diff --git "a/designv11-65.json" "b/designv11-65.json" new file mode 100644--- /dev/null +++ "b/designv11-65.json" @@ -0,0 +1,8430 @@ +[ + { + "image_filename": "designv11_65_0000495_j.1467-8667.1993.tb00215.x-Figure2-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000495_j.1467-8667.1993.tb00215.x-Figure2-1.png", + "caption": "Fig. 2. Generalized flux tube ge0metry.l' Fig. 3. Flat disc narrowing model.", + "texts": [ + " The representative spot thermal resistance is calculated considering around each spot a heat flux tube having a narrowing in the contact zone. The ideal geometry of the representative narrowing can influence considerably the effectiveness of the model. Different shapes of the narrowing have been investigated'.2.J.x*'3 and compared in Ref. 11. The heat flux problem with the boundary condition appropriate for the chosen geometry. The boundary conditions for the most general geometry have been proposed in Ref. 11, using an ancillary reference system (see Fig. 2). The contact between two truncated cones with variable angle is considered. In particular choosing the heights of both cones equal to zero we obtain the solution for the flat disc narrowing geometry (Fig. 3). Within the described hypotheses the boundary conditions are as follows a u r = 0 z > o r = b z > d a T -=() ar a T ar - = O Constitutive laws for a contact element 303 where k is the material conductivity, 4 is the heat flux and geometrical symbols a, b, 6, u, are described in Fig. 2. These conditions impose, respectively, a certain heat flux distribution in the contact zone, no heat exchange on the inclined surface of the cone, radial symmetry condition along the flux tube axis, no heat exchange on the flux tube vertical surface and a uniform heat flux distribution within the section at a certain distance from the narrowing. The definition of the function fir), describing the heat flux distribution in the contact zone is crucial for determining the solution of the problem. Hypotheses usually adopted consider a constant heat flux distribution in the contact zone 4 x a f ( r ) =2 Another very effective hypothesis considers the zone of contact as symmetry plane, hence the first boundary condition should be replaced by a condition imposing a uniform temperature value, T,,, on the contact zone In such a way we have a set of mixed boundary conditions which add new difficulties to the solution of the problem" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003775_3-540-29461-9_91-Figure5-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003775_3-540-29461-9_91-Figure5-1.png", + "caption": "Fig. 5. Suggested suction cup arrangements on a foot", + "texts": [ + " Due to that, elasticity is introduced into the robot\u2019s kinematics by standard passive suction cups in two ways. First, an elastic suction cup is prone to bending. Second, a cup increases its height depending on the pulling force. Hence, the robot has no rigid foothold which introduces elasticity into the kinematic structure and may cause control problems. But a simple constructive method eliminates the first problem and reduces the second one, namely to arrange multiple suction cups on a single foot in a distance larger than the diameter of a single suction cup (see Fig. 5). This distance acts as a rigid coupling which avoids bending and also like a lever which reduces the effective variation of suction cup height. Because severe bending is avoided, the danger of damaging a suction cup is also significantly reduced such that large torques and forces are compensated very effectively. Three passive suction cups (Fig. 5a) may be connected by short straps (black in Fig. 5). Hence, only a short stroke is needed to release the vacuum. In contrast, if four suction cups (Fig. 5b) are connected at one central point in order to simplify the lifting mechanism, the straps and hence the stroke have to be longer. The advantage of using four cups are an increased total adhesive force and stability against bending. Multiple suction cups are also useful to improve the safety of the robot by redundancy as well as by increasing the safety factor of the achievable adhesive force. Due to the simplicity and the inexpensiveness of passive suction cups, this is nearly free of additional costs and weight. For instance, a single suction cup used by DEXTER is capable of holding eight times the weight of the robot and has a weight of a few grams. And DEXTER has three of them on each foot (in an arrangement according to Fig. 5a). Of course, the problem of passive suction cups loosening by the time, only occurs when the robot is not in steady movement. Fortunately, climbing robots use a dexterous locomotion system. This allows a foot to be kept on the same place for a longer time by simply pressing it again and again to the surface in order to reevacuate the cup without moving the robot. An ordinary, light-weight RC-Servo plus a simple tearing mechanism, e.g. a rope, replace at least one vacuum pump, valves, tubes, muffles etc" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003890_imc.1990.687360-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003890_imc.1990.687360-Figure1-1.png", + "caption": "Figure 1 : Example of the determination of free sectors", + "texts": [ + "Thestepofdigitizationonthisaxisis: Ae\u2019 5; - 2 . Our goal is to obtain at least one free sector between two obstacles. A free sector will correspond in the histogram at the distance between two main maximum separated by an empty zone. If an empty zone exists, it is defined by 81 and eZ with the *-el> A@. where Wi is the weighting matrix of the i* measure. We have: We solve an equation of Constraints: 8, -KA8 5 el,, with K e N . sectorS for a simulated example; two obstacles were detected and the grid i=n Figure 1 is an illustration of the determination of free and occupied Wi'= [ Z(Ri)-'r'(RJ-' 1 h: was placed in order to obtain a h e path between the two obstacles. i l t where R = [ X(Rd-']-' is the covariance matrix of the estimation error i=l @ = p - d . When the vehicle moves f\" a position to another one, we want to establish a relation between the informations obtained at this two positions. Spatial Fusion At time k, the matching is possible only between high level informations @lanes, volw-;,, ...)", + " Cq = 0 be differentiated with respect to time, we ~ allows to rewrite the system (1): We can now derived a oneorder state equation representing the robot motion, and including all e mechanical constraints: n null matrix. form of the motion law: r obtained thru state feedback and representing by an artificial potential field created by I procedure,ithasbeenpo formalizetheMotionPlanning Problem for real robots as an Adaptiv lproblem. Thesedifferentmodules arenow Our approach consists in adapting the potential field in order to maintain a performance index as near as possible to a reference value (good behavior of the state trajectory for instance). This functionis devoted to a Copilot (Figure 1). Concerning the numerical integration of this model, let us choose a nxp m a ~ x A, composed of p column vectors in the null space of C, we can write Ch=O,. By multiplying the first equation of (1) by AT we obtain ArMq = ArF . Combining this equation with the second equation of (1) we obtain: Unfortunately, the errors due to this numerical integration are the causes z oftheviolationoftheconstraintequation(secondequationof(l)).Therefore Inorderto avoidthiserroraccumulation, wecanmodified thisschemeby it is impossible" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003261_2001-gt-0255-Figure2-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003261_2001-gt-0255-Figure2-1.png", + "caption": "Figure 2: 5DOF model parameters", + "texts": [ + " The inner race co-ordinate system (x, y, z, \u03b8x, \u03b8y) corresponds to the DOF of the rotor at this point and has its origin at the center of the bearing. The local rolling element co-ordinate system (r, Z, \u0398) defines the position of the inner race center of curvature for each rolling element and has its origin at the nominal position of the inner race center of curvature. The outer race center of curvature is used as a 3 loaded From: http://proceedings.asmedigitalcollection.asme.org/ on 02/03/2016 Term reference point. Figure 2 shows the various parameters used. Note that rp, the radial co-ordinate of the nominal center of curvature of the inner race, is not related to Dp. For the j\u2019th ball, the displacement of the inner race center of curvature, u, is related to the displacement of the inner race, d (treated as a known input quantity for the forcing function) according to: { } [ ] [ ]{ }TT r Z x yu u u x y z \u0398= = = \u03b8 \u03b8u R d R (9) where, [ ] j j p j p j p j p j j j 3 5 cos sin 0 z sin z cos 0 0 1 r sin r cos 0 0 0 sin cos \u00d7 \u03c6 \u03c6 \u2212 \u03c6 \u03c6 = \u03c6 \u2212 \u03c6 \u2212 \u03c6 \u03c6 R (10) All surfaces are assumed to be curves with only one radius of curvature" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002390_icsyse.1990.203207-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002390_icsyse.1990.203207-Figure1-1.png", + "caption": "Figure 1: Coordinate frames for crank turning task", + "texts": [ + "00 1990 IEEE forceGains: force control gains in COMPLY frame maxForceVel: maximum velocities in COMPLY frame DOFs springGains: position and orientation gains to pull COM- maxSpringVe1: maximum velocities due to virtual springs due t o compliance PLY frame back t o nominal motion trajectory Termination Condition Parameters: select: bit mask selecting which termination conditions to test for testTime: time over which to average termination conditions endTime: maximum time to comply and test termination conditions a t end of nominal motion endTransErr: translation error in COMPLY frame due to compliance endAngErr: orientation error in COMPLY frame due to compliance endTransVe1: time derivative of endTransErr end AngVel: time derivative of endAngErr endForceErr: COMPLY frame force error vector magnitude endTorqueErr: COMPLY frame torque error vector magni- endForceVe1: time derivative of endForceErr endTorqueVe1: time derivative of endTorqueErr tude Safety Parameters: forceThreshold: force threshold magnitude to stop on torqueThreshold: torque threshold magnitude to stop on positionThreshold: position threshold magnitude to stop on OrientationThreshold: orientation threshold magnitude to jointMargin: angular distance from joint limits and singustop on larities t o stop on The kinematic relationships between the WORLD frame, the robot, and the task motion and control frames are specified in the input parameters. These kinematic relationships are shown in figure 1. The manipulator position and motion destination are specified with respect t o a fixed Cartesian frame, WORLD. The transform trBase specifies the position of the manipulator base, the BASE frame, with respect to the WORLD frame. The transform trTn is the transform from the BASE frame to a frame fixed in the manipulator's terminal link, the TN frame, and is updated during task execution from joint angles by forward kinematics. The trNom transform is the transform from the TN frame to the nominal motion frame, NOM, where Cartesian interpolated motion is t o occur", + " ON-TERM-COND: stopped on satisfied termination conditions ON-TIME: timeout a t end of end motion - termination conditions not satisfied ON-FORCE-THR: stopped on force threshold ON-TORQUE-THR: stopped on torque threshold ON-POSITION-THR: stopped on position threshold ON-ORIENTATION-THR: stopped on orientation threshold ON-JOINT-THR: stopped on a joint threshold ENVIRONMENT XI. APPLICATIONS The Generalized-Compliant-Motion primitive has been utilized on the JPL Telerobot for many different compliant motion tasks. These tasks include pin insertion and removal, leveling grippers on grapple lugs, bolt seating and turning, pushing, sliding, space truss assembly, door opening, crank turning, and contour following. Figure 1 shows several of the parameters for turning a crank. Many more applications will be developed in future work as will enhancements to the primitive. XII. CONCLUSIONS The Generalized-Compliant-Motion primitive provides a planning level system a means to execute a planned compliant motion task. Its general functionality and large input parameter set allow i t t o be used for execution of a wide variety of compliant motion tasks. ACKNOWLEDGEMENTS The research described in this paper was performed at the Jet Propulsion Laboratory, California Institute of Technology, under contract with the National Aeronautics and Space Administration" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000302_cp:19991046-Figure6-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000302_cp:19991046-Figure6-1.png", + "caption": "Figure 6 Voltage vector in SRRF aligned with h,(k)", + "texts": [ + " In addition errors in R, and is will affect the accuracy of flux estimation. For input voltage vector selection, theBj s should be in stator field coordinate. As the stator flux vector at the k-th sampling time instant can be estimated using Eqn.(l) and expressed as - - - X,(k) =h,(k)e'Ok, the transformation is based on 8, . Thus i j , e jek = [vai t jvB,]cos e, t jsin e, 1 = lvai cos Ok + vfi sin 0, I t j l v p j cos o k i var sin ek J = vd, + jvsi The two orthogonal components, vd, and vq,, are parallel to and orthogonal to h,(k) respectively as shown in Figure 6. Whilst vd, determines the magnitude variation of h,(k), vqi changes its phase angle O k . Thus applying iji will increase h,(k) if vd, >O and decrease it when vdi O and decreased if vq; O (vdi 2 (see Ref", + " If, in addition, the protons released by these protonated groups to the SFR radicals are not completely replaced by other protons coming from the solution during the electrolysis time, the resulting negatively charged phosphate groups will create a local negative potential which may slow down the translocation of the SFR cations from the polar head region to the mercury surface. This attractive interaction between SFR cations and negatively charged polar heads will gradually increase as Qf approaches its maximum value Qf, max along a given Q(E,t= 50ms) vs. E curve. This may explain why the anticipation in SFR reduction, as observed in DOPS and DOPA films with respect to DOPC films, is confined to the lower portion of the Q(E,t = 50ms) vs. E plots over the pH range from 3.5 to 4.5 (see Fig. 3). The authors wish to thank Mr. Luciano Righeschi for valuable technical assistance. Thanks are due to the Ministerio de Educaci6n y Ciencia of Spain for a fellowship to R.H. during the tenure of which most of the present results were obtained. The financial support of the Ministero deli'Universita' e della Ricerca Scientifica e Tecnologica and of the Consiglio Nazionale delle Ricerche is gratefully acknowledged. [i] R.C. Prince, SJ.G. Linkletter and P.L. Dutton, Biochim. Biophys. Acta, 635 (1981) 132" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000834_000147919-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000834_000147919-Figure1-1.png", + "caption": "Fig. 1. Experimental setup. After the release of the platform by a sudden horizontal pull on the support, the platform (and differently, but not independently the subject) falls down. After a falling distance z of 55 mm. the fall of the platform is limited by a damper. All mea surements and calculations refer to the first 33 mm of free full. Dis tance and acceleration of the platform against the ground arc mea sured on the center line of the construction (dash-dotted line).", + "texts": [], + "surrounding_texts": [ + "Key Words Biomechanics Morphology, functional Elasticity Energy Muscle\nIntroduction\nEnergy storage in elastic elements of the locomotor ap paratus has been postulated to be one of the major domi nants of body shape and mechanical functions in a variety of species [cf. the overview in Alexander, 19881. The scal ing of vertebrate body proportions for many species can be described by \u2018elastic allometry' and \u2018elastically similar models\u2019 [McMahon. 1975. 19771. For some species the lo comotor modes, like for example the hopping of a kangaroo\nK A R G E R \u00a9 1997 S. Karger AG. Basel 0001-5180/97/1582-0106$ 12.00/0 I \u2022 - Mai I karger

10 mm would have been an exclusion criterion: none of the volunteers or measure ments had to be excluded on the basis of this criterion. The volunteers were not aware of the release of the platform. They were facing the diffuse, monochrome painting of a wall. She or he could not see the person releasing the platform, and could not hear the preparations for the fall, because the steel rope was tightened before and the support re leased by a sway of the \u2018releasor's' body and hand with the hand grip at the other end of the rope. Thus, the possible time available for a re action was approximately zero.\nIf the human body and the platform had fallen down as two inde pendent. stiff bodies, driven only by gravity g (air resistance, friction and other dissipative factors arc supposed to be negligible), the falling time t for a given distance z for both the human body and the platform should be\nl = ^ f ( I )\n(cf. Galilei. 1638. pp 75-76. 85-86: Newton, 1713. pp5-6 |.\nResults\nThe platform without any additional load fell the dis tance of 33 mm within 82.5 \u00b10.17 ms (n=20). Loaded with an additional, stiff external mass of 50 kg, the platform fell within falling times of 82.6\u00b10.17 ms. With 50 kg of wa ter (capsuled in cans), the average falling time was 81.8 \u00b10.18 ms (n=20). These measurements correspond rather well to the value predicted by (1):\nt(z = 0.033m) = a / 111 = V ~2 ' ?-3- !\u00b0 ~7 \u2122 ~82\u2022 10 's (la) v v -9.81 ms-\nWhen the volunteers stood on the platform, it fell reproducibly faster and the human subjects fell more slowly than calculated by the above formula which is based on the laws of stiff-body physics. In the upper part of figure 2, an ex ample of a plot of the vertical coordinate of the platform is\nshown. Time counting is started when the support is re leased. The platform continuously falls faster than pre dicted and measured without the volunteers.\nAs could be seen above, for stiff bodies, corresponding to a falling height of 33 mm. a falling time of 82 ms can be predicted. In contrast to this prognosis, we measured an av erage falling time of the platform of 50.1 \u00b10.46 ms (n = 50). Evidently, the platform is significantly accelerated in addi tion to the effects of gravity (Student\u2019s t-test: p <0.001). The average acceleration a of the platform can be derived from ( 1):\n2-z 2 (-0.033 in) \u201e , , m = ---------------- = -26.3 t: (0.0501 s)-\u2019 s\u2019 ( 2)\nIf we suppose this acceleration to act constantly during the whole fall (cf. fig. 2), the velocity v of the platform after a fall of z=-33 mm can be calculated by\nv = a \u2022 l = -26.3 \u2014\u2022 0.0501 s = - l .32 \u2014 (3) s- s\nThe platform contains a total energy E of\nE = EI,H+ E u\u201e= n v g -z + T m v % (4)\nwhere E^, is the potential energy and Ek,\u201e is the kinetic energy of the mass m, which is situated in gravity g at the vertical coordinate z with the velocity v. Before the release of the platform, it has a total energy of\n\u00a3 = \u00a3,*\u201e + Ekm = 40.1 kg \u2022 9.81 \u2122 \u2022 0 m + (\u25a0 40.1 kg \u2022 0 ^ \u2019 = 0 J (4a)\nAfter a fall of 33 mm. the total energy E of the platform is:\nE=Er\u00ab + Eta = 40.1 kg -9.81 - (-0.033 m)\u00b1T40.1 kg-1.32\u2019\u2014 =21.8 J (4b) s3 s3 The total energy of the platform increases, energy is\ntransferred from the human body to the platform. Our mea surements indicate a minimum energy transfer of more than 20 J within the first 50 ms, equivalently to an average pow er flow of 400 W from the body to the platform. On the basis of additional measurements of accelerations on diverse body surface landmarks like trochanter maior. os sacrum or vertex, the compensatory deceleration of the human body to a falling time of more than 82 ms can also constantly be observed, but as a consequence of the human body\u2019s obvi ous deviations from the ideal of mechanical stiffness and the different body masses of the subjects, an exact and re producible quantification of this effect for the center of gravity is difficult.\n108 Acla Anal 1997:158:106-111 Witte/Recknagel/Rao/Wiithrich/Lesch\nD ow\nnl oa\nde d\nby :\nK in\ng' s\nC ol\nle ge\nL on\ndo n\n13\n7. 