4471-4141-9_70-Figure70.5-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002241_978-1-4471-4141-9_70-Figure70.5-1.png",
        "caption": "Fig. 70.5 The CAD model of the proposed triangular prism mechanism module. a Theoretic model, b deployed configuration, c folded configuration",
        "texts": [
            " In order to build deployable mechanisms that can be deployed onto a surface structures, such as spherical surface or parabolic surface, the triangular prism mechanical modules with non-equilateral triangular base profiles are required, because any surface can be approximated by a set of triangular profiles [1]. Fig. 70.4 shows a conceptual model of a trussed surface structure that is consisted of triangular prism, ABC and A0B0C0 are the two bases for one module. In the complicated surface, however, the parameters for the modules in the assembled structure are different from each other, this fact makes the mobile assembly of large deployable mechanism very difficult. Using GLSBL as the bases of the prism, one can easily construct a tripod mechanism as shown in Fig. 70.5. GLSBL and three long links forms the bases of the tripod and joints ki\u00f0i \u00bc 1; 2; 3\u00de connect GLSBL and the long links. For simplicity, the direction of joint ki\u00f0i \u00bc 1; 2; 3\u00de is along zi\u00f0i \u00bc 1; 3; 5\u00de, respectively, as shown in Fig. 70.5. It can be deployed onto a triangular prism deployed profile from a bundle compact form with all the links being parallel and contact to each other. As we can also use the trihedral Bricard linkage to design a non-equilateral triangular base profile with different physical link lengths, but in the tripod mechanism, it is required that the lengths of all physical links being identical, therefore, only the GLSBL case can be applied to this deployable tripod mechanism. As shown in Fig. 70.6, DG; GF; FI; IE; EH; HD are the DH length of links, Q is the concurrent point of the axes z2; z4; z6; P is another concurrent point of axes z1; z3; z5"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003973_s12206-012-0811-y-Figure6-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003973_s12206-012-0811-y-Figure6-1.png",
        "caption": "Fig. 6. The finite element model.",
        "texts": [
            " CONTA173 is normally used to represent contact and sliding between 3-D \"target\" surfaces defined by TARGE170 and a deformable surface, defined by this element. The element is applicable to 3-D structural and coupled field contact analyses. This element is located on the surfaces of 3-D solid or shell elements without mid side nodes. It has the same geometric characteristics as the solid or shell element face with which it is connected. Contact occurs when the element surface penetrates one of the target segment elements on a specified target surface. The finite element model generated (Fig. 6) has 116224 elements. Out of this 115058 are SOLID 92 element, 693 are CONTA173 and 472 TARGE170. The three contact pairs are Piston \u2013 Piston pin, Piston pin \u2013 connecting rod and connecting rod \u2013 crank shaft. Material properties required for this analysis are modulus of elasticity, Poisson\u2019s ratio and density. The material properties of various elements of the analysis are listed in Table 1. The next step is to define loads and constraints required for the analysis. To account for an inertia effect like gravity, appropriate values for g (9"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002167_cca.2009.5280933-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002167_cca.2009.5280933-Figure2-1.png",
        "caption": "Fig. 2. Spindle weight balancing scheme: 1 \u2013 head of spindle; 2 \u2013 journal bearing; 3 \u2013 two balancing springs.",
        "texts": [
            " A universal joint kinematics is widely represented in the literature. But the rolling mills spindles weight balancing units have not been investigated enough in relation to transient torsional vibration control. The main problem is that spindles are equipped with the spring type weigh balancing units where a compensation force depends on work rolls vertical position which is variable with the rolls diameters and strip reduction. The two springs are fixed just by the screws at the bottom of supporting journal bearing (see Fig. 2). The controlled hydraulic type units are also used but in much less number of stands of the rolling mills. Stiffness of the weight balancing springs varies in the range of 0.5-3.0 KN/mm (for both springs) and spindle weight is about 20-200 KN for the different stands. So to balance spindle by 5% of its weight the spring position has to be regulated by the screws with accuracy of 1-2 mm. It is impossible during strip rolling and in maintenance practice and it leads to unavoidable backlashes gap opening twice per spindle rotation before transient"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000074_j.gaitpost.2006.09.079-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000074_j.gaitpost.2006.09.079-Figure2-1.png",
        "caption": "Fig. 2. Model II: A modified compass gait model is applied at the time instant of maximum CoM height and includes single support Knee flexion, leg inclination, pelvic obliquity, and single support heel rise.",
        "texts": [
            " For each trial the model geometry was defined using the 3-D position of each joint center determined from a subject\u2019s kinematics at the instant of minimum CoM. The isolated contributions of an individual determinant were computed as the difference between the CoM minimum height and the corresponding minimum calculated using the model with the individual determinant set to zero. These data were normalized by the clinically measured excursion of each subject to account for leg and step length differences. The second model (Model 2, Fig. 2), composed of three segments representing thigh, shank, and pelvis, was used to evaluate of the effects of determinants at the maximum CoM height (during single support). The effects of single support knee flexion, leg inclination (i.e. the antero-posterior distance between hip and ankle joint centers of the supporting limb), pelvic obliquity, and single support heel rise on the maximum CoM height were computed in the same manner as used for Model 1. For the purpose of normalization, a third simple compass gait model, was used to compute the excursion a subject would have for his or her average step and leg length if they had neither ankle or knee joints"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000935_156855208x336684-Figure4-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000935_156855208x336684-Figure4-1.png",
        "caption": "Figure 4. Proposed scheme for the formation of OPH\u2013QD bioconjugates. (Adapted from ref. [107].)",
        "texts": [
            " Therefore, the emergence of QDs in the toolbox of fluorescence experiments [103, 104] has opened ever-increasing perspectives in analytical detection [105, 106]. Leblanc\u2019s group reported a novel biosensor for the detection of paraoxon based on (CdSe)ZnS core\u2013shell QDs and an OPH bioconjugate [107]. The OPH was coupled to (CdSe)ZnS core\u2013shell QDs through electrostatic interaction between negatively charged QDs surfaces and the positively charged side-chain and end groups (NH2) of protein (as schemed in Fig. 4). The FL intensity of the OPH\u2013QD bioconjugates was quenched in the presence of paraoxon. The quenching mechanism was that the conformational change in the enzyme influences the degree of surface passivation of the QD. The detection limit of paraoxon concentration using OPH\u2013QD bioconjugates was about 10\u22128 M. The sensitivity of the biosensor could be increased through increasing the OPH molar ratio in the bioconjugates. Moreover, they also fabricated a polyelectrolyte architecture composed of chitosan and OPH polycations along with thioglycolic acid-capped CdSe QDs [108, 109]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000154_iros.2007.4399278-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000154_iros.2007.4399278-Figure2-1.png",
        "caption": "Fig. 2. Approach of PVTOL to control the roll position",
        "texts": [
            " The aerodynamics and thrust moments can be denoted by M b A,T = [ \u2113 m n ]T , they are shown in the Figure 1, then using the matrix of inertia and the moment vector, the equation (5) yields: P\u0307 = (Jy \u2212 Jz)QR Jx + \u2113 Jx (10) Q\u0307 = (Jz \u2212 Jx)RP Jy + m Jy (11) R\u0307 = (Jx \u2212 Jy)PQ Jz + n Jz (12) This section presents three decoupled stability augmentation control systems for the roll, the pitch, and the yaw positions of the vehicle in hover flight. These subsystems will be obtained using only the kinematics and moment equations from the general model. Several aerodynamic factors will be taken into account to obtain the transfer function that represents the dynamic of each system. To obtain the roll control system, it is assumed that the pitch and yaw rates are zero. Then, the vehicle can be analyzed in a similar manner to a PVTOL flight platform, as in [1]. This configuration is shown in Figure 2. Therefore, using the equations (7) and (10), the rotational dynamics for the roll angle can be represented by: \u03c6\u0308 = \u2113/Jx (13) where, the sum of moments \u2113 can be calculated as follows: \u2113 = F \u00b7 d\u2212 C\u2113\u03c6\u0307 \u03c6\u0307 (14) and F = f1 \u2212 f2 is the force difference between the right and left rotor and d is the distance from the center of mass to each rotor. The second term in the right side of equation (14) represents an aerodynamic moment produced by the change of the roll rate, normally opposing to the roll moment, that is why, the derivative, C\u2113\u03c6\u0307 = 0"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000545_pccon.2007.372977-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000545_pccon.2007.372977-Figure2-1.png",