73\n.1 44\n.1 38\n- 1\n/2 9/\n20 19\n4 :4\n7: 45\nP M" + ] + }, + { + "image_filename": "designv11_65_0002222_isie.2001.931567-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002222_isie.2001.931567-Figure1-1.png", + "caption": "Fig. 1. Cross section of the motor", + "texts": [ + " The Motor The three-phase 12/8 structure Switched Reluctance motor has also the structure of the double-salient poles. There are 12 teeth poles and 12 slots in the stator, and there are 8 teeth poles and 8 slots in the rotor. The two coils on the diametrically opposite stator teeth poles could be connected to make up a winding, the two vertical windings could be connected to become one phase winding in the stator. There are no brush, no magnet and no windings on the rotor. The cross section of the motor is shown in Fig.1. The rotor position detector could be installed on the no shaft extension end, which consists of the slotted disk and the photoelectric transducers. The slotted disk is fixed coaxial with the rotor, and there are three photoelectric transducers that are fixed at the case of the motor. There are eight teeth with 22.5\" widths per tooth and eight slots with 22.5\" widths per slot in the slotted disk. The three photoelectric ISIE 2001, Pusan, KOREA transducers are installed with a (15' +e ) interval" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002910_tt.3020080302-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002910_tt.3020080302-Figure1-1.png", + "caption": "Figure 1 Coordinate system in the bearing clearance", + "texts": [ + " On the basis of Stokes' couple-stress fluid model, the present paper considers the effects of couple stresses and inertia on the nature of squeeze-film behaviour in curvilinear thrust bearings with reference to synovial joints. It is hoped that such an analysis can aid in understanding the mechanism of human joint lubrication and the role of long-chain hyaluronic acid molecules in synovial fluids behaving as couple-stress fluids. ~~~ ~ TriDot&joirriin18-3, Mnrcli 2002. (8) 196 ISSN 1354-4063 $10.00 + $10.00 A bearing flow configuration is shown in Figure 1. The fixed surface is described by the function R(x) , which denotes the radius of the surface. The fluid film thickness in the bearing clearance is described by the function Iz(x,t), which denotes the distance between the fixed lower surface and the moving upper surface, measured normal to the fixed surface. An intrinsic curvilinear orthogonal coordinate system x,8,y linked with the fixed surface is shown in Figure 1. The field equations for the motion of an incompressible fluid with couple stresses are: div V = 0 (1) dV = -gradp+pV2V-qV4V P d t where V is the velocity vector, p the pressure, p the density, p the shear viscosity, and 17 the couple-stress viscosity. The physical parameters of the flow are the velocity components D,, uY and pressure p . With respect to the axial symmetry, these parameters are independent of the angle 8. After reduction allowed by assuming axial symmetry, and assuming that h ( ~ , t ) e R(x) the governing differential equations are:13p15 where the prime denotes differentiation with respect to x" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003605_mcs.2004.1337851-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003605_mcs.2004.1337851-Figure1-1.png", + "caption": "Figure 1. (a) Overhead view of the static friction coefficient testbed. (b) Detailed view of IR pair that detects movement of the test object. The test mass is placed on a plate whose inclination angle is varied by the dc motor. The IR receiver-transmitter pair detects the instant at which the test mass begins to slide. The corresponding inclination angle, which is measured by means of a potentiometer, is used to compute the coefficient of static friction.", + "texts": [ + " Next, we briefly describe two sample prototype mechatronics-enabled science experiments developed under the SMART program. Additional prototype science experiments developed under the program are described in Table 2 and [5], [9]\u2013[11]. October 2004 27IEEE Control Systems Magazine This testbed is designed to experimentally determine the coefficient of static friction between various surfaces. It consists of a horizontal base and a plate whose inclination can be varied using an H-bridge-controlled dc motor (Figure 1). A cube-shaped mass with four different types of surfaces serves as a test object. In particular, experiments are conducted to determine the coefficient of static friction between the plate and any one of the faces of the test object. An infrared (IR) transmitter and receiver are mounted on the opposing sidewalls of the plate. This IR transmitter-receiver pair is used to detect movement of the test object. A reference wall is placed at the free end of the plate so that, by placing the test object snug against the plate, the IR beam is tripped at the instant the test object begins to slide" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003918_j.jmatprotec.2005.02.163-Figure4-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003918_j.jmatprotec.2005.02.163-Figure4-1.png", + "caption": "Fig. 4. Used testing device.", + "texts": [ + " The testing devices are built in modules, which allow small manipulations without reaching the inside. The new generation of testing devices does not disturb elements above the surface while the horizontal arrangement avoids damages as well. During the production, the network remains in the horizontal position. This configuration allows a simpler attachment to the testing site. One of these new testing devices were the basis for this project. The system used to fix the plug and the other functions were analysed and documented. Fig. 4 gives an overview of the used parts. Fig. 5 shows the principle of the simplest testing device. The plug is fixed and the electrical connections are tested by appropriate software. If the light of the combined switch is on, then the testing device is ready to use. The plug is put through the front platform into the inlet. As soon as the plug reaches the switching contact, a signal goes to the electro-pneumatic valve and the cylinder presses the closing element, in this case a bar, into the empty space of the inlet and the plug is fixed by the bar" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002720_robot.1992.220099-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002720_robot.1992.220099-Figure1-1.png", + "caption": "Figure 1: Collision events defined from the shadow.", + "texts": [ + " Consider the corresponding robot coordinates DL and DA+l in the target frame of reference. Inertial bounds on the robot and target motion enable us to make a differential approximation and consider the robot trajectory during a time interval in the frame of reference of the target as arbitrarily close to a line segment Let us define the shadow of the target with respect to a point D as the workspace subset hidden by the target to an observer in D. We denote this subset by Sh(D). The collision probability associated with the robot displacement is (Figure 1) 2.2 Target Motion State Distribution Assuming the target motion is smooth enough, the kinematic state evolution of the target is modeled locally as a first order Markov time invariant model comprising the pose and attitude parameters and their first and second derivatives. S n + l = ASn + Vn (1) where we have: Sn = [U,, @,IT the target motion state vector, U,, = [Xn , in, $n] where Xn, kn, Z n are respectively the position, velocity and acceleration vectors of the target center of inertia, Xn = [zn,ynIT the position of the target center of inertia, 244 I 0 0, = [ C Y , , ~ , , & , ] ~ the target orientation angle 0 V, = [Vf, V:]' the error on the target motion model with zero mean Gaussian distribution and covariance equal to Tn , and independent , and its first and second derivatives, 0 A = the model transition matrix" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002472_b978-0-12-464965-1.50013-2-Figure8.2-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002472_b978-0-12-464965-1.50013-2-Figure8.2-1.png", + "caption": "Figure 8.2 Schematic diagram of a MOSFET", + "texts": [ + " Elec tronic aspects and chemical aspects of ISFETs and ChemFETs will be 325 326 Principles of ChemFET Operation 8 presented in Sections 8.1 and 8.2. We shall often use a Si02-based FET as a model in our discussions (although of little or no practical value), because it is a simple structure and allows us to illustrate the problems associated with ChemFETs in general. To understand the operating principles of the chemically sensitive FET devices, we will analyze the operation of an IGFET or a MOSFET, a FET with a metal layer (gate) over the oxide. A diagram of an IGFET is shown in Fig. 8.2. The resistance of the \"channel\" between the two n + regions is measured. This resistance is very sensitive to the potential applied to the gate electrode. The basic device parameters are the channel length L, which is the distance between the two n + p junctions (source and drain), the channel width \u0396 and the insulator thickness d. The example given is an \"^-channel\" device, \"^-channel\" because the inversion layer in this />-type silicon conducts by electrons. So the surface becomes \u00ab-type. The inversion layer in this device forms when a sufficiently large positive bias is applied to the gate (VG) of the structure shown" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001949_j.1749-6632.1986.tb34500.x-Figure4-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001949_j.1749-6632.1986.tb34500.x-Figure4-1.png", + "caption": "FIGURE 4. The unit cell of a filament. Dashed line: right-handed filament. Solid line: leftlhanded filament. The two unit cells, which are extracted from a helical net drawn at a radius of 70 A, are very similar. The change in twist manifests itself in the 7-,4 shift of the upper vertex. Each arrow denotes the length of the unit cell edge to the upper vertex.", + "texts": [ + " Calladine (1976)\u2019 predicted that the R and L transition would include a shortening by 0.8 A of the distance between subunits along the eleven-start rows and a transverse shift of 8 A in the intersubunit bonding. The former change would be responsible for the corkscrew shape of the filament and the latter for the twist. The first attempt to study these changes was made by Kamiya et af. (1979)\u201d using X ray diffraction. Their lattice results indicate a transverse shift of about 7 A as predicted, but a much smaller shortening than predicted of only 0.2 A. Our results (FIGURE 4), which are still preliminary, are similar regarding the transverse shift. We find, however, a 0.7-A t 0.3-A shortening, which is more in line with Calladine\u2019s prediction. Thus, the measured changes in the helical lattice are in agreement with Calladine\u2019s predictions. Salmonella typhimurium S J W 1 6 6 0 S J W 1 6 5 5 FIGURE 3. Diffraction patterns of images of the filaments in FIGURE 2. Lefi side: transform of a left-handed filament. Right side: transform of a right-handed filament. The order of each of the layer lines is indicated to the side" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000777_s0165-0114(96)00283-7-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000777_s0165-0114(96)00283-7-Figure1-1.png", + "caption": "Fig. 1. The applied membership function.", + "texts": [ + " Consider a PD-type discrete-time fuzzy controller whose inputs e(k) and ce(k) are defined as e(k)=sp(k)-y(k) , (1) c e ( k ) = e ( k - 1 ) - e ( k ) = y ( k ) - y ( k - 1), (2) where sp and y denote the applied set point and the plant output, and indices k and k - 1 indicate the present state and the previous state of the discrete-time system. The control rules with two inputs and a single output fuzzy variables E, CE and U representing the input and output variables e, ce and u of the controller are shown in Table 1. Seven fuzzy sets, applied for all the three fuzzy variables in the same way, are represented by the membership functions in Fig. 1. Among these fuzzy sets, the characters N, P, B, M, S denote negative, positive, big, medium and small, respectively, and ZE represents zero. Though the control rules in Table i, whose creations are based on the characteristic of the step response, state the state-action relationship in a linear way, adjusting the scaling factors and the control rules can result in considerable nonlinearity while applying them. During the control procedure, once the present plant output y(k) results, the values e(k) and ce(k) calculated by (1) and (2) are scaled by scaling factors (SFs), namely GE(k) and GCE(k), respectively", + " Figs. 3(a)-(c) depict the corresponding trajectories, in which case 1 is evidently the desired one. The trajectory in Fig. 3(a), which is fixed if c~, ri and the initial values are given, is called the original reference model. To generate a reference trajectory in the linguistic space, we notice that the input variables of the controller are E (output error) and CE (change in output error), and the universal discourses of the input fuzzy variables E and CE defined by the membership functions in Fig. 1 are both [ - 6, 6]. So, the initial values of the reference trajectory should coincide with this interval. Now, consider the discrete forms of (8a) and (8b) with zl and z2 on the right-hand side of the equations replaced by e and ce, which leads to Zl(k + 1) = (1 + ~zT)e(k) + riTce(k), (lOa) zz(k + 1) = - riTe(k) + (1 + aT)ce(k), (lOb) where e(k) and ce(k) are the present error and change in error, zl(k + 1) and z2(k + 1) denote the desired error and change in error in the next control step, and T is the sampling interval" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003131_naecon.2000.894959-Figure3-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003131_naecon.2000.894959-Figure3-1.png", + "caption": "Figure 3: PMSM and different coordinators", + "texts": [ + ", phase A, consist of three current ripples. During the time interval At, under Vc, the current change is Ai,; during the time interval At , under V d , the current change is ai,; and during the time interval At11 + At12 under zero-vector, the current change is Ail1 + 4 2 1 2 . If these current ripples are measured, the rotor position dependent and speed-proportional back-EMFs can be solved using the general PM machine model with the known voltage vectors applied. n L B. Algorithm A basic 2-pole PM synchronous machine is shown in Fig. 3, where as, bs, and cs axes are the stationary axes for the three phases. X-Y is the equivalent twophase stationary coordinate. In the PM machine, the d-axis is aligned with the North pole of the permanent magnet of the rotor and q-axis 90\" ahead of d-axis of the machine. In d-q coordinate, the voltage + [ ,& 1 Using transformation of X-Y to d-q as case sine In discrete form, the equation at k-th moment will be + [ :jkk; ] (5) where, T, is the sampling time period. The inductances [L] and back-EMFs [e] are assumed constants within one PWM cycle T," + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002993_1.1737377-Figure3-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002993_1.1737377-Figure3-1.png", + "caption": "Fig. 3 Planar 4 bar mechanism in a singular configuration", + "texts": [ + " Choose Ji , j such that d5dim Ci , j is the minimum of all possible choices 3! For the optimal choice, i.e. d is minimum a! compute SVD of CP5UTSV b! construct reduced matrix U\u0304 \u21d2Result: reduced coefficient matrix U\u0304PLPRd21,n The final result is a system of d21 velocity constraints, with d the lowest possible dimension of the constraint space. The following examples only contain revolute joints and the joint frames are such that the revolute axes are along the e3 axes of the reference frames. Planar 4 Bar Mechanism. The planar 4 bar mechanism in Fig. 3 is a standard example for the occurrence of dependent constraints. Any revolute joint chosen as cut joint yields 5 con- Transactions of the ASME 17 Terms of Use: http://www.asme.org/about-asme/terms-of-use Downloaded F straints of which 3 are identically zero. Evaluating the constraint space shows that Ci , j5se(2), the algebra of planar motions, for all possible cut joints Ji , j . That is, d5dim se(2)53. So d21 52 constraints are sufficient for the 4 bar mechanism. The GKC formula holds and the mechanism DOF is n2(d21)532251" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001186_s0007-8506(07)62882-0-Figure3-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001186_s0007-8506(07)62882-0-Figure3-1.png", + "caption": "Fig. 3. Schematic v i e w of expertmental setup for combined process of LCEMA. 1 - surface t o be cladded; 2 - alloying powder feeder nozzle, 3 - work ing gas; 4 - electro - magnet ic (EM) coil; 5 - focusing lens; 6 - EM power supply; 7 - laser beam; 8 - electric arc power supply, 9 - arc electrode - work ing gas nozzle; 1 0 - cladded layer: 1 1 - electric arc.", + "texts": [], + "surrounding_texts": [ + "( 1 ) Draper C., 1 9 8 2 , Surface Al loying: the State of A r t , Journal of Meta ls , 34 ,4 : 2 4 - 3 2 (2) Heuvelman C.J., et al, 1 9 9 2 , Surface Treatment Techniques by Laser Beam Machining, Annals of t he (3) Kovalenko V., Golovko L., 1 9 9 3 , Laser ThermoDeformat ion Mater ia l Hardening, Proc. Int. Conf Electron Beam and Laser Proc., Reno, USA. (4) Kovalenko V. et al, 1 9 9 3 , Wear resistant Laser Cladding with Boron Containing Powder Steel, Proc. ICALE0'93, Orlando, USA. ( 5 ) Kovalenko V., Haskin V., 1 9 9 5 , The Selection of Self-Fluxing Powder Mater ia ls fo r Laser Cladding, lnformat izat ion and N e w Technology, 1 : 3 6 - 3 9 . 9 9 9 (6) Kovalenko V., Lutay A., Anyakin M., Gas Powder Laser Cladding with Electro-Magnetic Agi tat ion, 1 9 9 7 , Proc. of ICALE0 '97 , San Diego, USA. (7) Mazunder J., Conde O., Villar R., Steen W., (eds.), 1 9 9 4 , Proc. of N A T O ASI, Sessimbra, Portugal CIRP, 4112: 6 5 7 - 6 6 6" + ] + }, + { + "image_filename": "designv11_65_0002353_iros.2000.895310-Figure2-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002353_iros.2000.895310-Figure2-1.png", + "caption": "Fig. 2. Coordinate systems.", + "texts": [ + "2 Coordinates systems and parameters Here we are concentrating on a situation, where the cameras are fixed, and the calibration object is attached into the robot hand (tcp). We do a set of calibration motions, and collect the corresponding pose values HI,, as well as the locations of the calibration point in the image planes, pimj. , j = 1, 2, ..., n denoting the camera. In addition, we need a fare initial guess for the hand-eye parameters HS,,, as well as the location of the calibration point in the robot hand, p I . .Fig. 2 illustrates this situation for one camera (to simplify the notation, we will leave out the index of camera j for a while). We have the following parameters for the estimated camera pose 5 : - 8 = ( c x c y cz ' ro l l 'pitch c y w ) (1) where c x , c,, G, give the location and crOl/, cpilch and cy,, give the rotation of the camera For the position of calibration point we set - PI = ( P x P y P z ) where pn p p p z , give the location in tool frame. The state vector (for cameras I to n) is then composed of camera pose parameters and the location of the calibration point: - 2293 - - - ; = @ I " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003437_0954406042690498-Figure8-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003437_0954406042690498-Figure8-1.png", + "caption": "Fig. 8 Pre-designed parabolic-like kinematic error function with magnitude x", + "texts": [ + " Mechanical Engineering Science C04304 # IMechE 2004 equation \u00bdJ 767 du=df0 1 db=df0 1 dv=df0 1 dy=df0 1 df1=df 0 1 df2=df 0 1 df0 2=df 0 1 2 666666664 3 777777775 761 \u00bc qF1=qf 0 1 qF2=qf 0 1 qF3=qf 0 1 qF4=qf 0 1 qF5=qf 0 1 qF6=qf 0 1 qF7=qf 0 1 2 666666664 3 777777775 761 \u00f046\u00de where \u00bdJ 767 is the Jacobian matrix and \u00bdJ 767 \u00bc q\u00f0F1,F2,F3,F4,F5,F6,F7\u00de q\u00f0u, b, v, y,f1,f2,f 0 2\u00de Substituting equations (44) and (45) into equations (42) and (43) yields the specific slidings of pinion and gear tooth surfaces respectively. The gear set discussed in this paper is provided with predesigned parabolic-like kinematic error as shown in Fig. 8. The magnitude of the parabolic-like kinematic error, denoted by x, is defined as the difference between the maximum and minimum values in a period of 2p=N1, where N1 denotes the number of pinion teeth. Within this period, only a single pinion tooth is in contact with another single gear tooth. Numerical settings for parameters of cutting tools determine entirely the magnitude of the parabolic-like kinematic error and the dimensions of elliptical contact areas. To acquire a certain magnitude and dimensions, the designer may vary the numerical settings of cutting tools and find out whether their expectations can be satisfied. However, this trial-and-error procedure is very inefficient. For automatically determining the numerical settings of cutting tools from a computer program, a system of governing equations is proposed. As shown in Fig. 8, the middle kinematic error curve crosses the leftand right-hand sides at f0 1 \u00bc jL and f0 1 \u00bc jR respec- tively. When meshing surfaces contact at f0 1 \u00bc jL, all parameters are assumed so that f0 1 \u00bc jL, b \u00bc bL, y \u00bc yL, f1 \u00bc f1L, f2 \u00bc f2L, u \u00bc uL, v \u00bc vL and f0 2 \u00bc xp=648 000\u00fe \u00f0N1=N2\u00dejL. Here, the constant 648 000 is used to balance units since x is in arcseconds and f0 2 and jL are in radians. When meshing surfaces contact at f0 1 \u00bc jR \u00bc 2p=N1 \u00fe jL, all parameters are assumed so that b \u00bc bR, y \u00bc yR, f1 \u00bc f1R, f2 \u00bc f2R, u \u00bc uR, v \u00bc vR and f0 2 \u00bc xp=648 000\u00fe \u00f0N1=N2\u00dejR" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000095_0957-4158(94)e0025-l-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000095_0957-4158(94)e0025-l-Figure1-1.png", + "caption": "Fig. 1. Fixed delivery rate cone winding.", + "texts": [ + " Often it is more convenient to draw the (low inertia) yarn supply over the end of the stationary package rather than rotate the (high inertia) package to unwind the yarn. This is commonly the case in knitting machines. In order that the yarn can be drawn off smoothly it is, therefore, wound into conical packages during its production. Whilst advantageous for the yarn user, this can present a problem for the yarn manufacturer. In those spinning processes which produce yarn at a constant rate (such as the modern open-end processes) winding the yarn onto a cone, as illustrated in Fig. 1, presents a cyclically varying mismatch between supply and take-up because of the difference in surface speed between large and small ends of the cone. Unless some kind of tension compensator is provided the resulting tension variations will cause a poorly wound package, yarn breakage, or both. At low production speeds the problem can be resolved by the use of a simple spring compensator, using a low-rate spring to maintain an adequately constant tension. Such a system is illustrated in Fig. 2; the two bollards around which the yarn passes are mounted on a disc which can oscillate back and forth to take in and let out yarn, thereby evening out the tension fluctuations as the yarn is wound from the small end to the large end of the cone" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001555_0301-679x(83)90057-9-FigureI-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001555_0301-679x(83)90057-9-FigureI-1.png", + "caption": "Fig I S;ooI piece o f the second set o![aiIed gear couplings", + "texts": [], + "surrounding_texts": [ + "Abnormal wear of gear couplings a case study\nT. Chander* and S. Biswas*\nA number of failures occurred in the gear coupling transmitting the torque from a steam turbine (rotating at 7500 r/rain) to a 1.5 MW generator through a 1 : 5 reduction gear box. Failure was found to be occurring in the form of rapid wear of teeth. Heavy wear steps were notices on both the hub and the spool pieces. Vibration measurements revealed that twice the running speed component was the most dominant frequency, with the maximum amplitude in the axial direction, indicating misalignment. Nonsynchronous components appearing in the frequency spectrum revealed that rubbing was taking place. Overall vibration analysis at different locations in the assembly indicated seating problems in the bearing pedestals with the bed plate.\nExcessive sludge accumulation was noticed in the gear coupling. Analysis of the sludge indicated the presence of large quantities of iron oxide, water, calcium and silica. Other elements of the gear coupling material were also present. However, no deviations in the physical and chemical properties were found through analysis of oil collected from the tank. Examination of the lubrication system showed lack of proper filters and centrifuge. The shape, size and orientation of the nozzle for the oil jet to gear coupling were found to be incorrect. Ferrographic analysis of the lubricating oil revealed a large quanti ty of wear particles. It was concluded that the wear was due to the misalignment aided by faults in the lubrication system\nKeywords: flexible couplings, gears, failure\nThe use of gear couplings in industrial steam turbine sets has been quite common in moderate to high speed applications, because they can acconnnodate a certain degree of misalignment while allowing freedom of movement in the axial direction. The common causes of gear coupling failures have been at tr ibuted to excessive shaft misalignment, incorrect lubricant flow, lubrication failure, excessive sludge formation, overloading and surge effects 1'2'3 .\nHeaw shaft vibrations have been found to indicate coupling wear due to misalignment 4 . Unsuitable lubrication systems, lubricant selection, nozzle size and shape have given rise to rapid wear rates in gear couplings a's . The problem of excessive sludge formation has also been reported 6 . Condition monitoring of gears using ferrography (a relatively new technique) has been carried out v.\nThe gear couplings reported in this case study had high failure rates due to rapid wear of the gear teeth. Although such failures are not common under normal circumstances in similar applications, a number of failures within a short time span gave rise to a critical situation. Excessive sludge accumulation was also occurring, and it was decided to investigate the problem from various aspects. Detailed vibration analysis, oil analysis and inspection of the lubrication system were carried out to find the causes of the unusual rate of wear. The methodology of the investigation and the results are reported in this paper.