        "caption": "Fig. 2. Inductance profiles of 12/8 SRM.",
        "texts": [
            " VOLTAGE EQUATION OF SRM where t, v, R , i, and SJ show the time, the applied voltage, the phase resistance, the phase current, and the flux linkage respectively. The relation between the phase inductance L and the flux linkage is given by U = L(i,O)i (2) where L and 0 stand for the phase inductance and the rotor position. From (1) and (2), the voltage equation is given by the following equation. dL dL)div-Ri+icd +(L+I ) dt dt dt (3) where co stands for the rotor angular velocity. The second term on the right hand side is the back-e.m.f. E, and can be neglected at standstill. 1-4244-0844-X/07/$20.00 \u00a92007 IEEE. 259 Fig. 2 shows the phase inductance profiles of tested 12/8 SRM for an electrical cycle. The phase inductance varies regularly with the rotor position as shown in Fig. 2. Each phase inductance has same waveform and different phase. The phase inductance increases as the rotor pole approaches the stator pole. When the stator and rotor poles are aligned, the phase inductance shows the maximum value, and thereafter the phase inductance decreases. Therefore, the rotor position is obtained by observing the value of the phase inductances of all phases. At standstill, the initial rotor position is defined as the angle between the stator and rotor poles their distance is the shortest of all pairs",
            " Therefore, assuming that the back-e.m.f. and the voltage drop of the phase resistance are ignored, the voltage equation is given by the following equation from (3). Each phase inductance can be considered as a vector. Fig. 3 shows an example of the composition of the inductance vectors when the initial rotor position is 160 electrical degrees. In this figure, the magnitude of the inductance vector L Z240' shows the value of the U- phase inductance LU when the initial rotor position is 160 degrees in Fig. 2. The direction of the inductance vector LU 240' shows the position where the stator pole of Uphase aligns with the rotor pole. The inductance vectors LVZO' and LWZ1200 show the same meaning. The composed vector L/OSe tn at the initial rotor position can be obtained using each phase inductance vector as follows: LZO ini = L /2400+L /00+L /1200 (7) where 0ein stands for the initial rotor position expressed by electrical angle. In this case, 0ini is 160 degrees. Fig. 4 shows the decomposition of the composed vector LZOetnt, In this figure, La and Lb stand for the horizontal component and the vertical component of the composed vector LI Oe ini and they are calculated by the following equations"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000549_tmag.2008.2001316-Figure8-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000549_tmag.2008.2001316-Figure8-1.png",
        "caption": "Fig. 8. Eddy current distribution of permanent magnet synchronous motor.",
        "texts": [
            " From the result of Fig. 5, the insulated layer is set as 1 mm and the resistivity at 5 m is varied from 10 to 1 m. The magnets in the motor are affected by various kinds of harmonic magnetic fields. In this case, 6th phase band harmonics and 12 th slot harmonics are considerable. As the frequency of the power supply is 60 Hz, the frequencies of these harmonics are 360 Hz and 720 Hz, respectively. Among them, the 12th harmonic field is largest because the number of the stator slots per pole is 12. Fig. 8 shows the calculated eddy current distribution in the magnets when the resistivity is 1 m. The analyzed region is reduced as half core length of one pole-pair due to the symmetry. The number of the finite elements is 2 177 760. The time domain formulation is applied and the number of the time steps per one time period is set as 256 to estimate the harmonics correctly. The total calculation time is about 700 h by Pentium 4\u20133.6 GHz PC. Fig. 9 shows the decomposed 6th, 12th, and 24th harmonic eddy current distributions in one magnet"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002043_pime_proc_1967_182_025_02-FigureI-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002043_pime_proc_1967_182_025_02-FigureI-1.png",
        "caption": "Fig. I4",
        "texts": [
            " In the tests described the ratio of slope amplitude to translation amplitude of the journal at a rotor resonance is exceptionally high, due to the very flexible rotor. It is not surprising therefore that the damping of the oil film is higher than that predicted by assuming pure translation of the journal. The rotor is certainly unrepresentative of, for instance, a large turbogenerator rotor in this respect. In the latter case, at its first modal shape supported upon typical bearings the ratio of slope amplitude to translation amplitude is of the order of 0.0408 rad/in, whereas in the present case this figure is approximately 19.3 rad/in. Vol182 Pt I No 13 at NANYANG TECH UNIV LIBRARY on June 9, 2016pme.sagepub.comDownloaded from INFLUENCE OF COUPLED ASYMMETRIC BEARINGS ON THE MOTION OF A MASSIVE FLEXIBLE ROTOR 263 This large slope to translation ratio at the journal may well result in another discrepancy between theory and experiment. The experimental critical speeds occurred at 7-75 Hi! within 0.01 Hz, as against a predicted value of 7.45 Hz. Now these critical speeds, for reasons explained above, are very insensitive to oil film stiffness, indeed the characteristic frequency for the pinned-pinned rotor supported at its bearing centres should in theory differ by much less than 0-01 Hz from the frequency for the rotor on supports of the order of flexibility of the bearings used in the experiments"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003713_978-1-4614-2419-2-Figure54.4-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003713_978-1-4614-2419-2-Figure54.4-1.png",
        "caption": "Fig. 54.4 Behavior of concrete under uniaxial loading in (a) tension & (b) compression",
        "texts": [
            " The behavior of concrete in the numerical simulations was incorporated using damage plasticity model available in the code. The model is based upon the concept of isotropic damaged elasticity in conjunction with isotropic tensile and compressive plasticity to represent the inelastic behaviour of concrete. The model can be used for the analysis of reinforced concrete structures subjected to monotonic, cyclic and dynamic loading. Under uniaxial tension, the stress\u2013strain response follows a linear elastic relationship until the value of failure stress (st0) is reached, Fig. 54.4a. The failure stress corresponds to the 526 M.A. Iqbal et al. 54 Response of Nuclear Containment Structure to Aircraft Crash 527 onset of micro-cracking in the concrete. Beyond the failure stress the formation of micro-cracks is represented macroscopically with a softening stress\u2013strain response which induces strain localization in the concrete. Under uniaxial compression, the response is linear until the value of initial compressive strength sc0, Fig. 54.4b. In the plastic region the response is typically characterized by strain hardening followed by strain softening beyond the ultimate stress, scu. When the concrete specimen is unloaded from any point on the strain softening branch of the stress\u2013strain curve, the unloading response is weakened. The elastic stiffness of the material appears to be degraded. The degradation of the elastic stiffness is characterized by the tension and compression damage variables, dt and dc respectively. The stress\u2013strain relations under uniaxial tension and compression loading are governed by the following expressions respectively; st \u00bc 1 dt\u00f0 \u00deE0\u00f0 et ~e plt \u00de (54"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000569_tie.2008.2009991-Figure2-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000569_tie.2008.2009991-Figure2-1.png",
        "caption": "Fig. 2. Inverted-pendulum model.",
        "texts": [
            " (2) The relations between workspace and joint coordinate space in velocity and acceleration are derived as follows by differentiating the earlier equations with respect to time x\u0307 =Jaco\u03b8\u0307 (3) x\u0308 =Jaco\u03b8\u0308 + J\u0307aco\u03b8\u0307 (4) where Jaco is the Jacobian matrix. The dynamics of the two-link manipulator is shown in the following equation, where T = (T1, T2)T is the torque of the motors, J(\u03b8) is the inertia matrix, and H(\u03b8, \u03b8\u0307) is a member of centrifugal and Coriolis forces, respectively, T = J(\u03b8)\u03b8\u0308 + H(\u03b8, \u03b8\u0307). (5) B. Inverted Pendulum The inverted pendulum in this paper has 2-DOF which means that it falls down in every direction. The inverted pendulum is modeled as shown in Fig. 2. The root of the pendulum is the origin of the pendulum coordinate. The pendulum has the mass mp and the length lp. \u03c6 is the angle from the z\u2032-axis. lpx and lpy are the lengths of the virtual pendulum projected on the x\u2032\u2013z\u2032 and y\u2032\u2013z\u2032 planes, respectively. \u03b1 and \u03b2 are the angles of the virtual pendulum from the z\u2032-axis. Furthermore, the directions of the x\u2032-, y\u2032-, and z\u2032-axes are identical with that of the x-, y-, and z-axes, respectively, and x\u2032\u2013y\u2032 plane corresponds to x\u2013y plane in the world coordinate"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0001546_j.jmatprotec.2010.12.002-Figure5-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0001546_j.jmatprotec.2010.12.002-Figure5-1.png",