\n*Failure Analysis Group, Corporate Research and Development Division, Bharat Heavy Electricals Limited, Ilvderabad - 500 593, htdia\nCase history Several failures occurred in a gear coupling transmitting the torque from a steam turbine (rotating at 7500 r/min) to a generator through a reduction gear box of ratio 1 : 5. The turbine set was in a sugar plant which had a more or less continuous production schedule. Twenty-one days after installation, the unit was shut down as the production season was over. Production started once again during the next season and the first failure of the gear coupling occurred after 45 days of operation. Heavy wear of the gear teeth was found, with some fretting. A high wear rate of the replacement gear coupling was observed after just 13 days of operation and subsequently the lubricating oil was changed. Centrifuging of the oil was carried out after the unit had operated for ten days with the new oil. However, this did not reduce the rate of wear and sludge formation. The third set of gear coupling parts was put into service after a month of operation. The wear rate of this coupling was similar, but it ran for another two months until the pro-\nduction season was over. Excessive accumulation of sludge, which was dark brown and grit ty, was noticed right from the beginning. It was a thick greasy substance. Approximately 50 g of sludge were removed per week from the gear coupling (by scraping) when the third coupling was in operation.\nInspection of the failed gear couplings Photographs of the spool and hub pieces on the turbine side of the second set of failed gear coupling parts are\nTRIBOLOGY international 0301 679X/83/03014! 06 $03.00 \u00a9 1983 Butterworth & Co (Publishers) Ltd 141", + "Chander ~nd B/swas -- Abnormal weer of gear couplings\nshown in Figs ! and 2. The coupl ing had 52 teeth , -which were no~ crowned, wi th a p i tch circle d iameter o f 130 m m The wear o f teeth of the hub and spool pleces on the turbine side was iess than that on the gearbox side (Fig 3). Figs 4 and 5 show more details o f one t o o t h cu~ from_ the hub piece, Here, wear wi th some fretti~:g can be observed In atl the coupling pieces, the t ee th were covered wi th a lacquer-Eke substance. This layer, adhering to the t ee th could no t be removed by scraping wi th a fingernail.\nI n v e s t i ~ a t [ o r ~\nBased on the case his tory and_ observations of the failed gear coup;ing pieces, the fol lowing course af invest igation w~s under taken ~:o analyse the o reb!em in its ent i re ty:\ns Metallurg~cal analysis of th~ c\u00a2,:~pnng \"~o identif j n-aer:~ lurgica] ~eficiences if' any, m :no m a t e r i \u00a3 o f t;~e coup l ing <~ Vib raucn spec : rum s.nalysls to ioe~ti?v discrepant;co such as misal ignment , maba!ance etc ~ in the\nmechanical assembly\u00b0 e Analy s~s of the lubricating oil fc,,r c h e m i c a compos i t ion ,\nwear debris analysis e~c * Evaluat ion of the design of the ]u~r.,caung oz~ syRem\n~o gear coupling,\nMetallurgical analysis\nThe material o f the gear coupling Umported~ was an Nio~. steel containing 0.8% Cr, 1.38% Ni. 0.26% Me 0.56% Y n 0.04% Cu. The micros t ruc ture consisted of temoered marten site wi th about ] 5% undissolved ahoy carbides Hardness measurement p roduced values of the ozder o753 R~. which (from reported. ~oreratu~%~ appeared to be in ~- ' -\nv\u00b0ibratio~ s~ect~'um anN ys~s\nThe weat o f \u2022he teeth of the gear coupling :is taz'gely :ieza~ra e m u pen the relative sliding velocities of the mam~g :;eetn. The sliding v d o c i t y responsible for :api~ wear of tee ~ee~b depends u.~on:\nThe inaccuracms m the macmnmg of me ;ooth ~;of ' i i~ ThemisaEgnmen t betwee;:, the cav ing and the :Irh,er' elements ;n the ws tem, The orbital m o v e m e n t of d~e rotor due t\u00a2 eacess~{e bes:o mg clearazces and large unbaiancc e-_s\nh is apparenL therefore r.i:ar anion i-~?e mc~ors ~espc, ns\u00a3mc %r the high sliding velocities of the tced<~ couid be a3?rer_~ riately identif ied, no correTave action could be o)annec.\n~,~ is well ~nown ma~ abnormahues such as {mbamnc: mJs alignmen-! e t c can be de tec ted b? \u2022 detailed ~4bra~on spect rum anab,'s~s !n view of this. the authors carried ou: vibrat ion measurements at al[ the bearings and the ~ea c o u / n n g no~.s~ g, as shown m Fig 6 <\u00a2veralJ v~brahcs\nFig5 Close 'p v~.ew o f the Iop oftJ~e rooi~ show~_ . ~'*,. .n-\u00a2~.s ..~,\n142 _uns !983 Vo i 16 No 3", + "levels were measured at all the locations shown using a portable B & K vibration meter and a triaxial accelerometer. The meter had the provision to measure in the three different modes, acceleration in m/s 2 , velocity in mm/s and displacement in microns. All measurements were made in the peak-to-peak mode with 1 s time averaging. Table 1 gives the overall vibration levels measured in the different modes in the horizontal, vertical and axial directions respectively.\nIn addition to the measurements by the vibration meter, the vibration signals obtained through a B & K Triaxial accelerometer were recorded on a 4-channel B & K charge preamplifier. The recorded signals were subsequently analysed on a Nicolet Fast Fourier Transformer analyser to find out the discrete frequency component that was probably responsible for this problem. Signals were analysed both in real time and the time-averaged mode to see whether any unsteady processes were present in the vibration behaviour. Several salient features of the overall vibration measurements and vibration signatures obtained on various bearings merit consideration.\n(i) Overall vibration displacement levels in the horizontal, vertical and axial directions on all bearings were found to be of the order of 40/xm (peak to peak), which is more than the limit of 25/xm prescribed by the company based upon its experience. It is interesting to note that the acceleration levels associated with the vibration were found to be as high as 5 0 - 7 0 m/s 2 (peak to peak) on the turbine rear bearings, indicating the dominance of high frequency components.\n(ii) Bearing pedestals, when erected properly with a good seating contact area, should show from the top of the pedestal to the bot tom a descending trend in the overall vibration levels in all directions. In the present case study, the vertical vibration - a fairly good indication of relative movements - was found to be ahnost constant at all the locations on both front and rear turbine bearings. Also, the tightness of the bolts fastening the pedestals to the bed plate was checked and found to be correct. From this it was concluded that the seating of the pedestals on the bed plate was not uniform and adequate. The axial vibrations in the displacement mode on all bearings were of the same order of magnitude as the radial vibrations. This showed clearly that misalignment existed in the system.\n(in the low frequency domain) showed dominance of a 245 Hz component which corresponded to twice the running speed. Second, the axial vibration levels in both the displacement and acceleration modes were substantial.\nPick-up Acceleration Veloc i ty Displacement location m/s 2 mm/s /~m\nHorizontal Vert ical Axial Horizontal Vertical Axial Horizontal Vertical Axial\n1 9 - 1 0 8 - 9 1 0 - 1 2 2 6 - 7 5 - 6 8 - 1 0 3 5 - 6 4 - 4 . 5 6 - 7 4 8 - 9 2 .6 -2 .8 1 0 - 1 2 5 17 -20 3 4 - 4 0 5 0 - 7 0 6 5 0 - 5 5 18 -22 3 0 - 4 0 7 10 -12 1 2 - 1 5 6 . 5 - 7\n4 .5 -5 .5 2 - 2 . 2 4 . 5 - 5 3 6 - 4 0 3 3 - 3 4 3 2 - 3 4 3--3.8 2.2 4 - 4 . 5 40 3 4 - 3 6 3 4 - 3 6 3 - 3 . 6 1 .8 -2 4 - 4 . 5 3 8 - 4 0 3 4 - 3 6 3 4 - 3 6 2 . 8 - 3 1 .5-1 .7 3 - 3 . 6 40 3 6 - 4 0 3 4 - 3 6 7 - 8 5 . 5 - 6 15 -17 3 6 - 3 8 3 4 - 3 6 3 6 - 3 8 9 - 1 0 4 - 5 6 - 7 3 6 - 3 8 32 3 6 - 3 8 2 .8 -3 2 . 6 - 3 2 .2 -2 .4 3 6 - 3 8 3 6 - 3 8 3 6 - 3 8\nTR I BOL OGY international 143" + ] + }, + { + "image_filename": "designv11_65_0002443_6.2002-1377-Figure4-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002443_6.2002-1377-Figure4-1.png", + "caption": "Figure 4 shows a picture of the specimen and the coordinates used in the experimental results. The deformations at the crease (residual distortion) are measured by the optical video microscope camera to see the characteristics", + "texts": [], + "surrounding_texts": [ + "Many research works on membrane space structures including inflatable structures have been studied because of their attractive features as light weight and simple deployment mechanism. However, to apply a very large membrane structure in the future space applications, the improvement of the static and dynamic characteristics is requested. For example, the collision and unexpected deployment of the membrane element of inflatable antennas and reflectors in the course of deployment have to be avoided, and the surface accuracy of membrane is significant for constructing a high performance antenna reflector. Thus, the active control of membrane is one of * Associate Professor, Department of Built Environment, AIAA senior member, e-mail: furuya@enveng.titech.ac.jp \u2020 Graduate Student, Department of Built Environment, AIAA student member \u2021 Graduate Student, Department of Built Environment Copyright c\u00a92002 by Hiroshi FURUYA, Yoko MIYAZAKI and Yoshihisa AKUTSU. Published by the American Institute of Aeronautics and Astronautics, Inc., with permission. the solutions to improve the performance and reliability of the membrane space structures. Some researches on control of membrane space structures proposed the control of tensile force induced in the membrane through the actuation of the rim and catenary which support the membrane.[1] Piezoelectric film actuators (PVDF) have been focused in the field of smart materials and structures because of its light weight and actuator/sensor properties. Several research works on vibration suppression control were performed by piezoelectric films.[2, 3] Some researches proposed the shape correction of inflatable structures by controlling the tensile force of the membrane by attaching piezoelectric films, however, many of these research works assumed the membrane was completely flat or ideal without crease or wrinkle.[4] Thus, the active control of creased and wrinkled membrane in the course of deployment, or in the case of small tensile force at the first stage of deployment, should be examined in detail. The mechanics of creased membrane has been analytically investigated in our studies.[5, 6] They showed that the residual distortion of the membrane is localized around creases. Also, the analytical research on shape control with locally attached piezoelectric films was performed in our previous research work[7], and the possibility for controlling residual distortion by the locally attached piezoelectric films has been investigated. This paper is an extension of our studies, and the static shape control of creased membrane using locally attached piezoelectric films is experimentally studied to investigate the performance and possibility of the shape control. The effects of geometrical configuration of membrane, tensile force for the membrane, and driving voltage to the piezoelectric films, on the performance of static shape control are investigated through the experiments." + ] + }, + { + "image_filename": "designv11_65_0000166_9.256404-Figure2-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000166_9.256404-Figure2-1.png", + "caption": "Fig. 2. An example of the closed chain robot system from [16].", + "texts": [ + "), the solution to the resulting LP problem for finding max.u(t) and min. ~ ( t ) can be obtained at any time instant while searching for the optimal trajectory on the phase plane (xl - xz). A numerical example from [16] is used to verify the theoretical result on the optimal solution structure and illustrate the computation of the solution. In this example, the closed chain system consists of two 3-DOF planar robot arms moving a common object from the start position to the end position along a straight line path in Cartesian space, as shown in Fig. 2. The joint variables are (ei, +i, &IT, i = 1 for arm 1, and i = 2 for arm 2, respectively. The control torque variables are (Oi, ai, Yri)T, i = 1 for arm 1, and i = 2 for arm 2. The upper bounds on the control torques are (8.0,4.0, 2.0)', and the lower bounds on the control torques are (-8.0, -4.0, -2.OIT for both arms. The 1652 IEEE TFUNSACITONS ON AUTOMATIC CONTROL, VOL. 37, NO. 10, OCTOBER 1992 results of the minimum time trajectories in the phase plane and in Cartesian space (n, y , a) are given in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001428_s0167-8922(00)80155-8-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001428_s0167-8922(00)80155-8-Figure1-1.png", + "caption": "Figure 1. Roll ing-sliding tempera ture rig", + "texts": [ + "1 Temperature Rise Mapping The original technique for mapping temperature in EHD contacts was developed by Winer, who used an infrared (IR) microscope to measured thermal emission from the steel surface and from the lubricant in a sliding contact between steel ball and sapphire disc (30)(31). These emission measurements were then converted to steel surface temperatures using a combined calibration/analysis approach. The current study also uses a steel ball on sapphire flat contact, but unlike most previous work, the disc and ball are independently driven, to obtain any desired sliding/rolling combination, as indicated in figure 1. Some tests were also carried out in a pure sliding system as shown in figure 2. In both cases, the contact is contained in a stainless steel chamber whose temperature can be controlled to + 0.5\u00b0C and the ball is half-immersed in lubricant to ensure fully-flooded conditions. A custom-built infrared microscope is mounted on an XY table driven by micrometers connected to computer-controlled stepper motors. The microscope is focussed onto the steel ball surface within the contact, and uses an lnSb detector to sample the level of IR emission from the focus area" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000012_piae_proc_1922_017_029_02-Figure17-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000012_piae_proc_1922_017_029_02-Figure17-1.png", + "caption": "FIG. 17.-Front hub for 25/50 car. Alternative arrangements.", + "texts": [], + "surrounding_texts": [ + "374 THE INSTITUTION OF AUTOMOBILE ENGINEERS.\nThe point arises here as to how the hand of the gear should be selected for the bevel or crown-wheel so as to impose the heavy thrust on one or the other. The conclusion just reached regarding the double-purpose unit at the crown-wheel is on the assumption that the light thrust load oome's in this position, and this is in accordance with general practice. In order to make full use of the possibilities of the single-row bearing, however, the hand of the gear might be reversed, and certainly there will, in many cases, be more opportunity for obtaining a large capacity in the crownwheel position than behind the bevel-pinion\nPilot Bemirig. Although the pilot bearing, Fig. 20, does not appear on any of\nthe graphs, it will be remarked that it is in many casea an& example of a bearing applied to relieve a heavily loaded one without kaking into account the possibility of its own failure, which frequently occurs.\nClutch. The bearing selection for the clutch is not so ainenable to the same treatment, but it will usually be found that a light type single-row bearing w i t b u t fillibg-slot has su5cient capacity .to deal with a load which is only in occasiond operation.\nFigs. 17 to 20 show the aotual arrangements of the front-hub, rear-hub. gear-box and rear-axle bearings of the car referred to as No. 12.\nSUMMARY. The difficulties in the way of fixing definite factors of safety which will meet all cases are fully appreciated, but it is hoped that the preceding recommendatipns will form a useful basis for investigating existing designs. For convenienoe, the factors have been collected in Table I., the letters in Fig. 15 qorres o d ' with those used thmughout the drawings of the Sizaire-brm? chassis.", + "THE ENDURANCE ,OF BALL HEARINGS. 375", + "376 THE INSTITUTION OF A UTOMORILE ENGINEERS.\nFIG. lg.-Gear-bux for 25/50 h.p. car." + ] + }, + { + "image_filename": "designv11_65_0002879_gt2002-30005-Figure5-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002879_gt2002-30005-Figure5-1.png", + "caption": "Figure 5 - Shaker/Loader Concept", + "texts": [ + " The ZCAB\u2019s which were described previously [10] were located outboard of the AMB\u2019s. Actuation was provided via an electromagnetically triggered spring-loaded closure system. For the purposes of this initial testing, the ZCAB\u2019s were configured such that rotation acts to open the ZCAB against the spring preload. Table III presents the ZCAB parameters. loaded From: http://proceedings.asmedigitalcollection.asme.org/ on 02/02/2016 Te The loader/shaker, with the lamination profile shown schematically in Figure 5, is a modified heteropolar magnetic bearing configuration located at the approximate center of the rotor to allow the application of both steady and dynamic loads. The lower poles provide a static load, while the upper poles sets at \u00b145 degrees provide dynamic excitation capability. The unused poles were removed to reduce cross-loading effects. The loader/shaker was operated in an open-loop configuration for the initial testing. Excitation is provided by PWM amplifiers similar to the radial bearings" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002246_iros.2000.893240-Figure7-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002246_iros.2000.893240-Figure7-1.png", + "caption": "Figure 7: Error bounds for coupled system with w n = 12.24. The error bounds in frequency and amplitude are 52 = {w = [10.5,13], D = [0.85,1.12]}, being the estimated values ~ = 12.2 and D = 0.94. The real values are w = 12.9 and D = 1.", + "texts": [], + "surrounding_texts": [ + "ure 5-a, the measured values for the frequency of oscillation are w = 9.88 and for the oscillator input amplitude D = 0.85, and the predicted ones w = 9.9 and D = 0.83, with 52 = {w E [9.6,10.2], D E [0.75,1]}. Note that the discretization of D is very small (scale is 0.25), and thus a higher resolution for D would improve the error bounds, making other curves crossing the unity, giving a more precise boundary. Figure 5- b shows the error bounds for the system introduced in example 2, which are R = {w E [11.4,13],D E [0.75,1.0]}.\nAnother way of checking error bounds is through the use of error discs, [3]. For which frequency, draw a disc of radius m, and draw circles on the -1/G locus as in Figures 6 and 7. The points where the tangent of the circles intersect the N ( D , w) locus at the same frequency give the range for w and D.\nExample 5 The graphical method for error bounds based in the use of error discs is shown in Figures 6 and 7. This method allows to visualize the variation of the uncertainty with frequency.\nThe degree given by (3) can also be determined by inspection. Considering the -1/G loci as a spring, and put pegs at the points correspondent to w,in, wmaz, Dmin and D,,,. Now pull the string from the pegs. The degree is nonzero if the strings still cross, otherwise the degree is zero and the analysis has failed, [3].\n4 Extension to Multivariable Systems\nElements of the transfer matrices for MIMO systems are complex transfer functions. Singular value decomposition (SVD) [8] provides the tools for a notion of the size of G(jw) versus frequency, equivalent to a MIMO gain. Direction information can also be extracted from the SVD. Indeed, for inputs along\nVi (columns of left singular vector of G), the outputs along V, will be amplified by a gain ui. Thus, Umaz of G(jw) define maximum amplification of a unit sinusoidal input a t frequency w, and the corresponding singular vector the direction for maximum amplification. One useful norm for such stable systems is the H , norm, defined by\nFor the derivation of the inequalities, consider X = X1 + X * as the input vector for the oscillators, X = -Gn(X), and that X * = P*X contains only harmonics whose indexes are in K*. Thus,\nX' = -G( jw)P 'n(X1+ X * ) (13) X I = - G ( j w ) P l n ( X i + X * ) (14)\nwhere n is a column vector whose elements are the dynamics of each neural oscillator. Writing (13) as X * = F ( w , D l , X * ) , it will be necessary to find an inequality that guarantees IIX*112 5 E + JIFll2 5 E, in such a way that F maps B(0, e) to itself, using the L2 norm for vectors. Thus,\n1 F(w, D1,X' ) = -Zmag ( kg* G(jkw)ckeJ(\"'+v\")\nfor n ( D l s i n ( w t ) + X ' ) = [mag (EksK ckej(kwt++'k)). By application of Schwartz's inequality,\nk \u20ac K *\n-1550-", + "This corresponds to a worst case analysis, and therefore it is independent of Di and 9,. The computation of p(w) requires only a finite number of terms if G is a stable strictly proper transfer matrix. From (14) and (16) results the phasor equation (18),\nwhere D1 = [d,l dX2e392 . . . d x n e j 9 n l T , and N is a diagonal matrix which elements are the DFs of each neural oscillator. For an invertible matrix G results\n[G-'(jw) + N ( D i ) - E ( D i , X * ) ] Di = 0. (18)\nTo estimate a bound, lets apply the Schwartz Inequality\nIlP*n(x)ll; 5 IIP*(n(X) - n(xl))ll;+ II P'n(X1) 11;\nThis last relation results from the Pythagorean equality on the Hilbert space of square-integrable periodic functions. From (15), one obtains the inequality (19):\nl l F l l 2 I P ( w ) ( d D l , E ) +P(Dl) ) I E (19)\nThe function that measures the error at the output of n is given by\nFinally, the resulting condition, using (18), is\n0 = ll[G-l + N(Di)]Dillz/~(Di,~) I1 (20)\nwhere\u20ac isdeterminedusingq(D1,E) = IJE(D~,X*)DIIJ~ and equation (19). A closed-bounded set 0, which contains the estimated frequency, system output amplitudes and phase-shifts - ($, Cy=, d;, Cy=2 @i), is found by all points in the neighborhood that satisfy the inequality (20). A better bound could be achieved through pole shifting, similarly to the SISO case.\nOf course, other norms could be used for p(w) in equation (8 ) , such as the H , norm, given by (12). This norm would be conservative for some cases, but not when equal to umaz(G(jkw), for any k. This would be equivalent to J I F ~ ~ ~ / E = y 5 1, where y is the L2 gain of AL, with A = P * n ( X ) and L = CkEK' G(jkw) (which could be represented as an interconnection diagram for the error dynamics, [8]). In addition, the L , norm could be used instead of the L2 norm to find the error bounds.\nThe singular value decomposition may also be very useful in other situations, such as the analysis of neural oscillators inter-connected in networks. Indeed, when a neural oscillator has more than one input (from other oscillators or from other plant state-variables), it is often difficult to infer which connections should be made to get a desired performance. Thus, inputs which drive the system along the direction of a minimum singular will have a small effect in driving the plant, and may even be negligible for very small umin. On the other hand, the direction associated to umas is useful to not only infer maximum propagated errors, but also to determine the optimal inter-connection of networks, so that the network outputs will drive the system with the maximum amplitude, along the maximum singular direction.\nFor the results thereafter presented the neural oscillators are connected to a MIMO system with transfer matrix G, without mutual connections among the oscillators, i.e., the oscillators are coupled through the dynamics of the MIMO system ( N , the approximated transfer matrix for the neural oscillators is diagonal, containing each element the approximated dynamics of each neural oscillator). If there were mutual connections, the computation of the error bounds p and q would be harder. In addition, it is assumed that all neural oscillators oscillate at only one resonance mode, [2] (indeed, a large spectrum of frequen-\n- 1551 -", + "cies would imply the use of a Describing Function Matrix, [5]).