        "caption": "Fig. 5. Finite elem",
        "texts": [
            " A sequence of trials has been conducted to select uitable number of elements particularly at the area closer to the , (b) tube ring of 30 mm length, and (c) weld geometry in the ring. weld line and in the thickness direction. More specifically geometries with 4, 6 and 8 elements through the thickness were tested for tube with 1.6 mm thickness. It was decided to use 8 elements for the 1.6 mm thick tube since this mesh required reasonable convergence and solution time with no significant loss of accuracy. A dense mesh was used in the area along the weld line, as shown in Fig. 5, and a coarser mesh for the rest of the structure. The final mesh was the result of compromise between computing time and accuracy. The finite element mesh was refined till mesh convergence was achieved, i.e. the variation in predicted temperature and residual stress values at some selected points became negligible. The mesh has 132,713 eight-node elements and 16,108 nodes. An appropriate time-stepping scheme was used for each analysis to achieve fast convergence of the solution and reasonable accuracy"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000530_t-pas.1976.32180-Figure5-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000530_t-pas.1976.32180-Figure5-1.png",
        "caption": "Fig. 5. Waveform of Induced and Terminal Voltages-Current Fed Induction Motor s ~ / U ~ ~ 5 . 7 , s=O.lO.",
        "texts": [
            ";= (33) Multiplying by J2/3 and taking the real part yields the \"exact\" form of the induced phase to neutral voltage For normal operation the real exponential is nearly unity over mode one and, since it is reset to unity at the end of each mode, can be neglected. The waveform then consists of segments of a sinusoid of rotor frequency. At each switchingtransition an additionalphase angle of ~ s / 3 is introduced. These discrete additional phase shifts cause a discontinuity at eachmodeboundary and also provide precisely the additional phase angle over one cycle to convert the rotorfrequency+ to stator frequency w. Figure 5 illustrates this waveform for s-0.10; the slip is chosenlarge to accentuate the discontinuity at each switching point for purposes of illustration. At more normal slip the discontinuity is quite small. Figure 5 also illustrates the voltage drop and the switching impulse voltages which m s t be added to the internal voltage to obtain the terminal voltage. For small slip, the magnitude of e can be expressed as a ea = ~ 2 / 3 (1- a) R wL I r s s (35) which is approximately (1-0) times the voltage calculated by using the fundamentalcurrent in the standard equivalent circuit (neglecting the stator resistance). Clearly a fundamental component approximation is adequate for CSI operation at small slip and frequencies satisfying w>>ao"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0003603_0954406212466479-Figure3-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0003603_0954406212466479-Figure3-1.png",
        "caption": "Figure 3. Generating processes of the conjugated straight-line internal gear pair: (a) pinion and (b) internal gear.",
        "texts": [
            " \u00bc s r1 \u00f014\u00de Therefore, the mathematical model of the pinion profile can be illustrated by R p 1\u00f0x1, s, , z1\u00de \u00bc Roa 1 Rab 1 Rbc 1 0 \u00f015\u00de Mathematical model of the rack-cutter for the pinion with straight-line profile This article assumes rack-cutter as the cutting tool for generating the pinion with straight-line profile. According to the generating principle, the pitch line of the rack-cutter is always tangent to the pitch circle of the pinion in the cutting process,1 which means that the profile of the rack-cutter is conjugated to that of the pinion. Therefore, the profile of the rack-cutter can be regarded as an envelope to the family of the pinion surfaces. In what follows, the profile of the pinion is assumed as the generating surface while the profile of the rack-cutter as the generated surface. Figure 3(a) shows the generating process of the pinion by a rack-cutter. Coordinate systems Sc(Oc, xc, yc) and Sg(Og, xg, yg) are built, which are rigidly connected to the rack-cutter and the ground, respectively. Here, the xc axis of Sc coincides with the pitch line of the rack-cutter and the origin of Sg with the pinion centre. The pitch line of the rack-cutter is tangent to the pitch circle of the pinion at point Pc.Figure 2. Tooth shape of the pinion with straight-line profile. at University of Bristol Library on January 6, 2015pic",