\nExample 6 For the system in Figure 4 of example 3, Figure 8 shows the graphs obtained from (20) varying D1, D2, cp and W . As shown, there are two resonance modes at w = 8.6 and at w = 10.9 with a well defined error interval of frequencies at [8.3,9] and [10.2,11.2]. Figure 8-b shows the plot for the system in Figure 4-b of example 3. The system does not oscillate at the second resonance mode, and therefore the error intervals are only determined for the first mode, i.e., w E [7,8], with 2ir = 7.5.\nThe graphic method of error discs (example 5) might also be extended to two oscillators connected to a MIMO system. However, instead of error discs, the method would now consist of error spheres, and, for higher dimensions, hyperspheres.\n5 Discussion and Conclusions\nA method was presented to infer limit cycle stability and error bounds on the estimates introduced by using describing functions. For neural oscillators connected to a low-pass system, the higher harmonics of the signal are removed and the estimate is accurate, and therefore the intervals of uncertainty are small. Stability of multiple oscillators connected to a MIMO system was also analyzed.\nThe work here presented is part of the humanoid robot project Cog, [4], shown in Figure 9-b. This analysis provides theoretical support to existing algorithms for the control of the robot arms using several biologically inspired Matsuoka neural oscillators (exemplified in Figure 9-a).\nThe error bounds determined for the estimated parameters correspond to the expectations. Further-\nmore, the algebraic equations (which were applied for parameter tuning) are solved rapidly, so that building a graph to check stability is simple and fast.\nA method was presented to estimate the multivariable error, which takes into account the maximum amplification of the error by the plant. Using this technique, error bounds are estimated for MIMO systems connected to multiple oscillators. SVD is also useful to check for the numberltype of connections necessary in a network of multiple neural oscillators to achieve a desired performance.\nAcknowledgments The author was supported by a Portuguese government grant PRAXIS XXI BD/15851/98. Support for this research is provided by DARPA/ITO, contract number DABT 63-99-1-0012 for Natural Tasking of Robots Based o n Hum a n Interaction Cues.\nReferences A. M. Arsenio, \u201dTuning of Neural Oscillators for the Design of Rhythmic Motions\u201d, accepted to IEEE Conf. on Robotics and Automation, 2000. A. M. Arsenio, \u201dNeural Oscillator Networks for Rhythmic Control of Animats\u201d, submitted to From Animals to Animats 2000. A. R. Bergen, L. 0. Chua, A. I. Mees and E. W. Szeto, \u201dError Bounds for General Describing Function Problems\u201d, IEEE Trans on Circuits and systems, 6 (29), June 82. R. A. Brooks, C. Breazeal, R. Irie, C. Kemp, M. Marjanovic, B. Scassellat and M. Williamson, \u201d Alternate Essences of Intelligence\u201d, AAAI-98 A. Gelb and W. Vander velde, \u201dMultiple-Input Describing Functions and Nonlinear System Design\u201d , McGraw-Hill, 68. K. Matsuoka, \u201dSustained Oscillations Generated by Mutually Inhibiting Neurons with Adaptation\u201d, Biological Cybernetics, 52:367-376, 85. J-J. E. Slotine and W. Li, \u201dApplied Nonlinear Control\u201d Englewood Cliffs Prentice-Hall, 91. K. Zhou and J. C. Doyle, \u201dEssentials of Robust Control\u201d.\n- 1552 -" + ] + }, + { + "image_filename": "designv11_65_0001226_20.717826-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001226_20.717826-Figure1-1.png", + "caption": "Fig. 1. Permanent magnet type stepping motor.", + "texts": [ + " This paper analyzes static characteristics of the permanent magnet type stepping motor with claw poles using three-dimensional finite element method, and shows the good agreement between calculated static torque and measured one. First, we calculate the magnetic flux distribution in the motor and clarify the flux distribution, especially in the air-gap, with and without exciting current. Then, we calculate cogging torque and stiffness cliaracteristics, and show the good agreement between calculated values and measured ones. Moreover, we show the difference of torque Characteristics developed by two types of magnetic situation of the permanent magnet. 11. NUMERICAL MODELING Fig. 1 shows the schematic diagram of a plastic magnet type stepping motor with claw poles. It consists of Manuscript received November 3, 1997. T. Ishikawa, 81-277-30-1742, fax 81-277-30-1707, e-mail tisikawa@cc.gunma-u.ac.jp; M. Matsuda; M. Matsunami, e-mail matsunamQtech.gunma-u.ac.jp a rotor and two stator sections. The rotor consists of a multi-polarized permanent magnet and a shaft. Each stator has a bifiler-wound coil and yokes with claw poles. Main parameters of the experimental motor are as follows; number of phases is 4, number of pairs of poles is 12, step angle is 7", + " In the case that Bm=Oo, the flux density produced by the excitation of phase A neglects that produced by the rotor magnet. This situation corresponds to the unstable position of the motor. The magnetic flux density in the air gap is smaller than that in the teeth. Then, it is not drawn in Figs. 4 and 5 . Investigation of magnetic flux density in the air gap is very important, because it mainly decides the torque and magnetic energy of the motor. Fig. 6 shows distributions of flux densities in the air gap with and without exciting current at several cross sections shown in Fig. 1. These figures also show the distribution of flux density calculated by two-dimensional finite element method to clarify the difference between two- and three-dimensional analysis. Excitation of coil A produces a slight change in flux density at lower section and produces no change in flux density at upper section. Describing in detail, excitation of coil A produces +B, at the region where 7.5\" 5 8 < 22.5\", -B, at the regions where 0\" 5 8 < 7.5\" and 22.5\" 5 0 < 30\", +Be at the region where 0\" 5 0 < 15\", -Be at the region where 15\" 5 8 < 30\", and +B, at the region where 0\" 5 8 < 30\", respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001332_rsta.1998.0268-Figure2-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001332_rsta.1998.0268-Figure2-1.png", + "caption": "Figure 2. Two equilibria: (a) stable and (b) unstable.", + "texts": [ + " The main importance of the stable and unstable manifolds is two-fold as follows. 1. They may define a basin of attraction (as explained later). 2. They may connect at a point other than the solution: such a connection is called either homoclinic if a manifold connects with itself, or heteroclinic if a manifold connects with another manifold. Such connections are responsible for global bifurcations. (f ) Saddle and repellors A solution may not be stable, as may be pictured by a ball on a surface (see figure 2). The cases in figure 2 are straightforward: both solutions are equilibria but figure 2a is stable and 2b is unstable. In two dimensions a solution may have the property of figure 3. In figure 3, the ball is \u2018stable\u2019 in one direction but \u2018unstable\u2019 in another. Clearly, however, the system is not \u2018robust\u2019: a small perturbation will move the ball away from the equilibrium solution and so the system is said to be unstable. For such an equilibrium point in two dimensions this would correspond to a system with one eigenvalue with real part positive and the other with real part negative. The eigenvectors give information about the direction of trajectories in the vicinity of the equilibrium point in phase/state space" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001783_00423118808968908-Figure2-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001783_00423118808968908-Figure2-1.png", + "caption": "Fig. 2. Components of freight truck model on curved track (rear view).", + "texts": [], + "surrounding_texts": [ + "The concept of Damper-Coupled Wheelset as a compromise between independently rotating wheels and conventional rigid axled wheelset has significant influence on the dynamics of railway vehicle. Pioneering and preliminary study by Bennington [I] on a single wheelset and primary suspension unit as a closed loop guidance control system showed potential of such concept for railway vehicle application. Recent work on the stability analysis of a single DCW [2], and a freight truck model with two DCW and a pseudo-car body [3] have demonstrated the significant influence of a wheelset coupler parameter on the stability of a wheelset as well as the primary and secondary hunting behaviour of a freight car system. The stability study of the truck model by Ahmed and Sankar [3] has shown that a speed dependent optimal torsional coupler damping between the wheels within an axle can provide over 100% improvement of the vehicle critical speed in comparison to the conventional system. A parametric study [4] of the freight truck model has shown superior stability performance of the system with optimal wheelset coupler damping in comparison to that of a conventional system. Performance of a freight car system with a concept such as DCW cannot be evaluated through tangent track stability analysis alone. It is a well known fact * Research Assistant Professor ** Professor and Director Concordia Computer Aided Vehicle Engineering, (CONCAVE) Research Centre, Department of Mechanical Engineering, Concordia University, Montreal Canada D ow nl oa de d by [ Q ue en sl an d U ni ve rs ity o f T ec hn ol og y] a t 1 5: 54 1 3 O ct ob er 2 01 4 296 A. K. W. AHMED A N D S. SANKAR that conventional freight cars have conflicting parameter requirements between stability and curving performance, and curved track in reality cannot be avoided. Therefore, to examine the potential of the concept of DCW for railway vehicle, it is necessary to evaluate the influence of the wheelset coupler damping parameter on the curving performance. The steady-state approach in curving analysis developed by Newland [5] and Boocock [6] is very efficient in providing vast information regarding the effects of model parameters, track curvature, cant deficiency (lateral unbalance), etc. on the curving performance. In the recent years, nonlinear and more realistic models for curving analysis have been developed. However, in this paper simplified approach with the assumption of linearity is adopted to aid in extensive computation for a comparative study of two systems, namely conventional and DCW. Linear models, in addition to ease of computation can provide useful and vast information pertaining to the system behaviour through accurate interpretation of results. The primary objective of this paper is to investigate the influence of the wheelset coupler damping parameter on the steady-state curving behaviour of a railway freight car system. An eleven degrees-of-freedom (DOF) freight truck model with two DCW and a pseudo-car body is utilized in this investigation. The model is validated by comparing results in its limiting case of a rigid coupler, with the results of an identical model with conventional rigid axled wheelsets. The curving performance is evaluated in terms of response to track curvature and track super-elevation, as well as slip and flanging boundaries. In all cases, the results are compared with those corresponding to rigid axle simulation of a conventional system. 2. MODEL DESCRIPTION The railway freight car model used in this investigation is similar to that used in [3] and [4] for tangent track stability analysis. For the steady-state curving analysis, the track is considered to be rigid and of constant center-line radius (R), with uniform angle of super-elevation. An eleven DOF, three-piece freight truck model on curved track is shown in Figs. 1 and 2, which includes two side frames, a bolster, two DCW, and a pseudo-car body. The various DOF include: lateral, yaw and relative spin between left and right wheels of each wheelset; truck frame lateral, yaw and warp motions; pseudo-car body lateral and roll motions. In curved track modeling various considerations include, addition of precession velocity in the expressions for creepage, and centrifugal force that acts on each component of the model. Further, due to super-elevation, depending on centrifugal force there is a lateral unbalance force. The lateral unbalance in turn results in a wheel load shift which affects the gravitational stiffness force and creep force. Detailed derivations for each of these considera- D ow nl oa de d by [ Q ue en sl an d U ni ve rs ity o f T ec hn ol og y] a t 1 5: 54 1 3 O ct ob er 2 01 4 CURVING OF RAILWAY FREIGHT TRUCK Fig. I . Components of freight truck model on curved track (plan view). tions are presented in [7], and are briefly discussed as follows: 2.1. Cant Deficiency Force In curving analysis of railway vehicles, the lateral force unbalance is usually expressed in terms of cant deficieny (cp,). The cant deficiency is defined as the angle between the resultant of centrifugal force (mV2/R) and the wieght (mg), and the normal into the rail plane. Making small angle approximation, the expressions for cant deficiency force on wheelset, truck and car body are: g, = miii!rn. (1) D ow nl oa de d by [ Q ue en sl an d U ni ve rs ity o f T ec hn ol og y] a t 1 5: 54 1 3 O ct ob er 2 01 4 A. K. W. AHMED AND S. SANKAR 2.2. Wheel Load Shift In the presence of cant deficiency, there is an unbalance force in the plane of the track, which introduces transfer of load between left and right wheels. From equation (I), an unbalance cant lateral force WApp (qd) acts through the CG of the vehicle at a height hc- from the rail plane. By taking moments about left and right wheel contact point, the resulting expressions for normal load at the left and right wheels are: WAPP *L.. 7 (1 + h@*) (2) Where h is substituted as an approximate value of h,--/a. 2.3 Gravitational Stiffness Force Gravitational stiffness force (GSF) is the lateral component of the normal load that acts along the rail plane due to the contact angle between wheel and rail and angle of wheelset roll. Summing the forces in the lateral direction and substituting for normal loads from equation (2) leads to the expression for GSF as: GSF = - W D ow nl oa de d by [ Q ue en sl an d U ni ve rs ity o f T ec hn ol og y] a t 1 5: 54 1 3 O ct ob er 2 01 4 CURVING OF RAILWAY FREIGHT TRUCK 299 2.4 Creep Forces and Moments The creep forces are defined as the product of creepage and creep coefficient, where the creep coefficients are a function of normal load. For a curving model under the influence of cant deficiency, the normal load is not the same for left and right wheels. The expressions for creep coefficients are derived using load proportionalities acoording to Kalker's theory, and expressions for normal loads on the left and right wheels as given by equation (2). The creep forces and moments are first obtained at the wheel/rail contact plane, which then are transofrmed into fixed body reference system of the truck frame. For steady-state curving analysis, the derivative terms in the expressions are dropped, except for terms which represent wheel spin velocity. In doing so, the expressions for creep forces and moments on the left and right wheels of leading and trailing wheelset are obtained as: r a F L = [ f , , [ l + i h $ d ] [ f +-v L b ] ] ? L + 0 (4) where F/R is + e/R for leading wheelset, and - e / ~ for trailing wheelset. The term e*/R appears due to the chosen coordinate system, where qr = - b/R indicates radial alignment of the wheelsets with the track. 2.5 Suspension System and Wheelset Coupler Other forces and moments acting on the freight truck model include the force and moment due to primary and secondary suspensions, and moment due to the wheelset coupler. In deriving these expressions it is assumed that both wheels within an axle move together in all directions except in the spin mode. The primary suspension system modeled as parallel stiffness and damper, is D ow nl oa de d by [ Q ue en sl an d U ni ve rs ity o f T ec hn ol og y] a t 1 5: 54 1 3 O ct ob er 2 01 4 300 A. K. W. AHMED AND S. SANKAR referred to the connection between wheelsets and truck frames. Each wheelset can move in the lateral and yaw directions with respect to the truck frame resisted by the primary suspension. The wheels within an axle also have a relative spin D O F resisted by coupler damping (DAx). The secondary suspension system is referred to the connection between truck and car body. Various elements of the secondary suspension include stiffness and damping that resist lateral, longitudinal and warp motions of the truck frames with respect to the bolster-car body assembly, and centre plate connection that resists the truck-bolster yaw motion with respect to the car body. All elements in this investigation are assumed linear." + ] + }, + { + "image_filename": "designv11_65_0002961_05698198108983034-Figure3-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002961_05698198108983034-Figure3-1.png", + "caption": "Figures 3 and 4 show how the moving-wave concept was implemented in the seal test apparatus. The drive ring is designed so that the primary ring can float and align itself but cannot rotate about the shaft axis. The primary ring thus takes its alignment from the face of the rotating secondary ring. The waviness-drive cylinder is driven by a worm at low speed (2 revolutions per 24 hours). The waviness-drive cylinder seals into the housing with two O-rings", + "texts": [], + "surrounding_texts": [ + "D ow\nnl oa\nde d\nby [\nU ni\nve rs\nity o\nf G\nla sg\now ]\nat 1\n6: 35\n1 8\nD ec\nem be\nr 20\n14", + "A Test Apparatus for Measuring the Effects of Waviness in Mechanical Face Seals 373\nacts as the support for the vessel. The pressure vessel is designed to be readily taken apart by removal of bolts. The secondary-ring support rotor is removed by removing the capscrew. T o minimize axial runout at the seal face, the end of the shaft is ground true with the bearings. The rotor is ground both where it mates with the shaft and the secondary-seal ring. The two secondary-seal ring faces are ground parallel.\nThe primary ring is made of Pure Carbon P658RC. The unusual proportions of the ring result from adjusting the geometry so that sealed pressure causes no rotation about a circumferential axis. The primary ring is driven by lugs which engage two notches. Two, as opposed to a greater number of notches, were used because this arrangement gives the primary ring maximum freedom to float radially to align itself. Second harmonic waviness is minimized by driving the ring through the centroid of its cross section. Any remaining second harmonic waviness produced by drive forces will be more easily flattened than the third harmonic waviness built into the ring, so is not considered important. The secondary seal is located at the left end on the inside diameter. The balance ratio for the design shown is 1.0. Springs to the left of the seal provide a preload of about 0.2 MPa (30 psi).\nThe secondary-seal ring is fabricated from tungsten carbide. This ring is also of a zero-pressure moment design. An unusual feature of the particular design is that the mechanical force at the right-hand side of the seal has been reduced to nearly zero by placing the O-ring seal on the right-hand face as opposed to the outside diameter. It is useful to minimize the axial force to reduce coning effects - caused by friction at the load bearing point in conjunction with radial deformation due to pressure and temperature. The secondary-seal ring is held in place by spring clips (not shown) which engage notches in the ring.\ndesigned so that gas pressure can be supplied through the end plate and into the waviness cylinder. Each of the six holes on the waviness-drive cylinder is connected to this gas pressure. When the primary ring is positioned over these six holes and sealed off by six O-rings, then the primary ring becomes subjected to six small regions of gas pressure on its inside. These forces produce moments about the centroid of the ring. The staggered location causes alternating moments of opposite sign. Face waviness and an alternating-face taper are produced by these moments. As the waviness cylinder is turned by the worm and gear, the waviness pattern moves relative to the seal primary ring.\nThe torque sensing element is shown in Figs. 2 and 3. The axial load that must be carried by this member is large, 53 000 N @ 6.9 MPa (12 000 Ibs @ 1000 psi) whereas the torque to be measured is quite small (down to a few N.m (in-lb)). The axial load is carried by ten 3 x 25 x 102-mm (0.120 in x 1 in x 4 in) long elements. These elements have a very high axial stiffness and strength but low torsional stiffness. The torsional load is sensed by two beams which have a low axial stiffness. The axial and torsional loads are in effect isolated so that torque can be measured independent of axial load. Strain gauges are mounted on the two beams. Test results verify the designed behavior. There is no measurable influence of pressure on torque at least up to 300 psi where the test was performed.\nThe torque sensing element is sensitive to temperature. A temperature increase of 50\u00b0F causes a zero shift of several in.-lb. This effect is repeatable and has been corrected in the data reduction.\nDATA ACQUISITION\nThe test apparatus is designed to operate continuously and unattended most of the time. Data collection, data storage, graphic display, machine control, and safety monitoring are performed primarily by a Hewlett Packard 9835 computer, 6940 multiprogrammer and 7225 plotter.\nTorque and seal primary-ring temperatures are measured several times a minute and plotted and stored on magnetic tape approximately one time per minute. The torque measurement has been described previously. Seal temperature is measured by a thermocouple located 0. I in behind the center of the carbon seal face. Leakage is measured by a counting device which dumps and actuates a relay each time 2.5 rn of leakage has accumulated. The relay, in turn, interrupts the computer which then determines the time of the interrupt and calculates the time rate. This device has the capability of measuring rates ranging from a fraction of a mllmin to 50 mflmin. Leakage is recorded and plotted whenever new data become available or one time a minute at higher leakage rates.\nSystem pressure is checked once pel- second by the computer. When pressure falls below the set level, an air-driven piston pump is activated for one stroke. This adds about 2 m P of fluid to the system. Using this type of control in conjunction with a gas-charged accumulator results in system-pressure control to about 5 psi with no significant\nFig. 4--Exploded view surges.\nD ow\nnl oa\nde d\nby [\nU ni\nve rs\nity o\nf G\nla sg\now ]\nat 1\n6: 35\n1 8\nD ec\nem be\nr 20\n14", + "Wear measurements are made using a surface analyzer and gauge pins. Only the 0.125-in nose length of the carbon is included in the wear measurement in order to minimize the effects of dimensional changes of the carbon due to water absorption.\nMeasurements of waviness under zero pressure conditions have been made using an inductance type proximity probe.\nWAVINESS\nThe waviness introduced into the carbon ring by the gas pressure can be computed using ring deflection theory such as presented in Ref. (9). It can readily be shown using Fourier analysis that 6 alternately spaced forces, as shown, produce a waviness with a predominate n = 3 (third harmonic) characteristic. Waves of n = 9, 15, etc, are also present but their magnitudes are insignificant. Approximating the pressure distribution by a uniform distribution having the same total force as the circular-pressure distribution, a Fourier analysis of the moment produced by the gas pressure gives\nwhere for this particular seal\nn = 3-number of waves p, - gas pressure\nd = 12.7 mm (0.5 in)diameter of the gas pocket e = 4.85 mm (0.191 in)axial distance between gas pockets and centroid of cross section r, = 48.3 mm (1.90 in)-\nseal inside radius 1; = 52.8 mm (2.08 in)-\nseal centroidal radius 711, - distributed moment in N.m/m of\ncentroidal circumference.