            " Mathematical model of the internal gear with conjugated profile As mentioned previously, the profile of the internal gear is conjugated to that of the pinion. Therefore, the profile of the internal gear can be regarded as an envelope to the family of the pinion surfaces. Then, the pinion is assumed as a shaper-cutter to generate the internal gear and the mathematical model of the pinion is regarded as the generating surface. Coordinate system S2(O2, x2, y2) is built to demonstrate the relationship between the pinion and the internal gear during the generation (Figure 3(b)). Here, S2 is rigidly connected to the internal gear and the origin of S2 coincides with the gear centre. The pitch circle of the internal gear, which has a radius r2, is tangent to that of the pinion at point P. The rotation angles of S1 and S2, \u20191 and \u20192, are related by equation (32). r1\u20191 \u00bc r2\u20192 \u00f032\u00de Here r2 \u00bc 1 2 mlz2 \u00f033\u00de where z2 is the tooth number of the internal gear. Then, according to the theory of gearing, the profile of internal gear can be obtained by equations (34) and (35). Ri 2\u00f0\u20191, \u20192,x1, s, , z1\u00de \u00bcM21\u00f0\u20191, \u20192\u00deR p 1\u00f0x1, s, , z1\u00de \u00f034\u00de N p 1 v \u00f0 pi\u00de 1 \u00bc 0 \u00f035\u00de where Ri 2\u00f0\u20191,\u20192, x1, s, , z1\u00de is the envelope to the family of the pinion surfaces, M21\u00f0\u20191, \u20192\u00de the matrix for coordinate transformation from S1 to S2, as represented by equation (36), and v \u00f0 pi\u00de 1 the relative velocity between the pinion and the internal gear, as represented in S1 by equation (37)",
            " Parameter Value z1 28 z2 38 s 5 mm 25 h a 1 h d 1 at University of Bristol Library on January 6, 2015pic.sagepub.comDownloaded from profilemeans the modified profile in the following part of this article. As mentioned above, shapes of the tooth spaces of the rack-cutter and the internal gear can be obtained, as shown in Figure 6(a) and (b), respectively. With Figures 2 and 6, the basic geometric dimensions of the rack-cutter, the pinion, and the internal gear can be calculated by the parameters of z1, z2, s, , h a, and h d, as listed in Table 2. Based on the kinematic relationships shown in Figure 3, Figure 7 demonstrates the generating processes of the pinion and the internal gear by their at University of Bristol Library on January 6, 2015pic.sagepub.comDownloaded from corresponding cutting tool. Here, the rack-cutter is located outside the pinion, while the shape-cutter is located inside the internal gear. Obviously, the generated profile is the envelope of the families of the cutter surfaces. Moreover, the geometric models can be built with the help of the computer-aided software SolidWorks, as shown in Figure 8",
            " According to some literatures,7,22 overcutting can be determined by the appearance of singular points on the generated surfaces. In this approach, the generated gear surfaces are usually derived by the given surfaces of the cutting tool. However, our study presents a contrary case, in which the surfaces of the rack-cutter are obtained by the given surfaces of the pinion, as mentioned in the \u2018Mathematical model of the rack-cutter for the pinion with straight-line profile\u2019 section. Considering the kinematic relationship shown in Figure 3(a), the pinion can be assumed as a cutting tool for the rack-cutter. Meanwhile, the pinion is regarded as the shaper-cutter for the internal gear, as shown in Figure 3(b). Then, based on these assumptions, the profiles of the rack-cutter and the internal gear can be obtained, as demonstrated in Figure 9. Here, the pinion has a tooth number of 17 and other parameters are listed in Table 1. Obviously, singular points have appeared in the profiles of the rack-cutter and the internal gear. In at University of Bristol Library on January 6, 2015pic.sagepub.comDownloaded from addition, with envelopes shown in Figure 10, it is clear that overcutting has occurred to the profiles in the positions inside the circles"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0002733_0954410012464002-Figure4-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0002733_0954410012464002-Figure4-1.png",