\nUsing ring deflection theory, for a cosine moment, the waviness (axial displacenlent of face) is given by\nThe alternating face tilt or rotation is given by\nEJ, (n2 - 1)2\nE = 2.07 x loi0 Pa\n(3 X lo6 psi) - Young's Modulus\nG = 0.86 x 10'0 Pa\n(1.3 x 106 psi) - Shear Modulus\nJ , = 18 350 mm4 (0.0441 in4) [51\n- Moment of inertia about a radial axis\nJ , = 593 1 mm4 (0.0143 in4)\n- Torsional Constant\nAt p, = 6.9 MPa (1000 psi), Eqs. [ I . ] through [5] give\nv = - 7.64 p,m (301 @in) cos no\n6 = 1165 x 1 0 . 6 ~ 0 ~ no\nThus, the face waviness is combined with an alternating radial taper; that is, when v is negative (the gap is large), the faces are convergent and vice versa. In the converging portions, a considerable hydrostatic-fluid pressure component is developed which acts to support most of the seal load. As the wave moves around the seal, the alternating radial taper is first modified by wear and then reaches an equilibrium shape.\nTEST CONDITIONS\nFor the tests reported herein, the operating conditions were as follows:\nSeal environment temperature 38\u00b0C (100\u00b0F) Pressure - 3.45 MPa (500 psi) Speed - 1800 rpm Water - tap water\n60, 30 20 TEST NO. 40\n15 -;. .- E . 5 - lo u c3 u Y\n9 5 -1\n3.45 MPa (500 psi) WATER PRESS RPM I\nwhere TIME ( h ) Fig. C T e s t results--Flat face\nD ow\nnl oa\nde d\nby [\nU ni\nve rs\nity o\nf G\nla sg\now ]\nat 1\n6: 35\n1 8\nD ec\nem be\nr 20\n14" + ] + }, + { + "image_filename": "designv11_65_0003190_cca.2000.897544-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003190_cca.2000.897544-Figure1-1.png", + "caption": "Figure 1: Planar Vertical Takeoff and Lauding Aircraft.", + "texts": [ + " An iterative off-line computational algorithm is then proposed to approximately generate the flat output reference trajectories in terms of the desired non-minimum phase outputs trajectories. Section 3 presents the simulation results. Section 4 is devoted to present some conclusions and suggestions for further research. 2 lkajectory macking for the PVTOL Aircraft Example 2.1 The PVTOL aircraft model The simplified description of the dynamics of a planar vertical take-off and landing (PVTOL) aircraft is given by the following magnitude and time normalized model (see Figure 1 ) ii = -~ ls ine+Eu2cose e = u2 (2.1) 2 = u 1 c o ~ B + a ~ 2 s i n 8 - g where z and z are the horizontal and vertical coordinates of the center of gravity of the aircraft, respectively measured along an orthonormal set of fixed horizontal and vertical coordinates. The angle 0 is the aircraft\u2019s longitudinal axis angular rotation as measured with respect to the fixed horizontal coordinate axis. The controls u1 and u2 represent normalized quantities related to the vertical thrust and the angular rolliig torque applied around the longitudinal axis of the aircraft respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002998_esej:20020106-Figure3-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002998_esej:20020106-Figure3-1.png", + "caption": "Fig. 3 Second generation 1998", + "texts": [ + " In a competitive demonstration situation this becomes very boring for both the spectators and the assessors. It was at this stage that the TCS Programme started. Early work involved a market survey, which identified that the mechatronics workbench was a viable idea and indicated the typical kit cost that the market could sustain. It also identified that the largest market areas would be those that would require a simpler concept than the layered architecture shown in Fig. 1. The prototype developed (Fig. 3) incorporated a number of chassis panels (with a matrix of punched holes to permit design variations), both DC and stepper motors, for propulsion and steering, and a selection of wheels. The PIC 17C44 board was still used for control. The water pump system was discontinued and a 5 kg weight was utilised to compare vehicle performance unloaded and loaded. This version showed a substantial improvement, especially in audience engagement during the competitive assessments. There were still problems concerning speed (0-3 d s ) , and the cost of the motors and wheels, which were sourced fbm industrial components" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001311_6.1997-2631-Figure8-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001311_6.1997-2631-Figure8-1.png", + "caption": "Figure 8 Flow field at the bearing hole location for 0.016\" axial clearance", + "texts": [ + " A force balance indicates that the seal is in equilibrium at the clearance and pressure conditions being analyzed. The pressure differential across the seal and the seal/rotor air gap were selected based on measured values obtained during full scale testing. For the test conditions analyzed (7.1 psid and an air gap of 0.016\"), the calculated flow rate of 0.97 Ibm/s agrees very well with the measured value of 1.03 Ibm/s. The seal was analyzed at the measured test clearance of 0.016\", and the forces were shown to be balanced to within 2%. Figure 8 shows the flow field in the plane of the orifice hole, and Figure 9 shows the flow field in the plane of the vent slot. In both locations, there is mixing of the flows from the dam and air bearing regions of the seal, and the strong radial component originating at the dam appears to prevent the formation of a hydrostatic film at the air bearing. It appears that the radial flow from the dam slows upon entering the hydrostatic bearing region, resulting in increased pressure at the bearing face. This increased pressure would inhibit the seal's ability to close" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002100_pime_proc_1986_200_139_02-Figure3-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002100_pime_proc_1986_200_139_02-Figure3-1.png", + "caption": "Fig. 3 Geometry of the contact ellipse", + "texts": [ + "comDownloaded from A1 Original inner race centre A2 Deflected inner race centre B Fixed centre of outer race C Deflected ball centre or (1 3) wi(d, - D cos ai) + w0(d, + D cos a,) w, = (dmD/2)(cos a. - cos ai) The epicyclic ratio is always defined with respect to a stationary outer race; thus, (14) w -6.J A=-- - IL+ - 0, (d , - D cos a,) (d , D / ~ ) ( c o s a, - cos mi) From Jones (1) the rolling axis p is given by D sin a, 2 0 cos a, + d, p = tan-\u2019 the ball speed is ( d , + D cos a0)(w, - 0,) wb = D COS(CL, - p) and the spin speed is u s p i n = 6.Jb cos(ai - p) 2.3 Spin power Consider the geometry shown in Fig. 3. The ellipse is divided into m elements on the major axis and n on the minor. The element i, j has the dimensions 6 y x hx, where a 6x = - m (( 1 - x2/a2)b2) 6y = n However, (20) a m x = ( i - 4) 6x = ( i - 4) - therefore, (21) ([I - { ( i - $)/m)2]b2)\u201d2 n 6 y = Proc Instn Mech Engrs Vol 200 No C5 @ IMechE 1986 at Purdue University on March 13, 2015pic.sagepub.comDownloaded from The coordinates of the centroid of the element i,j are ((i - 4) sx, 0\u2019 - +) 6y) thus the radius r is given by ri j = [ { ( i - +) Sx}\u2019 + {(j - 4) 6y}]\u2019/\u2019 and the angle 4 by + i j = tan-\u2019{( 0: - $1 6Y } i - $) Sx For the coordinate system shown in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002565_0022-0728(89)87280-8-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002565_0022-0728(89)87280-8-Figure1-1.png", + "caption": "Fig. 1. Schematic representation of the electrochemical ceil. (a) Plexiglass cover (6 \u00d7 5 \u00d7 1 cm); (b) rubber gasket ( 6 \u00d7 5 \u00d7 0 . 1 cm); (c) stainless steel base with inlet and outlet for water circulatio~ t 6 \u00d7 5 \u00d7 1 cm); (d) channel and inlet for the calomel reference electrode; (e) platinum counter electrode (~ \u00d7 0.2 cm); (f) cell chamber (0.1 c m \u00d7 1 cm diameter); (g) platinum working electrode (1 \u00d71 cm).", + "texts": [ + " The initial ra tes of oxygen evolution by native or i m m o b ~ thylakoid membranes were monitored with a Clark electrode as described previously [17]. The 393 reaction m e d i u m conta ined 50 m M sodium phosphate , p H 7.1, 10 m M NaCI, 5 m M MgCI2, 600 p M 2,6-dichlorobenzoquinone and thylakoids at a chlorophyl l concentrat ion of 11 / . tg/ml. Immobil iTed prepara t ions were homogenized in a mor ta r wi th the react ion m e d i u m before use. W e ,lsed the horizontal single compar tmen t electrochemical cell descr ibed in Fig. 1. The .cell chamber (80/~1) was fi l led wi th a suspension of thylakoid m e m b r a n e s or samples of immobi l i zed m e m b r a n e s sliced at the proper d imens ions f rom the layer of sponge-like mat r ix mater ia l prepared in the petr i dishes. The electrolytic m e d i u m contained 50 m M sodium phospha te buffer, p H 7.0, 0.15 m M NaCl , and 1 m M MgCI2. The chlorophyl l concentra t ion was 250 / t g / m l unless otherwise specified. The cell was equi l ibra ted at 2 3 \u00b0 C before measurements " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003159_12.470666-Figure12-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003159_12.470666-Figure12-1.png", + "caption": "Fig. 12 Intermediate gear wheel", + "texts": [], + "surrounding_texts": [ + "As mentioned above the liquidus point of brass (CuZn37) is at about 920 \u00b0C whereas the stainless steel S 20AP melts at a much higher temperature (about 1400\u00b0C). Hence, the process has to be controlled in a way that the steel part reaches first the melting temperature and the energy is transferred by means if heat conduction to the brass. This is shown in Fig. 1. Due to the lower absorption coefficient of brass this effect is achieved by positioning the laser spot slightly on the steel axis. Most of the energy is absorbed in the steel and the brass is heated by heat conduction only. Material Axis S20AP Wheel CuZn37 \u2205 0.3 mm Material Inner wheel S20AP Wheel CuZn37 \u2205 1.9 mm Proc. SPIE Vol. 4637 579 Downloaded From: http://proceedings.spiedigitallibrary.org/ on 06/16/2016 Terms of Use: http://spiedigitallibrary.org/ss/TermsOfUse.aspx" + ] + }, + { + "image_filename": "designv11_65_0001653_s0003-2670(00)84617-6-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001653_s0003-2670(00)84617-6-Figure1-1.png", + "caption": "Fig . 1 . Schematic diagram of glow discharge apparatus : (a) lateral view ; (b) cross-section ; (c) polypropylene film sandwiched between two perforated masks situated between two plate electrodes.", + "texts": [], + "surrounding_texts": [ + "An enzyme immunoelectrode for the amperometric determination of serum insulin is described . The device consists of an immunoreactive membrane combined with a hydrogen peroxide electrode . The surface of a microporous hydrophobic polypropylene membrane is modified by water vapour plasma treatment to make it partially hydrophilic . Subsequent treatment with octamethylenediamine and glutaraldehyde enables the surface of this membrane to interact with various proteins . Anchoring of the antibody to Protein A immobilized on the membrane was effective for immunoreactivity. For more than a decade, various enzyme immunoelectrodes for assaying different antigens such as serum albumin [ 11, insulin [ 1,2 ], chorionic gonadotropin [3], a-foetoprotein [4] and hepatitis B surface antigen [5] have been described, but none was suitable for practical use. Problems associated with the use of such electrodes have been discussed [6] . With regard to the membrane itself, there are also problems related to both preparation and availability . Recently, a reproducible membrane applicable for immunoassay, especially for use on an enzyme immunoelectrode, was prepared by means of water vapour plasma treatment. In this paper, details of an enzyme immunoelectrode consisting of a microporous, partially hydrophilic polypropylene membrane and a hydrogen peroxide electrode are described . This electrode has permitted the assay of serum insulin with a sensitivity comparable to those of enzyme immunoassay and radioimmunoassay ." + ] + }, + { + "image_filename": "designv11_65_0003074_cdc.1996.572892-FigureI-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003074_cdc.1996.572892-FigureI-1.png", + "caption": "Figure I Aircraft with canard", + "texts": [], + "surrounding_texts": [ + "Output tracking in sliding modes is considered for nonminimum phase nonlinear systems. A sliding mode controller design with a dynamic sliding manifold is developed. The desired linear output tracking is provided in a dynamic sliding manifold. The dynamic sliding mode controller combines features of a conventional sliding mode controller and of a conventional dynamic; compensator. It is insensitive to matched nonlinearities and disturbances, and it rejects unmatched disturbances regardless of the nonminimum phase nature of a plant. The Approach developed is applied to a sliding miode controller design used to track the angle of attack of aircraft or missiles configured with canards.\nIntroduction\nMany missiles and a growing number of aircraft are controlled in pitch by aerodynamic surfaces placed ahead of the gravity center, referred to as\ncanards. This configuration is advantageous compared to\nthe conventional elevator or rear stabilator configuration. The reason is that the lift Fcanard produced by the canard adds up to the wing lift, F- instead of being subtracted in the case of conventional configurations. The problem posed by the canard configuration is that the control of the angle of attack is nonminimum phase [ 111, and its internal dynamics are unstable [l]. A linear SI[SO control system with an input-output transfer functiion having zeroes in the right half of the complex plane: is nonminimum phase. The nonminimum phase nature of a plant restricts the application of such powerfid nonlinear control techniques as feedback linearization control [l] and sliding mode control [2-.3]. The redefinition of the output [4] is usually required in order to apply feedback linearization control and sliding mode control to a nonminimum phase ourput tracking. This output redefinition leads to a system with a stable zero dynamics (minimum phase). Provided that the new output is accurately tracked, the original output is also adequately tracked [ 1 11. This work proposes to combine sliding mode controllers with dynamic sliding manifolds [5-71 to perform directly the traclung of the nonminimum phase output. This proposed approach does not require any rediefinition of the output. The designed dynamic slidmg mode controller combines features of a conventional sliding mode controller and of a conventional dynamic compensator. As such it is insensitive to matched disturbances and nonlinearities and it rejects unmatched Idisturbances regardless of the nonmintmum phase nature of a plant in a nonminimum phase output tracking. The tracking of the angle of\n207 1 0-7803-3590-2/96 $5.00 6> 1996 IEEE", + "attack in aircraft with canard configuration is considered. Canard configuration leads to a nonminimum phase in tracking of angle of attack [ 11). The sliding mode controller with a dynamic sliding manifold is designed. The computer simulation confirms the desired tracking of angle of attack in the dynamic sliding manifold.\nThe following nonminimum phase plant is considered:\nx=Ax+F(x,t)+bu; y=Gx. (1)\nIn Equation (l), x ER\" is a state vector, U ER' is a control function, y E R' is the controlled output, A, b, G are constant matrices of corresponding dimensions, {A,b] is is a controllable pair, and finally y*(t) ~ R ' i s\nthe given current reference input. F(x,Q e R n is a nonlinear time-dependent vector-function, it is split as F(x$ = FI(x)+ F2(t) , where FI(x) : ll&(x)ll< M is a matched unknown nonlinear function and F2(t) :\nIlF2(t)llI N is an unmatched and smooth enough unknown disturbance. The sliding mode controller is specified as:\n3(x, e, t) = 0 is the equation of the slidmg manifold. It is defined as a dynarmc operator, operating upon the state variable vector xand upon the output tracking error\nThe control is defined by discontinuous control function, uz+ U; of x, t . The following goals need to be met: + The output y ( t ) is required to track asymptotically the given reference output trajectory y*( t ) . That is:\n~i+*(t) - y(i/ l= lidle(t)ll= o as t -+ cc, . + The behavior of the tracking error e( t ) in the sliding mode must be linear and desirable. + The existence of the slidmg mode in the designed dynamic sliding manifold must be provided.\ne(t)=y*(t)-y(t).\nProblem solution\nThe system is transformed to a block-canonical regular form [2,3], using the well-known nonsingular\ntransformation, { x ~ , x l J ' = ~ x . After the\ntransformation, the system (1) is written as follows:\nI In Equation (3), x1 E Rn-',x2 E R , b2 f 0 , and the transfer function of Equation (4), has zeros at the right hand half of the complex plane.\nA direct introduction of a sliding mode in the system (3) is impossible due to the nonminimum phase nature (4) of the plant [ 1-41. The algorithm of a dynarmc slidmg manifold design for a nonminimum phase plant (3) is developed as follows:\nDynamic sliding mode controller\nStet, one consists in the designing of the sliding manifold as represented by Equation (5).\nThe condition of existence of sliding mode [2-31 3- 5 < -dZf must be met in the vicinity of the sliding manifold (5). This is obtained for the system (2-3) as follows:\nIn Equation (6) b, is assumed to be positive and p is any positive numkr. Remark 1: It is obvious that U+ and U - , which satisfy the existence conditions (6) are bounded (realizable) if o is differentiable. The sliding motion in the manifold ( 5 ) is stable, and II&l(x, X, )I/ is bounded in the sliding domain. The equations of sliding of the system (3) in the slidmg manifold ( 5 ) are obtained as follows [2-31:" + ] + }, + { + "image_filename": "designv11_65_0000785_s0094-114x(98)00074-3-Figure3-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000785_s0094-114x(98)00074-3-Figure3-1.png", + "caption": "Fig. 3. Contact lines, envelope Er to contact lines on generated surface Sr and tangent Te.", + "texts": [ + " Equation of meshing (19) yields that ru, ry, and rf lie in the same plane, and therefore rf \u00ff aru \u00ff bry 0: 32 It can be proven as well [8] that ff \u00ff afu \u00ff bfy gr N r r 2 0; 33 gf \u00ff agu \u00ff bgy hr N r r 2 gu gy gf r2u ru ry ru rf ry ru r2y ry rf 1 N r r 2 \" # : 34 Here, hr$0 if Hr$0, and a ru rf ru ry ry rf r2y N r r 2 ; b r2u ru rf ry ru ry rf N r r 2 35 where Nr (r)= ru ry. Finally, we can obtain that Te hr ru ry 2 fu fy gu gy Tr 36 where Tr fyru \u00ff fury 37 is the tangent to the contact line on Sr. Eq. (36) con\u00aerms that curve (21) formed by singular points on Sr is the envelope to the contact lines. Figure 3 shows the contact lines on the generated surface Sr, and curve Er that is formed by singular points on Sr. The drawings show that Er is the envelope to the contact lines We emphasize that an in\u00aenitesimal displacement along curve Er is not equal to zero since Er is a regular curve. However, an in\u00aenitesimal displacement over surface Sr in any direction that di ers from Te is equal to zero, since a point of curve Er is a singular point of surface Sr. Note: The proof that envelope Er to contact lines on generated surface Sr is simultaneously the edge of regression is based on the following considerations: 1" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001017_0021-8928(96)00031-7-Figure2-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001017_0021-8928(96)00031-7-Figure2-1.png", + "caption": "Fig. 2.", + "texts": [ + "1) Also, let the final ,desired state of the system be xj(0)=xlo, x2(0)=0, x3(0)=x3o, x4(0)=0 (2.2) We shall now fon~ulate the problem of finding the control momen t s ~tl(X), ~2('1;) which ensure that the two-link mechanism transfers from its initial state (2.1) to its final state (2.2) in the minimal time 0. We shall assume that x3o =-X3o, X3o/> 0 (2.3) that is, that the initi~d configuration (2.1) and the final configuration (2.2) of the two-link mechanism are symmetric, one to the other, with respect to the line OS, the equation of which has the form Xl = x10/2 (Fig. 2) in polar coordinates. Problems of the optimal control of a two-link mechanism have been studied in many investigations. H=~/I x2-1~4~ + \u00a52~tl + ~1/3x4 + ~1/4/.~ ~2 _ ..~.~ ~tl [-~L tJt ~ + T ( x 2 C s i n x3 2 + 13~x42 ) | q/ (2.4) The conjugate variables satisfy the equations ~1 =0, ~/2 =- -~L-2C~4x2 sinx 3 et8 o d(1 / 8) ~lJ\" 3 -~ --~IX2 + ~lllX 4 dx 3 d(13 / 8) _ ~1/4~2 d(8 / et) dx 3 dx 3 + ~ 4 1 x I - - d(l~ / ~) dx 3 d (sinx~'~ x2 d fl3ysinx~ (2.5) 13y sin x 3 ~'4 = \u00a52 - \u00a53 -2C~4x4 t~8 It follows from the maximum principle that the optimal control must satisfy the conditions pq (x) = sign ~t/(x), la2(x) = ~t20 sign ~4(x) (2" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000314_bf00167941-Figure9-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000314_bf00167941-Figure9-1.png", + "caption": "Fig. 9a, b. Stochastic simulation of receptor-mediated cortical actin: transient polarity with mode 1 dominating. Numerical integration results for the Gauss limit SDE system are for //2 (Table 1, n = 20). a is plotted as in Fig. 8 for a a* = 0.7 and b a* = 0.8", + "texts": [ + " 2a that the uniform solution a(t, x) = a* to (20) is unique and linearly stable for HI, and that modes 1 and 2 are equally least stable, i.e. 21 = 22 ~ - 2 (21 = - 1.86, 22 = - 2.16). The effect of the fluctuations in the distribution of bound receptors and, hence, a*, from its expected value (constant and spatially-uniform) is seen to induce a complex spatialtemporal pattern of the cortical actin distribution involving a mixture of predominantly mode 1 and 2 pattern as predicted from the eigenvalues. We present simulation results in Fig. 9a for/I2, a* = 0.7 and Fig. 9b for /72, a* = 0.8. Recall from Fig. 2b that for/I2 the uniform solution a(t, x) = a* is again unique and linearly stable, with mode 1 least stable, but only for a* < a* r ~ 0.875 (Fig. 3). For a* = 0.7 (Fig. 9a), the results are qualitatively the same as for Fig. 8 but with mode 1 dominating. However, the results are qualitatively different for a* = 0.8 (Fig. 9b). Even though (a*) = 0.8 < 0.835 (below the turning point value, a*p ~ 0.835, seen in Fig. 3), the fluctuations in a* clearly excite the polar solution that exists for a* > a~ in a sustained way over t ~ 20 - 30 and again after t ~ 35 (note that the same \"seed\" value was used in the stochastic simulations generating Figs. 9a and 9b, the seed value completely defines the random deviate sequence needed to approximate a realization of dw(t) and integrate (32), and with the Cauchy-Euler method, the same realization of dw(t) is generated when the same values for the seed, h, to, and t I are used). The fluctuations of a* are calculated to have a r m s deviation of 0.050 from (18), and upon inspection of the bifurcation plot (Fig. 3) are consequently often large enough to transiently raise a* above a~ at any position, x; hence, the \"excitation\" of the polar solution. Finally, simulation results in Fig. 10 are presented for the same parameter values as in Fig. 9a (//2, a* = 0.7) but for a different seed and values of d that are also 0.01 (Fig. 10a) and 100 (Fig. 10c) times the value used in the Fig. 9a simulation. It is clear that the same realization of dw(t) induced very different spatial-temporal pattern characteristics, with both the amplitude and longevity of fluctuations appearing to decrease as d increases. In spite of the complex spatio-temporal pattern characteristics, the mode 1 shape of the a-profile is maintained in each simulation, so that the position of its minimum, go(t), corresponding to the leading lamella, defines an empirically observable process of directional shifting, i.e. a persistent random walk with rapid excursions superimposed on transient drift to either side. This suggests that the speed of the shifting process, s(t) = dgo/dt, has a zero mean value but nonzero variance that increases as d increases. 4.3.1.2 Quantitative behav ior - correlation functions. The large amplitude variations in a(t, x) in Fig. 9b indicate an essential influence of the nonlinearities of (20); however, for Figs. 8 and 9a, the transient small amplitude fluctuations suggest that the linearized form of (20) can be analyzed to yield ins ight in to the s p a t i a l - t e m p o r a l cor re la t ion s t ructure under ly ing these representa t ive s imula t ions . (In fur ther suppo r t of these s ta tements , whereas the con tou r p lo t for a s imula t ion co r r e spond ing to Fig. 9a for 7 reduced by a fac tor of 10 is essent ial ly ident ical , t ha t co r re spond ing to Fig. 9b shows a less p ronounced peak at t ~ 24; the max imum peak-to-peak values of a were reduced by more than 90% in each case for the reduced value of 7.) It is impossible to discern any systematic relation between the plot for a and the plot for a* (none shown) in any given simulation, so such an analysis is, in fact, necessary. (a) State diffusion approximation (Gauss limit). Parametr ic dependencies of the s ta t ionary statistics of the stochastic variables a(t), f(t) and b(t) can be analytically determined in terms of two-point auto/cross-correlation functions, as was done in characterizing the properties of the stochastic receptor model alone (Fig", + " For clarity, we omit the subscript \"s\" that denotes stationarity from all variables in the remainder of this section. ~O,Q 5.0 / II, 2.5 ~ 0.0 i 3.o\" ,\u2022 1.0\" 0.0. ~\" .~o. Fig. l la , b. Two-point stationary correlation functions for receptor-mediated cortical actin variables: transient polarity with mode 1 dominating. Analytical results in the Gauss limit are for the//2 parameters (Table 1, a* = 0.7, n = 10). See Fig. 5 for the related receptor correlation functions Relevant two-point correlations involving yai(t) and ybi(t) are presented in Fig. 11 first for the simpler case of Fig. 9a corresponding to 112, a* = 0.7 (n = 10 was used to obtain all correlation results in the Gauss limit because A' is ill-conditioned when n = 20). Symmetry noted previously for the receptor correlations in Fig. 5 is evident again in the two-point correlations for Ya in Fig. 1 la. Based on the simulation where a polar morphology dominates, the expected dependence of the two point correlation for Ya as a function of spatial distance at any correlation time, z, is obtained: (Yal (0)yai(z)) decreases from a maximum value for i = 2 and 10 (adjacent to the (arbitrary) reference compartment) to a minimum that is negative at i = 6 (the position halfway around the cortex)", + " Thus, the (linearized) morphogenetic 4th-order parabolic equation serves as a spatial-temporal filter selecting, for the parameter values investigated here, the (s)lowest modes. The two-point correlations for the case of Fig. 8 corresponding to H1 (n = 10) are now presented in Fig. 12. Recall that in this case modes 1 and 2 are equally least stable. The contribution from mode 2 is apparent from the nature of spatial correlations at z = 0, although it appears to decay rapidly in time with respect to the mode 1 contribution. Although the pattern properties are perhaps more interesting in this case, we return to the more relevant case for P M N of Fig. 9a (i.e. a predominantly polar pattern) and present one-point correlations (i.e. only autocorrelations for i = 1) assessing the influence of varied d. Fig. 13 presentsthe dependence on d for values 0.01 and 100 times the base case//2 value of 0.152 (for n = 10). Recall from Fig. 6a that the variance (y21) is not affected (as expected from (50)), but the autocorrelation for Yba is affected (as expected from (49)). Consistent with this \"smoothing\" effect of enhanced receptor diffusion relative to binding, both the joint second moment and cross correlation between Yal and Ybl decreases as d is increased [13b], and the same is true for the variance and autocorrelation time for Yal [13a]", + " 14 presents results to show the dependence on ( a * ) below the bifurcation value alp ~ 0.875. It is well-known that the magnitude of deviations increase dramatically and become correlated extremely long in time when a phase transition point is approached (Gardiner et al. 1976). Since alp represents a phase transition point (stable nonpolar to stable polar cortical actin distribution and extension), the results of Fig. 14 are consistent with this phenomenon. They are also consistent with the trend indicated in Fig. 9. (b) State and space diffusion approximation (Poisson limit). In order to understand the properties of the correlation functions computed in the Gauss limit in analytical terms, we derived formulas in the Poisson limit (n ~ R r -~ oo ) that reveal parametric dependencies. The linearized stochastic partial differential equation governing a(x, t) from (20) then involves a continuous diffusion process for a*(x, t), which is proportional to B(x, t), the density of bound receptors. Let ( be the Fourier transform variable of corre la t ion time, z" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000602_978-94-009-1718-7_41-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000602_978-94-009-1718-7_41-Figure1-1.png", + "caption": "Figure 1. Typical construction of a fully parallel manipulator with planar base and end effector platforms.", + "texts": [ + " Introduction Parallel manipulators have received growing interest over the last decade, due to some superior characteristics with respect to serial manipulators: rigidity, dynamic characteristics, accuracy, etc., Fichter (1986), Hara and Sugimoto (1989), Hunt (1983), Kerr (1989), Merlet (1989, 1992), Nair and Maddocks (1994), Shi and Fenton (1994). Fully parallel manipulators con sist of six serial kinematic chains (\"legs\") between a fixed \"base\" and a moving \"end effector.\" Each leg has six joints, but only one is actuated. In most designs, the actuated joint is prismatic (Fig. 1). The inverse position kinematics for a fully parallel manipulator finds the lengths of the six legs when the end effector position and orientation are given. As is well known, this problem turns out to be simple, and yields unique solutions. The forward position analysis, on the other hand, tries to find the end effector pose when the leg lengths are given. In general, this problem involves highly non-linear equations, and closed-form solu tions are only known for a couple of special designs. Moreover, \"closed form\" most often means that polynomials of order eight or sixteen have to be solved, Griffis and Duffy (1989), Husain and Waldron (1994), Inno centi and Parenti-Castelli (1989), Nanua et al", + " The above-mentioned class of platforms with similar base and end effector, Lee and Roth (1993), Sreenivasan et al. (1994), Wang (1992), is a special case of the more general class presented in this paper. Overview of the Paper. Section 2 describes the mathematical representa tion of a general manipulator's geometry, and of its loop closure equations. Section 3 then explains which relations between the design parameters of the manipulator give rise to closed-form solutions of the forward kinemat ics, and how these solutions are constructed. Finally, Section 4 gives some examples. 2. Geometry Figure 1 shows a general fully parallel manipulator with six legs, and pla nar base and end effector platforms. The right-handed, orthogonal ref erence frames {bs} and {ee} are fixed to the base and to the end ef fector, respectively. The unit vectors along the axes of these frames are ex, ey, ez, and f x' f y, fz, respectively. Only the prismatic joints along the legs are actuated, and only their joint positions (\"leg lengths\") are mea sured. Their attachment points to the base are denoted by the position vec tors bi , i = 0, " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000854_s0389-4304(99)00024-7-Figure12-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000854_s0389-4304(99)00024-7-Figure12-1.png", + "caption": "Fig. 12. Reference \"gure for lever ratio of double-wishbone suspension.", + "texts": [ + "erence in values obtained between cases executed from a direction that eliminates the e!ects of such bush friction and cases when the vehicle body is simply pushed downward from the initial condition. 3.3. Vehicle application requiring leverage consideration (such as double-wishbone suspension vehicles) In such cases, application is possible by adding the correction equation described below. Damping force Spring constant\"Output of unit/g2 (for product with double-wishbone suspension) (Here, leverage g\"B/A) In Fig. 12, A length of lower arm B length from lower-arm body-side pivot position to shock absorber unit installation position Strut-type and other suspensions tend to have leverages (lever ratios) that are close to 1; thus, output can be obtained that is nearly equal to the values measured during single-item tests (Fig. 13). Speci\"cally, the strut-type shock absorber has a slight inclination (about 5 to 15 degrees), but the e!ect on vertical movement is only approximately between 1 and 4 percent; and because the inclination angle is small, the e" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000836_bf02481485-Figure7-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000836_bf02481485-Figure7-1.png", + "caption": "Fig. 7 Hierarchical trajectory planning for a robot manipulator", + "texts": [ + " However, if the robot has no motions and skills, it generates new motions by motion planning based on inverse kinematics and inverse dynamics. Thus, a robot has no skills in its initial state, but gradually acquires skills and motion through interaction with the environment. Redundant manipulator with structured intelligence A robot basically generates new motions by GA. We have proposed various trajectory planning methods for redundant manipulators. H One of them is hierarchical trajectory planning, composed of configuration generation as lowlevel planning and trajectory generation as high-level planning (Fig. 7). The hierarchical trajectory generation method for intelligent robots is based on the concept of external and internal evaluations (Fig. 8). External evaluation is basically used for generating primitive motions (configurations) through interaction with the environment. Internal evaluation is a criterion for binding the robot's primitive motions recursively. This criterion dynamically changes according to the total evaluation, that is, the criterion for selecting a configuration becomes high, as the best trajectory gives high performance" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003977_146441905x9980-Figure2-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003977_146441905x9980-Figure2-1.png", + "caption": "Fig. 2 Switching sections and basic mappings in the phase plane: (a) x1 \u00bc+E; (b) x2 \u00bc+E", + "texts": [ + " S \u00fe P4: S \u00fe ! S , P5: S ! S \u00fe , P6: S ! S\u00fe (11) for x2 \u00bc +E. From the definition of mappings, the mappings Pq (q \u00bc 1, 2, 4, 5) relative to one switching section are termed the local mapping, and the mappings Pq (q \u00bc 3, 6) relative to two switching sections are termed the global mapping. The global mapping transports the motion from one switching section into another switching section. The local mapping is the self-mapping in the corresponding switching section. Six mappings are illustrated in Fig. 2. To describe the complicated motion, the mapping structure of dynamical systems in equation (1) is used herein. For simplicity, the notation for mapping is introduced as Pn1n2...nk ; Pn1 WPn2 W WPnk (12) where Pni [ {Pqjq \u00bc 1, 2, . . . , 6} and ni \u00bc {1, 2, . . . , 6}. Note that the rotation of the mapping of periodic motion in order gives the same motion (i.e. Pn1n2...nk ,Pn2...nkn1 , . . . ,Pnkn1...nk 1 ), and only the selected Poincare mapping section is different. The motion of the m-time repeating of mapping Pn1n2 nk is defined as Pm n1n2 nk ; (Pn1 WPn2 W WPnk )W W(Pn1 WPn2 W WPnk )|fflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflffl{zfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflffl} m sets (13) To extend this concept to the local mapping, define Pm 12 ; (P1WP2)W W(P1WP2)|fflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflffl{zfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflffl} m sets and Pm 45 ; (P4WP5)W W(P4WP5)|fflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflffl{zfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflffl} m sets (14) Symmetry of steady-state solutions of non-smooth dynamical systems 111 K02704 # IMechE 2005 Proc" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003752_s0141-6359(04)00045-5-Figure8-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003752_s0141-6359(04)00045-5-Figure8-1.png", + "caption": "Fig. 8. The experimental system setup.", + "texts": [ + " (1) and (2) can be represented by phase 1, for |\u03c3|n/\u03bb > xh, (the creep phase): d dt xs x\u0307s xh x\u0307h xp x\u0307p = x\u0307s( 1 + \u03b1n \u03bb |\u03c3|n\u22121Cs )\u22121 ( F \u2212 \u03c3 m \u2212 q ) \u03b1|\u03c3|n \u03bb \u2212 \u03b1xh sgn(\u03c3) ( 1 + \u03b1n \u03bb |\u03c3|n\u22121Cs )\u22121 ( \u03b1n \u03bb |\u03c3|n\u22121Cs F \u2212 \u03c3 m + q ) sgn(\u03c3) ( \u03b1|\u03c3|n \u03bb \u2212 \u03b1xh ) ( 1 + \u03b1n \u03bb |\u03c3|n\u22121Cs )\u22121 ( \u03b1n \u03bb |\u03c3|n\u22121Cs F \u2212 \u03c3 m + q ) (3) phase 2, for |\u03c3|n/\u03bb \u2264 xh, (the hardened phase): d dt xs x\u0307s xh x\u0307h xp x\u0307p = x\u0307s F \u2212 \u03c3 m 0 0 0 0 (4) where \u03c3 = Csx\u0307s + k1(xs \u2212 xr) +sgn(xs \u2212 xr) k2 \u03b2 (1 \u2212 e\u2212\u03b2|xs\u2212xr |)+ \u03c3r (5) q = \u03b1n \u03bb |\u03c3|n\u22121(k1 + k2 e\u2212\u03b2|xs\u2212xr |)x\u0307s \u2212 sgn(\u03c3)\u03b1x\u0307h (6) and Cs is the damping coefficient of the system. Then, the measured output displacement x is: x = xs + xp (7) Combining the above equations with the reverse point rules, the response of system subjected to static friction can be obtained. The parameter values presented in the equations can be identified from the experimental data. The experimental system is the stage shown in Fig. 8 and the experiment is performed following the test method proposed in [7]. A typical experimental response of static friction compared with the theoretical data is shown together in Fig. 2. The result shows consistency between the real system and the model. In this study, experiments were conducted several times and the parameter values were varied w.r.t. position and time. This supports the time and position dependency reported in In fact, the static friction behavior is similar to the mechanical structural behavior and can be explained by the asperity model [4] in Fig", + " (28) and from the definition of sliding surface, the closed-loop system equation can be obtained as: e\u0308+Kve\u0307+Kpe = s = \u2212W \u222b sgn(s) dt + 1 m \"Ff (32) This reveals that the characteristic of the closed-loop system is now governed by e\u0308+Kve\u0307+Kpe. The corresponding approaching condition can be satisfied by: ss\u0307 = s ( \u2212W sgn(s)+ 1 m \"F\u0307f ) = \u2212W |s| + 1 m s\"F\u0307f \u2264 |s| ( \u2212W + 1 m \"F\u0307f ) \u2264 0 (33) That isW > \"F\u0307f/m. The experimental system used in this study is a two-axis ball-screw-driven stage actuated by servo motors as shown in Fig. 8 associated with the control devices. A schematic diagram of a single axis stage is shown in Fig. 9. The stage displacement is measured by a grid encoder mounted on top of the table surface. A low-pass filter was introduced to reduce the noise effects and a 10 nm resolution could be obtained during operation. The control was implemented on an industrial computer with a sampling rate of 2000 Hz. The corresponding controller parameters are listed in Table 3. Two controllers specified as the static and the dynamic friction controllers were developed in the preceding sections" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000007_piae_proc_1922_017_075_02-FigureI-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000007_piae_proc_1922_017_075_02-FigureI-1.png", + "caption": "FIG. I8 . -ha r hub for 25/50 h.p. car.", + "texts": [], + "surrounding_texts": [ + "374 THE INSTITUTION OF AUTOMOBILE ENGINEERS.\nThe point arises here as to how the hand of the gear should be selected for the bevel or crown-wheel so as to impose the heavy thrust on one or the other. The conclusion just reached regarding the double-purpose unit at the crown-wheel is on the assumption that the light thrust load oome's in this position, and this is in accordance with general practice. In order to make full use of the possibilities of the single-row bearing, however, the hand of the gear might be reversed, and certainly there will, in many cases, be more opportunity for obtaining a large capacity in the crownwheel position than behind the bevel-pinion\nPilot Bemirig. Although the pilot bearing, Fig. 20, does not appear on any of\nthe graphs, it will be remarked that it is in many casea an& example of a bearing applied to relieve a heavily loaded one without kaking into account the possibility of its own failure, which frequently occurs.\nClutch. The bearing selection for the clutch is not so ainenable to the same treatment, but it will usually be found that a light type single-row bearing w i t b u t fillibg-slot has su5cient capacity .to deal with a load which is only in occasiond operation.\nFigs. 17 to 20 show the aotual arrangements of the front-hub, rear-hub. gear-box and rear-axle bearings of the car referred to as No. 12.\nSUMMARY. The difficulties in the way of fixing definite factors of safety which will meet all cases are fully appreciated, but it is hoped that the preceding recommendatipns will form a useful basis for investigating existing designs. For convenienoe, the factors have been collected in Table I., the letters in Fig. 15 qorres o d ' with those used thmughout the drawings of the Sizaire-brm? chassis.\n2016 at Kungl Tekniska Hogskolan / Royal Institute of Technology on June 5,pau.sagepub.comDownloaded from", + "THE ENDURANCE ,OF BALL HEARINGS. 375\n2016 at Kungl Tekniska Hogskolan / Royal Institute of Technology on June 5,pau.sagepub.comDownloaded from", + "376 THE INSTITUTION OF A UTOMORILE ENGINEERS.\nFIG. lg.-Gear-bux for 25/50 h.p. car.\n2016 at Kungl Tekniska Hogskolan / Royal Institute of Technology on June 5,pau.sagepub.comDownloaded from" + ] + }, + { + "image_filename": "designv11_65_0002040_(sici)1099-1581(200002)11:2<69::aid-pat938>3.0.co;2-v-Figure5-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002040_(sici)1099-1581(200002)11:2<69::aid-pat938>3.0.co;2-v-Figure5-1.png", + "caption": "FIGURE 5. CVA for Si electrodes in the reaction solution after 12 104 cycles in PP regime: (a) on p-Si \u00d0 in the presence of CAA (An:CAA = 50) and - - - in the absence of the catalyst; (b) on n-Si \u00d0 in the presence of the catalyst (An:PDC = 2000) and - - - in the absence of the catalyst; v = 20mV/sec.", + "texts": [ + " 4d, curve 2). Figure 4 exhibits the CVAs for resulting polymer coatings registered in the working solution for p-Si (Fig. 4a) and n-Si (Fig. 4b) during the catalytic synthesis in the absence IrCl2\u00ff6 (curves 1) and the presence of IrCl2\u00ff6 (curves 2). Potential Pulse (PP) Synthesis of PAn. This regime provides 0.3\u00b10.8 V pulses with 1 10\u00ff2 and 2 10\u00ff2 sec duration for 3\u00b112 104 cycles. The application of the PP regime allowed the preparation of the visually most dense homogeneous coatings. It can be seen in Fig. 5 that the introduction of the catalysts (CAA and PDC) to the electrolyte accelerated the formation of a polymer coating in the catalytic PP synthesis. Copyright \u00e3 2000 John Wiley & Sons, Ltd. Polym. Adv. Technol., 11, 69\u00b174 (2000) The study of the electrochemical behavior of PAn coatings on porous n- and p-Si electrodes in acidic media reveals some peculiarities. Firstly, current runs through naked p- and n-Si electrodes only at anodic scan: cathodic scan yields no current. Conductivity at cathodic scanning appears only with the formation of PAn coating and grows as the quantity of the polymer increases" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001850_ma00232a014-Figure3-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001850_ma00232a014-Figure3-1.png", + "caption": "Figure 3. Coordinate system used for describing the bending motion. Each segment makes an angle 0 with the z axis and the z and y axes are in the plane that bisects the angle between the segments.", + "texts": [], + "surrounding_texts": [ + "Numerous authors have dealt with the motions of a once-broken rod and the effect of these motions on the dynamic light scattering correlation functions, primarily on a theoretical However, those papers assumed that the break point was a universal joint, allowing all possible angles between the two segments. Furthermore, the dynamic light scattering theories deal only with the isotropic scattering and intramolecular scattered light interference effects. In this work, an approximate theory is presented for the scattered light intensity time correlation functions from a dilute solution of once-broken rods. Both the isotropic and anisotropic components are found, assuming negligible intramolecular interference, and the angle between segments is restricted to some maximum value. An example of a once-broken rod with a restricted intersegmental angle may be the myosin rod. The myosin molecule (Figure 1) is composed of three basic functional units.