        "caption": "Figure 4. Flight path angle.",
        "texts": [
            " A picture describing these angles for a particular case ( A \u00bc 0) is given in Figure 2, where Fresultant is the sum of the gravitational force (W) and the centrifugal force (Fcf). Helical turn is a maneuver in which the helicopter moves along a helix with constant speed (Figure 3). In a helical turn, the flight path angle is different than zero being given by sin\u00f0 FP\u00de \u00bc sin\u00f0 A\u00de cos\u00f0 F\u00de cos\u00f0 F\u00de sin\u00f0 A\u00de cos\u00f0 A\u00de sin\u00f0 F\u00de cos\u00f0 A\u00de cos\u00f0 A\u00de sin\u00f0 F\u00de cos\u00f0 F\u00de \u00f04\u00de A picture describing the flight path angle for a particular case, A \u00bc 0, F \u00bc 0, is given in Figure 4. Note that _ A 4 0 is a clockwise turn and _ A 5 0 is a counterclockwise turn (viewed from the top) while FP 4 0 is referring to the ascending flight and FP 5 0 is referring to the descending flight. Note: the numerical results reported from here on (i.e. for trim, linearization, and control design) were obtained using Puma SA 330 data.17,40 In this article, trim is defined as the condition for which level banked or helical turn with constant VA and zero sideslip is achieved. For our helicopter model, 29 trim equations and 29 unknowns were at University of Bath - The Library on June 11, 2015pig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_12_0000155_pime_proc_1976_190_050_02-Figure3-1.png",
        "original_path": "designv11-12/openalex_figure/designv11_12_0000155_pime_proc_1976_190_050_02-Figure3-1.png",
        "caption": "Fig. 3: The contact of spherical asperities during sliding",
        "texts": [],
        "surrounding_texts": [
            "A CONTRIBUTION TO THE THEORY OF FRICTION AND WEAR AND THE RELATIONSHIP BETWEEN THEM 485\nF i n a l l y i t mus t b e r e c o g n i s e d t h a t any p l a s t i c d e f o r m a t i o n w i l l mod i fy t h e g e o m e t r i c a l n a t u r e o f t h e s u r f a c e and t h e r e s u l t a n t work h a r d e n i n g w i l l a l s o m o d i f y t h e form o f t h e s t ress s t r a i n c u r v e . T h i s wear law i s t h e r e f o r e o n l y s t r i c t l y t r u e f o r a s i n g l e p a s s o f o n e a s p e r i t y of t h e u p p e r s u r f a c e . I n p r a c t i c e t h e n a t u r e o f t h e c o n t a c t s w i l l b e a mixed e l a s t i c and p l a s t i c p o p u l a t i o n which d u r i n g \u201c r u n n i n g i n \u201d w i l l move t o w a r d s a b a s i c a l l y e l a s t i c p o p u l a t i o n . Thus t h e two e x t r e m e f o r m u l a t i o n s o f t h e wear l a w i n d i c a t e t h e r a n g e o f p r a c t i c a l b e h a v i o u r .\nC O N C L U S I O N S\nThe f o r e g o i n g p r e s e n t s a m a t h e m a t i c a l model o f t h e f r i c t i o n and wear r e s u l t i n g f rom t h e i n t e r a c t i o n o f p o p u l a t i o n s o f s p h e r i c a l a s p e r i t i e s . The e n s u i n g e q u a t i o n s a re a p p l i c a b l e t o a wide r a n g e o f e n g i n e e r i n g s i t u a t i o n s a l t h o u g h t h e wear r e s u l t s mus t b e r e s t r i c t e d t o cases where f < 0 . 5 . The r e s u l t i n g e q u a t i o n s a re i n e v i t a b l y e x p r e s s e d i n f i n i t e mathem a t i c a l fo rm b u t t h e i r r e a l v a l u e i s i n i d e n t i f y i n g t h e i m p o r t a n t p h y s i c a l p a r a m e t e r s and p r o v i d i n g a f e e l f o r t h e i r r e l a t i v e s i g n i f i c a n c e . Thus t h e c o e f f i c i e n t o f f r i c t i o n i s s e e n t o b e d e p e n d e n t on terms d e f i n i n g d e f o r m a t i o n and s h e a r s t ress . I n t h e wear e q u a t i o n s t h e i m p o r t a n c e o f s u c h f a c t o r s a s t h e d u c t i l i t y and t h e mean s l o p e o f t h e a s p e r i t i e s a r e d e m o n s t r a t e d a l t h o u g h t h e b a s i c wear law h a s t h e g e n e r a l l y a c c e p t e d form.\nA P P E N D 1 X\nThe f o l l o w i n g a r g u m e n t s a r e r e s t r i c t e d t o l i n e t r a c i n g s of a s u r f a c e wh ich is b a s i c a l l y t h e model u s e d i n t h e p r e c e e d i n g a n a l y s i s . Fo r s u c h a l i n e t r a c i n g p h y s i c a l i n t u i t i o n s u g g e s t s t h a t t h e p a r a m e t e r s 0 , r) , B are n o t i n d e p e n d e n t s i n c e c l e a r l y any change i n rl m u s t a f f e c t t h e o t h e r p a r a m e t e r s . T h e i r r e l a t i o n s h i p may b e most e a s i l y d e m o n s t r a t e d by e x p e r i m e n t s on a b r a d e d s u r f a c e s which h a v e d i s t r i b u t - i o n s s i m i l a r t o t h o s e assumed f o r t h e d e r i v a t i o n o f t h e w e a r law. Of t h e t h r e e p a r a m e t e r s , 0 a n d r) a r e m o s t e a s i l y measu red a c c u r a t e l y a n d show t h e r e l a t i o n - s h i p , f i g u r e A l .