\"12 The light meromyosin fragment (LMM) is believed to be rather stiff and rodlike. Subfragment 2 (S-2) is more flexible and connects the LMM with the head group. The head group consists of two subfragment 1 (S-1) moieties. The myosin rod is the myosin molecule with the head group removed. Both electric birefringenceg and electron rnicroscopy1'J2 experiments on the myosin rod and the myosin molecule indicate a considerable amount of flexibility at the joint between the LMM and S-2 fragments, with a possible maximum intersegmental angle of 145' (ref 12, where an angle of 0' means a stiff rod with no bend). Here, the results from dynamic light scattering experiments13 are compared with our theory of the oncebroken rod in an attempt to extract the translational and 0024-9297/82/2215-1023$01.25/0 rotational diffusion coefficients and the maximum intersegmental angle." + ] + }, + { + "image_filename": "designv11_65_0000700_978-3-642-97646-9_13-Figure12.35-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000700_978-3-642-97646-9_13-Figure12.35-1.png", + "caption": "Fig. 12.35. Stator and magnetising current vectors of current source inverter drive during a speed reversal at 200 l/min", + "texts": [ + " The motor was again connected to a DC dynamome ter at no load. The behaviour of the drive at low speed is demonstrated by the transient in Fig. 12.34 where a slow speed ramp causes the pulse-width modulation to take effect around zero speed. This is clearly seen from the trace of the switching state ~(t) of the ring counter; as soon as the speed leaves the narrow band around standstill, the modulation ceases and the angle controller continues with a unidirectional switching sequence. Characteristic loci of the current vectors are seen in Fig. 12.35 during a speed reversal at very low speed. Initially the stator current vector is stepping around the six-pointed star; when, a reversal of the speed is commanded, the magnitude increases sharply and the direction of the rotation is inverted. After a few steps the speed reversal is completed and the magnitude of the current vector is reduced to its former no-load value. The locus of the magnetising current vector is again nearly unaffected by the switching operation, even though the radius of the locus is not quite as perfectly constant as in the case of the voltage source inverter drive (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001103_s0736-5845(97)00033-1-Figure8-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001103_s0736-5845(97)00033-1-Figure8-1.png", + "caption": "Fig. 8. Weighted damped least-squares approximation to unreachable commands with the Kraft slave manipulator.", + "texts": [ + " To move from c1 to c2, requires that joints 2 and 4 rotate clockwise, while joint 3 rotates counter clockwise. The result, r2, of the command c2, is not intuitive to the operator. If the command is extended further out from the base, the manipulator has a tendency to move further down, creating confusion for the operator. While the damped least-squares solution may perform satisfactorily for bondaries defined by singularities, it is clearly unacceptable for boundaries defined by joint limits. Fig. 8 uses the kinematic model of the Kraft slave to illustrate the convergence of a resolved rate controller using the weighted damped least squares methods proposed. The result of extending the command, c1 in Fig. 7, to c2 in both figures, is r2 in Fig. 8. The resulting solution \u2018\u2018points\u2019\u2019 to the commanded position. As the command is moved through the unreachable space, as shown by c3 and c4, the robot tracks along the workspace boundary, finding approximate solutions in the task space. For a human operator, the result r2 in Fig. 8 is a much more intuitive solution to the command c2 in Figs. 7 and 8 than is the result r2 in Fig. 7. Furthermore, with a bilateral controller where the force-feedback in the master tends to drive the master to the slave, the operator is further confused by the result in Fig. 7. Here the master will be driven downward rather than along a shorter path back to the workspace of slave, as results from the weighted solution. It has been determined by all of the operators of the teleoperational system, that the controller using the weighted damped least squares methods is a very significant improvement over the re-alignment and re-indexing solutions to unreachable commands", + " The motion of the slave is both smooth and stable when the manipulator contacts the environment and/or when the manipulator is in several joint limit and/or singularity regions. Operation near a singularity results in approximate velocity solutions with damped motion for the joints which swing about dangerously if an exact solution is used. Operation on the workspace boundaries results in approximate task space position solutions that track along the workspace boundaries in a manner that is easily understood by the operator. We may consider the behavior of the damping factor and joint weights using the situation depicted in Fig. 8. As the manipulator approaches the work space boundary defined by the limit on joint 3, the distance from the joint limit, l 3 , is decreasing as determined from Eq. (19). And, if this distance is smaller than the l 0 parameter, then the joint weight w q3 ramps down temporally to 0.01#0.99l 3 /l 0 as determined by Eqs. (20) and (21). Furthermore, a increases as determined by (20) and (22). This smaller joint weight and nonzero damping factor increase the importance of dq 3 in the minimization condi- tion (15)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001102_tt.3020030402-Figure7-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001102_tt.3020030402-Figure7-1.png", + "caption": "Figure 7 Schematic view of the reciprocating stick-slip tester: (a) in its original state; (b) after modification", + "texts": [ + "00 Comparison of two stick-slip testers and recommendations for repeatable and significant stick-slip testing 373 Hence, the way in which the vertical load is applied in the test equipment influences stick-slip results. A vertical load applied, for example, by means of a spring (influencing k,) or a lever system (influencing m,), can lead to different results in comparison with a load applied by simple equivalent sliding mass (dead weight) which gives the same contact pressure. The Cincinnati reciprocating stick-slip tester was used in the corresponding ASTM test until this procedure was withdrawn because of poor repeatability. The test rig (see Figure 7a) consists of a base block (l), driven in reciprocating motion by a lead screw (Z), sliding against a top block (3). The vertical load is applied by means of a calibrated spring assembly (4,8). The friction force between the two blocks is transmitted by means of rods (5) to leaf springs (6) the deflection of which (proportional to the friction force for stationary situations) is measured by dial indicators (7). RECIPROCATING STICK-SLIP TESTER: The Cincinnati Milacron test rig Tribotest journal 3-4, June 1997", + " During tests, it was observed that this spring tilts with the very small reciprocating motion of the top block. This tilting leads to a change in spring length and thus also of the vertical load acting on the top block during a test. To solve this problem, it was decided to apply the normal load by means of dead Tribotest journal 3 4 , June 1997. (3) 374 1354-4063 $10.00 + $4.00 Comparison of two stick-slip testers and recommendations for repeatable and significant stick-slip testing 375 weights (8 in Figure 7b), producing a normal load of 220 N in accordance with the ASTM test procedure. As a result of this modification, a constant normal load is guaranteed during the tests. However, the vertical stiffness (k,,, Figure 3a) disappears by removing the spring. This important change of the vertical stiffness can influence the stick-slip results (see above). For example, in the case of smooth sliding (F, smaller than Fk) reduction in vertical stiffness will tend to increase the stick-slip number. The kinetic friction force F k is measured in a situation of larger normal separation of the rubbing surfaces compared to the static friction force F , (as described earlier)", + " If the stick-slip numbers to be determined are greater than 1 (risk of stick-slip), electronic measurement of mass displacement by means of strain gauges is recommended. With electronic measurement, the mass displacement signal produced by the strain gauges can be sampled at a sufficiently high frequency (e.g. by a data-acquisition card) to record the stick-slip motion. As an extra advantage, the behaviour of lubricants during start, stop or velocity reversal (see below) can be observed. A load cell (9 in Figure 7b) was mounted between the frame and the top block of the test rig. In fact, a load cell is a relatively stiff linear spring on which strain gauges are fixed which record its deflection. Hence, the load cell also measures the mass displacement and not the friction force during stick-slip. The replacement of the leaf springs by the load cell increases the tangential stiffness k, (from 80 kN/m to 425 kN/m). This modification affects stick-slip numbers greater than 1 (see above), which could not be determined accurately by means of the dial indicator" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001579_0022-4898(87)90009-7-Figure6-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001579_0022-4898(87)90009-7-Figure6-1.png", + "caption": "FIG. 6. Center of turning of the track and the resistant force under turning motion.", + "texts": [ + " 7) B - - Distance between tracks P - - Ground contact pressure Pi - - Ground contact pressure for inside crawler Po - - Ground contact pressure for outside crawler P - - Lateral force acting on elemental area of crawler # - - Longitudinal frictional coefficient #t - - Lateral frictional coefficient F - - Thrust force Fo - - Rolling resistance for outside crawler in straight travel - - Rolling resistance for inside crawler in straight travel Mt - - Moment against turning Mf - - Moment for front part of crawler about the point ~O' (Fig. 6) Mr - - Moment for rear part of crawler about the point O' (Fig. 6) O - - Geometrical center of the crawler (Fig. 6) Zo - - Thrust for outside crawler Zi - - Thrust for inside crawler Mo - - Moment for outside crawler Mi - - Moment for inside crawler Wo - - Load acting on outside crawler Wi - - Load acting on inside crawler Zo' - - Thrust for outside crawler when L is changed to L' *Presented at the First Asian-Pacific Conference of ISTVS, Beijing, China, August 1986. \"['Professor, Depar tment of Agricultural Machinery, Mie University, Tsu, Mie 514, Japan. 251 Zi' - - Thrust for inside crawler when L is changed to L' Wo' - - Load acting on outside crawler when L is changed to L\" Wi' - - Load acting on inside crawler when L is changed to L' 1 HE TRACKED vehicle is popular for performing operations on soft, loose ground because of its better manoeuverability", + " Some of the examples of practical mechanisms of controlling the contact length of the crawler are shown in Figs 3-5. The turning resistance moment can be varied depending on the location of the center point FIG. 2. OUTSIDE TRACK dx p 112 1 /2 I U S I D E ( b r a k e d ) TRACK Resistant force acting about the inside (braked) track under steering motion. IDLER SPROCKET TRACK ROLLER , . . . . ~ J / ~ . . . . H - b FIG. 3. Method of varying the contact length of the track. of turning, even if the contact length is kept same during the turning. Referring to Fig. 6, let us consider the turning resistance moment Mr where the center point of turning is located at the point O' being a distance x' away from the point O, the geometrical center of gravity of the crawler. Substituting l = L / 2 - x ' and l=L/2+x' into equation (2), equations (3) and (4) respectively can be obtained. Mr = #tpb(L/2-x ' )2/2 (3) Mr = #~pb(L/2+ x')2 /2. (4) Therefore the total turning resistance moment about the point O' can be expressed as the sum of Mf and Mr. M = Mr + Mr = u~pb((L/2-x') + (L /2 + x')2)/2 = u : b ( ( L / 2 ) + 2x'2)/2" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000431_s0094-114x(97)83002-9-Figure2-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000431_s0094-114x(97)83002-9-Figure2-1.png", + "caption": "Fig. 2. A schematic of the test manipulator with sensor locations shown.", + "texts": [ + " In particular, r/~ k- ~) and r/) k) represent solutions (deflections) at the beginning and end of the kth interval, respectively, and a) k), b~ k) and c) k) [13] are dependent upon the modal parameters, modal forces and time length for step k. The displacements at physical stations, or coordinates, then become {x} (k, ~ r/(k) , r ~ (k) = J t ,J ( 5 ) j = l The procedure discussed above was applied to the investigation of a robotic manipulator with a revolute and prismatic joint, as sketched in Fig. 2. The aluminum arm, considered to be an elastic beam (2\" x 0.0625\") due to its slenderness, can move axially with respect to the sliding constraint, which is shaped as a slot. Attached to the revolute joint are the arm and all necessary driving mechanisms and actuators. The two joints move in coordination to produce the desired motion and trajectory of the arm tip. E x a m p l e 1: A n a l y t i c compar i son In order to assess the accuracy and reliability of the above described solution procedure, a comparative study was first made analytically. The required modal data used in the solution w a s 682 identified, using a linearization-based methodology for modal analysis[10], from the timeresponse numerically obtained through a finite element solution. As a result of processing this modal extraction, there were 10 identified discrete sets of modal parameters available in this example. Figure 3 shows the time history of the lateral displacement at the tip of the arm (see Fig. 2) due to the rigid-body inertia arising from imposed trapezoidal velocity profiles for both the revolute and prismatic joints. The revolute joint was assumed to rotate 180 \u00b0 and the prismatic joint was assumed to simultaneously slide the arm from an initial length of 9.9in. (251.5 mm) to a final length of 8.1 in (205.7 mm). The acceleration and deceleration durations were set to a thousandth of the total duration of motion (0.5 s). In the figure, the dashed curve denotes the solution resulting from this mode-based procedure which utilizes the analytically extracted modal models, in comparison with the original time-history solution (solid curve)", + " The specified motion at the revolute and prismatic joints has simple harmonic motion profiles as follows: O(t)=~ - ~ c o s 0 ~ < t ~ < 0 . 5 s(t) = 1.6 -- ~cos ~ 0 ~< t ~< 0.5 (7) Dynamic response prediction 685 The modal parameters for the three lowest modes utilized in the mode-based solution were experimentally identified [14] through modal testing. For illustration, the mode frequency and mode shape for the first mode are shown in Fig. 6 where the mode shape is normalized such that the mode amplitude at station 1 (see Fig. 2) is equal to unity. Applying the solution procedures outlined in Fig. 1, the transverse deflection due to the specified motion results, as shown in Fig. 7. These curves are considered to be elastodynamically-induced errors of the manipulator under investigation, i.e. the unwanted deviation of the robot 's position (or trajectory) because of its flexibility. Example 3: Continuous multi-straight-line trajectory Figure 8 depicts a top view of the RP manipulator (dashed), with a prescribed multi-straight-line continuous path to be followed" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000577_aim.1999.803236-Figure2-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000577_aim.1999.803236-Figure2-1.png", + "caption": "Figure 2: Schematic diagram of the shaft dynamometer for the human-powered submarine. Shown are the propeller driveshaft and the magnetic encoders, including typical output waveforms for the Hall Effect sensors.", + "texts": [ + ", Matlab, Simulink, Mathematica, Maple, and AutoCAD). Shaft Dvnamometer for a Human-Powered Submarine. As part of the human-powered submarine design competition, in which the ME Department has competed for several years, a need exists to monitor in real time the speed, torque, power, and total energy associated with the vehicle propulsion shaft. One mechatronics project team developed a dynamometer, based on a pair of custom-designed magnetic encoders at opposite ends of the propeller driveshaft. As shown schematically in Fig. 2, a pulse timing technique was used to infer both angular velocity and shaft torque (based on shaft twist) sixteen times per revolution of the shaft. From these measurements, the other output variables of interest were computed. To converge on an optimal design, the students used a model relating the mechanical and computational design parameters to the expected accuracy and bandwidth of the dynamometer. The final design included a scheme that used an automatic lookup table to compensate for the manufacturing tolerances in the magnetic encoders" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002726_iros.1994.407422-Figure5-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002726_iros.1994.407422-Figure5-1.png", + "caption": "Figure 5: The motion generation scheme.", + "texts": [ + " Let 6t be the time increment of the motion generation process (St = AT/p), s\u2019 be the state of A corresponding to the configuration q\u2019, and let s\u2019(n6t) be the state of A obtained after having applied n elementary motion steps from sa (i.e. after having applied a sequence of n controls on the \u201cphysical effectors\u201d of @(A)). Determining the required sequence of controls to apply to A can be done by executing an iterative algorithm involving two complementary steps. The,first step consists in hypothesizing a nominal sub-path r; between the current sub-configuration @(nbt) and the next sub-goal represented by geXt, and the second step allow to track I\u2019k while processing the physical vehicle/terrain interactions, as illustrated in figure 5. In the current implementation of the algorithm, ri is constructed using a technique derived from the Dubins\u2019 approach [6]. The obtained sub-path is a smooth curve made of straight line segments and circular arcs (the choice of the gyration radius to be applied at a given step of the algorithm depends on the length of the involved path4 and on the mechanical characteristics of A, see [5]). The trackin function operates on the physical models @(A) and a(%). It takes as input the velocity controls applied on each controlled wheel during a time increment S t ", + "Jj corresponding to the wheels). Since the motion generation step accounts physical phenomena like sliding or skidding, the state s:(nSt) of A obtained after having applied n successive controls may be different from the nominal state si(n6t). The processed motion generation step will be considered as a failure when the corresponding sub-configuration c( nSt) is too far from its nominal value qi(n6t). The previous algorithm is iterated until the neighbourhood of sLeZt is reached or until a failure is detected (see figure 5). Figure 6 shows a local trajectory obtained when A is controlled to cross an irregular area of the terrain. *The control U is given by the linear velocity and the gyration angle of the robot (u,q5) E, { - V O , V O } x {-q5mal,0,~ma2} where q5maz is the maximum steering angle of A (as described in [l]). 31n this case, 2r = 0 and q5 E {-q5mQ2,q5mQ5}. 41f this length is large enough, the velocities of the controlled wheels are positive; otherwise, the velocities of the opposite controlled wheels have opposite signs and the vehicle will skid and use more energy to execute the motion" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002912_s0093-6413(02)00249-5-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002912_s0093-6413(02)00249-5-Figure1-1.png", + "caption": "Fig. 1. Coordinate system and Lewis parabola.", + "texts": [ + " For these reasons, usage of j in calculation of J is a direct, easy to implement and flexible method without needing any numerical algorithms. Profile of an involute spur gear tooth is composed of two curves. The working portion is involute and the fillet portion is trochoid. Coordinates of the points on working and fillet portions can be determined by using parametric equations previously derived by the author of this paper (Arikan, 2000; Arikan, 1995). Used coordinate system, important points and tooth profile are shown in Fig. 1. Theoretical limit radius (rL) is an important parameter when gear kinematics is considered. It is the radius at which the involute profile on a gear should start, in order to make use of the full length of the involute profile of the mating gear. If it is smaller than the radius at which involute tooth profile starts (rti), involute interference occurs. When gear teeth are produced by a generating process like using a rack cutter or hobbing, interference is automatically eliminated, since the cutting tool removes the interfering portion", + " B \u00bc B m\u00f0NP \u00fe NG\u00de cos/m cos/c \u00f06\u00de Involute function, inv is defined as follows: inv/ \u00bc tan/ / \u00f07\u00de When the gear pair is going to be operated at an exact desired center distance Cm, /m can be found by making use of the following equation: cos/m \u00bc Cs cos/c Cm \u00f08\u00de where, Cm is the operating, Cs is the standard center distance, and can be calculated as given below: Cs \u00bc m\u00f0NP \u00fe NG\u00de 2 \u00f09\u00de For spur gears, definition of the AGMA geometry factor J reduces to the following equation (AGMA Standard 218.01, 1982; AGMA Standard 2001-B88, 1988): J \u00bc Y Kf \u00f010\u00de Following reduced equations can be used for tooth form factor Y and stress correction factor Kf . Tooth dimensions used in calculation of Y are shown in Fig. 1. /L is the load angle. Y \u00bc 1 cos/L cos/m 6hF s2F tan/L sF \u00f011\u00de Kf \u00bc H \u00fe sF qF L sF hF M \u00f012\u00de where, H \u00bc 0:331 0:436/c \u00f013\u00de L \u00bc 0:324 0:492/c \u00f014\u00de M \u00bc 0:261\u00fe 0:545/c \u00f015\u00de Height of the Lewis parabola hF, and corresponding tooth thickness at critical section sF can be calculated by finding the coordinates of the tangency point of the Lewis parabola and the tooth root fillet curve. As seen in Fig. 1, y-coordinate of the vertex of the parabola yL can be found by making use of the equation of the line whose slope is tan/L and passing through a point on the involute tooth profile \u00f0xi; yi\u00de. Values of xi and yi can be found by using the parametric involute equations: yL \u00bc yi xi tan/L \u00f016\u00de Following equations can be used to find the load angles for the loads acting to the tooth at diameters dx and Dx, which are for the pinion and the gear, respectively. Load angles are measured from an axis perpendicular to the tooth centerline" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003674_j.mechmachtheory.2004.05.002-Figure6-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003674_j.mechmachtheory.2004.05.002-Figure6-1.png", + "caption": "Fig. 6. Rigid body displacement associated with a prismatic joint.", + "texts": [ + " Moreover, according to the shown procedure, it can be concluded that $j is nothing but the screwen $i after an infinitesimal change of the independent parameter of motion. Thus, and analyzing the terms involved in Eqs. (30) and (37), we can finally state that o$i o/ \u00bc $ $i\u00bd \u00f038\u00de thereby proving that the screw derivative of an arbitrary screwen, when the screwer is associated with a revolute joint, can be expressed in terms of the Lie product. This particular case of rigid body displacement occurs when the screw axis coincides with the motion axis of a prismatic joint. A geometrical description of such a displacement is shown in Fig. 6. According to the nature of the case under study, we have now that the screws associated with the corresponding motion axis, the reference configuration and the current configuration are, respectively, as follows: $ \u00bc 0 e ; $i \u00bc ei ri ei \u00fe pei ; $j \u00bc ej rj ej \u00fe pej \u00f039\u00de Then, resorting to the definition of the Lie product, Eq. (5), and for the case under study, we obtain $ $i\u00bd \u00bc 0 e ei \u00f040\u00de On the other hand, considering the geometry shown in Fig. 6, we can observe that the translational vector parallel to the screw axis is now dk \u00bc se, being now s the independent parameter of motion. Then, we have that ej \u00bc 1ei \u00bc ei \u00f041\u00de rj ej \u00bc ri ei \u00fe s\u00f0e ei\u00de \u00f042\u00de In analogy with the procedure presented in Section 4.1, the corresponding partial derivatives of the above expressions take now the following form: oej os \u00bc 0 \u00f043\u00de o\u00f0rj ej\u00de os \u00bc e ei \u00f044\u00de Thus, the partial derivative of screw $j can be arranged as follows: o$j os oej=os o\u00f0rj ej\u00de=os \u00bc 0 e ei \u00f045\u00de Analyzing the foregoing expression, it can be noted that Eq" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003261_2001-gt-0255-Figure7-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003261_2001-gt-0255-Figure7-1.png", + "caption": "Figure 7: Photograph of test rig", + "texts": [ + " changing the input angle of inclination of the rotor (\u03b8x) resulted in different solutions for the higher speeds. Figure 6 shows the run up solution at 18,000rpm compared to a solution obtained by using an all zero initial condition. Here, the zero initial condition solution is chaotic. These multiple solutions Copyright \u00a9 2001 by ASME ms of Use: http://www.asme.org/about-asme/terms-of-use Down result from the bearing non-linearity and should be avoided, as far as possible, by appropriate selection of system parameters. Figure 7 shows a photograph