\nC J n = C o n s t a n t A 1\nFrom t h e s e r e s u l t s t h e c o n s t a n t a p p e a r s t o be t h e mean s l o p e o f t h e a s p e r i t i e s , b e i n g s m a l l e r f o r t h e s m o o t h e r s u r f a c e s , and h a v i n g v a l u e s wh ich a r e c o n s i s t e n t w i t h expeir ience.\nIn an earlier paper lo it was shown that t h e mean s l o p e f o r s u r f a c e s , w , i s g i v e n by I\nA2\nSuch a r e s u l t i s c o n s i s t e n t w i t h s e v e r a l o b s e r v a t i o n s i n t h e l i t e r a t u r e on wha t i s now c a l l e d t h e P l a s t i c i t y I n d e x .\nGreenwood a n d Williamson\u2019\u2019 u s i n g a s p h e r i c a l a s p e r i t y mode l , -\n- P l a s t i c i t y I n d e x = gl/i H\nEl - ( 1 . 7 -+ 2 . 4 ) s i n w H\nK r a g e l s k i i \u2019 l u s i n g a c o n i c a l a s p e r i t y mode 1 , P l a s t i c i t y I n d e x = E\u20191.8 t a n w\nH a l l i d a y \u201d u s i n g e x p e r i m e n t a l v a l u e s ,\nP l a s t i c i t y I n d e x = ; (1.2+2.5) t a n w\nH\nE l\nCombining A 1 and A 2 s u g g e s t s an e q u a t i o n\nf r) 0 = ( 0 . 4 + 0.6)\nA 3\nI n t h e m e a s u r i n g o f s u r f a c e p r o f i l e s t h e v a l u e of B i s most s u s p e c t due t o t h e e f f e c t o f s a m p l i n g l e n g t h b u t F i g . A 2 shows some s u c h measu remen t s wh ich s u g g e s t t h a t ,\n0 = 0 .4 (P}\u2019 = mean s l o p e of B a s p e r i t i e s . A4\nT h i s r e l a t i o n s h i p h a s b e e n u s e d i n t h e f o r m u l a t i o n of t h e wea r l a w s i n t h e p r e c e e d i n g a n a l y s i s .\nREFERENCES\n1.\n2.\nBOWDEN,F.P. & T A B O R , D.\nBOWDEN,F.P. & T A B O R , D\n3 . MOORE, D.\n4 .\n5 .\nH A L L I N G , J & N U R I , K.A.\nHALLING, J.\nThe F r i c t i o n and L u b r i c a t i o n of S o l i d s . V O l . I C l a r e n d o n P r e s s , O x f o r d 1950.\nThe F r i c t i o n and L u b r i c a t i o n o f S o l i d s V O l . I1\nC l a r e n d o n P r e s s , Oxfo rd 196 4.\nThe F r i c t i o n and L u b r i c a t i o n o f E l a s tome rs . Pergamon, O x f o r d , 1972\nC o n t a c t o f Rough S u r f a c e s o f WorkH a r d e n i n g M a t e r i a l s\nI . U . T . A . M . Symposium, u n i v e r s i t y o f Twen te , H o l l a n d , 1974.\nA C o n t r i b u t i o n t o t h e T h e o r y o f M e c h a n i c a l Wear.\nWear 34, 239, 1975.\n@ IMwW 1976 Roc lnstn Mech Engrs Vol190 43/76\nat WEST VIRGINA UNIV on June 5, 2016pme.sagepub.comDownloaded from",
            "J . HALLING 486\n6 .\n7.\n8.\n9 .\nHALLING, J. A C o n t r i b u t i o n t o t h e 1 2 . TAVERNELLI I \" A C o m p i l a t i o n a n d T h e o r y o f F r i c t i o n . J .F . & COFFIN, I n t e r p r e t a t i o n o f\nU.S.M.E. /T/40/75. T e s t s \" L.F. C y c l i c / S t r a i n F a t i q u e\nTABOR, D . T h e H a r d n e s s o f M e t a l s O x f o r d U n i v e r s i t y P r e s s ,\nT r a n s . A m SOC. Me ta l s , 5 1 , 4 3 8 , 1 9 5 9 .\nR e l a t i o n s h i p b e t w e e n 1 9 5 1 . 1 3 . SUH,N.P. &\nSRIDHARAN,P. t h e c o e f f i c i e n t o f JEFFERIS,M.A. & JOHNSON,K.L. T r a c t i o n s i n E l a s t o -\nF r i c t i o n a n d t h e Wear R a t e o f Meta ls . h y d r o d y n a m i c s L u b r i c a t -\ni o n .\nP r o c . I . Mech.E., 1 8 2 , B u t t e r w o r t h P r e s s , P t . l , N o . 1 4 , 2 8 1 , 1 9 7 0 . L o n d o n , 1 9 6 5 .\n1 4 . K R A G E L S K I I , I . V . F r i c t i o n a n d Wear\nTIMOTHENK0,S T h e o r y of E l a s t i c i t y 1 5 . HALLIDAY,J.S. S u r f a c e E x a m i n a t i o n b y & G O O D I E R , J . N . M c G r a w - H i l l , 1 9 5 1 . R e f l e c t i o n E l e c t r o n\nMicros C O D Y - 10. HALLING,J. & A S t a t i s t i c a l M o d e l f o r\nEL-REFAIE,M E n g i n e e r i n g S u r f a c e s I n s t n . Mech. E n g r s . T r i b o l o g y C o n v e n t i o n , 1 9 7 1 .\nProc .1 .Mech.E. 1 9 6 9 , 7 7 7 , 1 9 5 5 .\n11. G R E E N W O O D , J . A . C o n t a c t s Of N o m i n a l l y This paper is presented for written discussion. The MS was received on & WILLIAMSON, F l a t S u r f a c e s . J .B .P . 1st November 1975 and accepted forpublication on 10th June 1976. 33\nP r o c . Roy. SOC. , A , 2 9 5 Communications are invited for publication in the hceedings. 3 0 0 , 1 9 6 6 . Contributors should read the instructions on page ii of cover.\nRoc Instn Mech Engrs Vol190 43/76 IMechE 1976\nat WEST VIRGINA UNIV on June 5, 2016pme.sagepub.comDownloaded from",
            "Roc Imtn Mech Engra Voll90 43/76\nat WEST VIRGINA UNIV on June 5, 2016pme.sagepub.comDownloaded from"
        ]
    }
]