of the modeled test rig which is currently being commissioned. The test rig model is the same as that in [10] except three more DOF were added at each bearing as required by the 5DOF modeling. Table 3 gives the relevant parameters for the rotor system while the lumped mass model and DOF are shown in Fig. 8. Note that DOF 25 and 26 are added axial DOF that have no stiffness, but have masses of 0.1kg. The rotational DOF at masses 2 and 3 have Id values of 7.130\u00d710-3kgm2 and 3.995\u00d710-3 kgm2 respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001637_0146-6380(85)90004-x-Figure2-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001637_0146-6380(85)90004-x-Figure2-1.png", + "caption": "Fig. 2. Schematic of the analytical train showing that used for thermal reduction analysis (upper), and aqueous scrubbing analysis (lower).", + "texts": [ + " Other weight loss data obtained were for acidified samples originally intended for CHN determinations. While ambiguous due to the calcareous nature of the CB sediments, these data and data for acidified CCS sediments are included for comparison with other parameters. Replacement of COl by CI~- should result in little change in ash weight of the CB sediment, The CCS sediments were not calcareous sands and the weight loss data for acidified CCS sediments are probably equivalent to the organic matter content. The analytical apparatus is shown schematically in Fig. 2. For thermal reduction analyses, 3-100mg quantities of pulverized, freeze-dried sediment were weighed and placed inside a small gold foil boat. Inside a 10 mm o.d. quartz tube wrapped with nich- rome heating wire, and under a carrier flow of 200 ml/min (H2:He, 2:3 v/v), the sample was heated to 480\u00b0C for 10min by applying 15V through a rheostat. The evolved sulfur gases were swept into a liquid N2-cooled U-trap (6mm o.d. Pyrex, 35 cm long) packed evenly with 60/80 mesh Chromosorb 102 for collection", + " The PM signal was amplified by a Heath Photometric Readout (EU-701-31) operated at 10 -6 or 10-7A and recorded on a Linear Instruments model 252 integrating recorder operated at 3000 cts/min and 1 in/min. A typical thermal reduction chromatogram is shown in Fig. 3. All thermal reduction data are presented as means of triplicate determinations in which precision averaged _+ 20% with variability due to the small size and heterogeneity of the samples. Aqueous scrubbing analysis was also performed for each sampling interval. Here, measured volumes of wet sediment from polyethylene bottles were taken by syringe corer and transferred to a bubbler/scrubber apparatus (Fig. 2) containing 50 mi of degassed, deionized water. The scrubber was connected to the carrier line, purged of 02 for one minute, and connected to the sample collection line. The sediment plug was dispersed by activating a magnetic stirrer and liquid N2 trapping began. Five hundred #1 of acidic Cr(II) solution was then injected into the slurry through the scrubber side arm and scrubbing continued for 30 min. The Cr(II) solution was prepared in a Jones reductor apparatus. A 30 min scrub time was experimentally found to give quantitative recovery of MESH" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002898_bf02441311-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002898_bf02441311-Figure1-1.png", + "caption": "Fig. 1 E02 s Principle o f pCO2-electrode with a monocrystalline antimony electrode as pH sensing element", + "texts": [ + " First received 17th June and in revised form 12th October 1980 0140~)118/81/040447 + 10 SO1-50/0 9 IFMBE: 1981 The antimony electrodes were spark-cut from a large single crystal in a fixed direction to obtain monocrystalline electrodes with the measuring surface parallel to the trigonal (1i0) crystallographic plane. The electrodes were cast in epoxy resin (Araldite H/hardener HY951, Ciba-Geigy) to form plastic cylinders fitting into the conventional Radiometer a = inner electrolyte b = reference electrode c = salt bridge d = spacer e = membrane f = electrode housing g = Araldite socket h = monocrystalline antimony i = O-ring Medical & Biological Engineering & Computing July 1981 447 p C O 2 jacket (Model E5037), with the measuring surface exposed at the bottom end, see Fig. 1. This surface (A =0\"8mm z) was polished to optical smoothness using 1 pm diamond paste as the final step. Antimony crystals of two different purity grades were used; pure antimony (Materials Research Corp.) and highly purified antimony (Studsvik Energiteknik AB). The purity grades were Sb > 0\"9999 (4N) and Sb > 0.999999 (6N), respectively. The highly purified antimony was obtained by further zone refining of the pure material. The design of the electrode is shown in Fig. 1. Carbon dioxide from the test sample diffuses through a 25pm silicone membrane (Type D606, Radiometer AS) into the inner electrolyte. A nylon mesh, 50 #m thick, was used as a spacer in front of the electrode to keep the electrolyte film at a constant thickness. The corresponding volume of the electrolyte film was approximately 1#1, whereas the bulk electrolyte volume was approximately 0\"5ml. An Ag/AgC1 reference electrode was used in both configurations shown in Fig. 1; as a coaxial electrode inside the measuring cell, or connected to the measuring cell with a salt bridge. The internal coaxial reference electrode was a chloridised silver foil, whereas the external reference electrode was a commercially available Ag/AgC1/(3 M KC1) half cell (Model 373, Ingold). The complete pCOz-electrode was installed in a bloodgas analyser (Radiometer BMS 3) thermostated at 37~ Potentiometric measurements were made with a digital electrometer with high input impedance (1015f~)", + " Thereby a potential that is stable and reproducible for a sufficiently long period of time for several applications can be gained. Pits, cracks and other inhomogeneities in the surface have to be avoided, since they will give rise to local corrosion cells. These cells result in rapid and irreproducible changes in current densities and cause potential level instability. W h e n used as a pH sensor in a Stow-pCO2 configuration, the antimony electrode is immersed in the inner electrolyte of the pCO2-electrode consisting of NaHCO3 and NaC1 (Fig. 1). The carbon dioxide diffuses from the sample through the membrane and causes a change in the pH of the electrolyte according to C O 2 + H 2 0 ~ HzCO 3 ~ H + + H CO 3 so that K \" [ H C 0 3 ] pH = p ,+ log - - - - - . . (4) Sco2\" pCO2 K. = first dissociation constant of H2CO 3 S = solubility coefficient According to Fig. 2 the antimony potential is under mixed control 9 Nevertheless a change in pH at constant pO2 results in a linear potential response. A change of the equilibrium potentials caused by a change in pH will, however, not necessarily result in an equal change of corrosion potential" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003742_bf02324896-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003742_bf02324896-Figure1-1.png", + "caption": "Fig. 1--Types of Wildhaber-Novikov gear", + "texts": [ + " a = p r e s s u r e ang le , deg pl = prof i le r a d i u s of t h e p i n i o n too th , m m p~ = prof i le r a d i u s of t h e w h e e l too th , m m Zxp = d i f f e r e n c e b e t w e e n t h e p ro f i l e rad i i , m m ac = f l ank s t ress , P a r -\" b e n d i n g s t ress , P a Introduction W i l d h a b e r - N o v i k o v g e a r s h a v e c o n v e x a n d c o n c a v e c i r c u l a r - a r c prof i les o n t h e f l anks of t h e m a t i n g t e e t h of t h e p i n i o n a n d w h e e l gear . T h e c o m b i n a t i o n 116 t March 1976 of these circular arcs chosen for the mat ing teeth defines the type of Wi ldhaber -Novikov gear, namely, al l -addendum, a l l -dedendum and addendum-dedendum (Fig. 1). These gears have a contact surface larger than that of involute gears owing to the high conformity of the mat ing profiles. Though Wi ldhaber -Novikov gears are not in great commercial use as yet, they are used for power transmission in marine engines, a i rcraf t engines and agricul tural equipment, where heavy loads are to be transmitted. 1 Reluctance to use these gears on a wider scale is due to the higher accuracy needed for the manufacture of these gears and due to the reported unsui tabi l i ty of these gears for h igh-speed applications" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001103_s0736-5845(97)00033-1-Figure7-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001103_s0736-5845(97)00033-1-Figure7-1.png", + "caption": "Fig. 7. Damped least-squares approximation to an unreachable command with the Kraft slave manipulator.", + "texts": [ + " In the particular control system described however, they are not controlled using a joint space mapping between the master and slave. Instead, a task space mapping, with scaling and indexing, is used. Weighted damped least squares is used to deal with unreachable commands and operation near singularities, in real time within the local controller of the slave. The remainder of this section is a discussion of the how this system performs, giving a specific example of the approximate solutions obtained outside a workspace boundary defined by joint limits Fig. 7 uses the kinematic model of the Kraft slave to illustrate the convergence of a resolved rate controller using a damped least squares solution to Eq. (2). The command, c1, and the result, r1, are on the boundary of the workspace. Joint 3 is at a limit and cannot move in the required direction to satisfy the command, c2. To move from c1 to c2, requires that joints 2 and 4 rotate clockwise, while joint 3 rotates counter clockwise. The result, r2, of the command c2, is not intuitive to the operator", + " If the command is extended further out from the base, the manipulator has a tendency to move further down, creating confusion for the operator. While the damped least-squares solution may perform satisfactorily for bondaries defined by singularities, it is clearly unacceptable for boundaries defined by joint limits. Fig. 8 uses the kinematic model of the Kraft slave to illustrate the convergence of a resolved rate controller using the weighted damped least squares methods proposed. The result of extending the command, c1 in Fig. 7, to c2 in both figures, is r2 in Fig. 8. The resulting solution \u2018\u2018points\u2019\u2019 to the commanded position. As the command is moved through the unreachable space, as shown by c3 and c4, the robot tracks along the workspace boundary, finding approximate solutions in the task space. For a human operator, the result r2 in Fig. 8 is a much more intuitive solution to the command c2 in Figs. 7 and 8 than is the result r2 in Fig. 7. Furthermore, with a bilateral controller where the force-feedback in the master tends to drive the master to the slave, the operator is further confused by the result in Fig. 7. Here the master will be driven downward rather than along a shorter path back to the workspace of slave, as results from the weighted solution. It has been determined by all of the operators of the teleoperational system, that the controller using the weighted damped least squares methods is a very significant improvement over the re-alignment and re-indexing solutions to unreachable commands. Also, the damping of the resulting joint velocities makes it possible to operate within the workspace near a singularity" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000935_elan.1140090215-Figure2-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000935_elan.1140090215-Figure2-1.png", + "caption": "Fig. 2. A schematic diagram of apparatuses for depositing the LB film with adsorbed GOD.", + "texts": [ + ", Tokyo, Japan) to improve adhesion of the gold film to the glass substrate, and a l5Onm gold film was deposited on the chromium film using an ULVAC evaporating system (EVC-SOOA, Tokyo, Japan). A contact metal mask was used Elecfroanalysis 1997, 9, No. 2 0 VCH Vedagsgesellschafl mbH, 0-69469 Weinheim, 1997 1040-0397/97/0202-161 $ 10.00+.2S/0 162 H. Tsuji, K. Mitsubayashi 2.2. Deposition of LB Film LB films with adsorbed glucose oxidase (GOD) were deposited on the substrate using apparatuses shown in Figure 2. A Lauda film balance (Langmuir trough), Model FW-1, and a Lauda film lift (vertical film dipping apparatus), FL-1, were used to deposit LB films. Amphiphilic compounds with plus, minus or neutral charge at a head group (see Scheme 1) were used to form LB films. Arachidic acid (AA), stearylamine (SA) and dimethyl-dioctadecylammonium bromide (DDAB) were purchased form Wako Chemical Co. (Osaka, Japan), Sigma, and Sogo Pharmaceutical Co. (Tokyo, Japan), respectively. L-Glutamic acid dioctadecyl ester (GDE) was synthesized by Katayama Chemical Co" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001017_0021-8928(96)00031-7-Figure6-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001017_0021-8928(96)00031-7-Figure6-1.png", + "caption": "Fig. 6. Fig. 7.", + "texts": [ + " The control ~tl(X) has the form (4.3) (the notation 0 has to be replaced by 7). The control lX2(~) has the form (xl = 0.21776, x2 = 0.44238) I . t 2 (X) = - 1 when 0 ~< x ~< xl, ~t2(x) = 1 when xl < x ~< x 2 la2(x) = 0 when z 2 < x ~< T - X 2 P-2(X) = - i whenT - x2 < x ~< T - \"c I, tx2(x ) = 1 when T - xl < x ~< T (4.7) It is shown schematically in Fig. 5. Just when x2 ~< x ~< T - I; 2 the links are folded and the moment of inertia of the two-link mechanism with respect to the hinge O is minimal. However, the motion shown in Fig. 6 (\u00a51 -- 1, ~3(0/2) = 0.08873), which satisfies the maximum principle, takes less time in the case of conditions (4.6). In the case of such a motion, the link KL completes oscillations around the link OK, that is, the angle x3 does not decrease strictly monotonically (x4(0/2) > 0). The time of this apparently optimal motion 0 = 0.96830. It is less, although insignificantly, than the time T = 0.98656 of the motion which contains the singular mode, the gain in time being just 2.6 x 10 -4. The control ~1('17) has the form of (4" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001827_s0263574700004641-Figure2-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001827_s0263574700004641-Figure2-1.png", + "caption": "Fig. 2. Collision-free trajectory for the spherical manipulator.", + "texts": [ + " Therefore, the problem reduces to finding the shortest path from vertex x0 to vertex xg, consisting of arcs of minimal loading. This task is classical and solvable by many known algorithms.'\u20226912~15 On solving the task one obtains an optimal one-step trajectory xs equal to sequence of states f+((x0, \u00ab,): / = 0, 1, . . . , n - 1) produced by a sequential use of input letters from the optimal sequence un. An example of optimal sequence of states and corresponding to them unequivocally input letters is given in Figure 2. The figure illustrates the raster, set Y and state space X for a 2RIT manipulator, as well as objects of the scene placed in the raster. As far as the global function is concerned it is assumed that G(R) = G(L) = 3, G{U) = G(D) = 2, G(F) = G(B) = 1, G(N) = 0. The initial state has been chosen as x0 = (i, j , k), and the terminal state as xg = (i + 4, j + 3, k + 3). Optimal trajectory takes the form xs = ((i, j , k), j) ( 3, j + 2, k), j) 1), (i1 ) 2,k + 2), ( j ) ( j ) ( j 3)). To such a trajectory there corresponds a sequence uw = R, R, R, U, U, F, F, U, F, R", + " Function P determines a subsequence x*tzxs whose elements must be reached successively by the manipulator. The transition between them does not have to be realized by means of the segments of sequence xs having been removed, as the transition obtained by simultaneously performing the movement corresponding to them does not lead to a collision with objects on the scene. Obviously, the space trajectory between the states is not determined precisely and the only thing being certain is that the trajectory lies entirely in appropriate elements of output set Y' (see 3.1). Figure 2 shows an example of collision-free trajectory x* for the optimal sequence of states xs chosen at the first step of planning. The function P has selected the following set as a sequence x* = ((i,j, k), (i + 3, j + 3, k + 2), (i + 4, j + 3, k + 3)). This shows that the trajectory found in the first step consists of ten successive movements of the manipulator, while the collision-free trajectory x* consists of two movements corresponding to the simultaneous performance firstly the sequence u, = R , R , R , U, U, F, F t h e n , a f t e r t h e s t a t e x 7 = (i + 3,j + 3, Collision-free movements 293 k + 2) has been reached, a simultaneous sequence u, = U, F, R" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002872_1.1456456-Figure8-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002872_1.1456456-Figure8-1.png", + "caption": "Fig. 8 Ball bearing parameters", + "texts": [ + " (12) Once the radial stress p is found, the radial displacement of the bearing outer race can be found by d[u1(a) through ~1! to ~4!. Finally, the radial displacement d can be used to calculate the contact angle and the bearing stiffness at elevated temperature through the following three steps. In this section, quantities with primes refer to quantities at elevated temperature. First, the radial clearance Pd8 is determined by Pd85do82di822D8, (13) where the inner race diameter di8 and ball diameter D8 ~Fig. 8! are predicted through the coefficient of thermal expansion and temperature change. Moreover, the outer race diameter before and after the temperature change is related by do85do12d . (14) Second, the new contact angle a8 can be estimated through a simple kinematic model as shown in Fig. 9. As a result of the temperature change, the outer race expands with a displacement d8 relative to the inner race, where 798 \u00d5 Vol. 124, OCTOBER 2002 rom: http://tribology.asmedigitalcollection.asme.org/ on 08/23/2017 Term d85 1 2@~do82do" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000175_0016-0032(95)00056-9-Figure5-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000175_0016-0032(95)00056-9-Figure5-1.png", + "caption": "FIG 5.", + "texts": [ + " The mode of numbering equations, although unorthodox, is adopted so as to clarify appropriate groupings and, thereby, to facilitate the reader's appreciation of the governing equations of linkages treated. In this work, as in many other places, we abbreviate sine to s, cosine to c and tangent to t. It is difficult to visualise any of the octahedra of Bricard without the aid of a model or a graphic depiction, such as Fig. 11 of (1), the essence of which is reproduced here as Fig. 4. In order to illustrate our notation, however, we must divest the portrayal of physical interpretation and employ Fig. 5, which has the appearance of a convex polyhedron. It allows us, as well, the convenience of referring to the \"outside\" of the representation, as we shall have cause to do. In assigning vertical angles to the octahedron for either of the symmetric cases, our decisions are facilitated by the obvious nature of the property, and eight notionally independent angles are readily specified (24, 25). The matter is not straightforward for the remaining type which is definable, broadly, by its peculiar characteristic of Journal of the Franklin Institute 658 Elsevier Science Ltd Bricard's Doubly Collapsible Octahedron Vol. 332B, No. 6, pp. 657-679, 1995 Printed in Great Britain. All rights reserved 659 two collapsed configurations; then, as stated by Bricard, the opposite vertical angles at each node must be equal or supplementary. More precisely, as observable from a working model, the property of supplementarity applies to each of two opposed nodes and that of equality to the other four. In Fig. 5 we show how such an arrangement might be achieved. There are again the eight nominally defining angles (~1, (~2, i l l , /~2, ~1, \u20222, 1~1, 62 and we put, as in (24, 25), 0~ = 0~ 1 -}-~2 7 = 71 +?2 = 01 \"1- \u20222 ' A primed symbol indicates the supplement of the angle denoted by the unprimed letter. By introducing the constraints 7=c~ 6= /~ , (B-i) we reduce the number of independent angles to six; a further reduction will be seen to follow from assembly of the octahedron. It is evident that vertices B and E are the two which exhibit the property of supplementarity between opposite pairs of vertical angles", + " We have yet to consider, in subsequent sections, the implication of the octahedron's mobility. The procedure about to be described applies to all faces of our octahedron. Figure 6 represents the plate CAB, and there is indicated a clockwise direct ion o f c i rcula t ion as viewed from the topological outside of the figure. A dot at a vertex means that the notional link between adjacent hinges there is directed \"out of the plane\" and a cross stands for a sense \"into the plane\"\u2022 (It may be observed that there is just one primed symbol for each face shown in Fig. 5; the vertex so identified will be the one to carry the cross.) Hence, under the usual right-hand convention, for the plate depicted, the three skew angles involved are cq, c~z and a = 7z + a2. We now imagine that the face is identified with a diametral plane of a notional sphere (Fig. 7), the hinge-axes rendered by vectors emanating from its centre. The heads of the vectors then define a two-part arc which can be called the spherical indicatrix (13) of the plate. If all plates are treated in the same fashion on one notional sphere, the \"arc-form\" linkage (25) of Fig", + " Strangely, given Bennett's familiarity with \"paradoxical\" linkages and techniques of synthesis, he makes only passing mention of the issue of the six-bar forms contained in the networks discussed in this paper. VI. The Octahedral Six-bar Linkage and a Generalisation We turn at last to the six-bar loops incorporated in the octahedron itself. There are twelve hybrid chains of the double-Hooke's-joint type obtainable from neighbouring pairs of spherical four-bars; they call for no attention from us. There are, as well, four non-hybrid loops, each realisable by removing a pair of symmetrically opposed plates from the polyhedron and traversing the remaining hinges; they can be defined, with reference to Fig. 5, by ABCDEF, ABFDEC, ACBDFE, AECDBF. Closure equations for the non-hybrid octahedral linkages of all three kinds are developed in (12), but are based in part there on standard sets of displacement relationships which are difficult to manipulate, and so some of the resulting equations are rather unwieldy. The recent advances made (24, 25) in treating the symmetric octahedra allow for a more direct determination of angular relationships, whence the governing equations of the relevant six-bars can be expressed in simpler formats, as in (26)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002959_s0021-9290(00)00216-5-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002959_s0021-9290(00)00216-5-Figure1-1.png", + "caption": "Fig. 1. (A) Walking model. Our model consists of eight massive rigid links, which represent upper and lower torsos, thighs, shanks, and feet, respectively. The whole body is supported by either one or two legs at the same time depending on the supporting condition. (B) Main symbols. Link number 0 shows the ground being in contact with the foot and the index letter i counts up to the upper links. Values of physical parameters of the walking model were obtained in our previous paper (Tagawa and Yamashita, 1983).", + "texts": [], + "surrounding_texts": [ + "Simulating abnormal walking is difficult because quantitative characteristics of human abnormal walking have not been collected enough to be described as dynamic models. Our method of the study and mathematical equations are described in concise forms in the following. 2.1. Basic walking model and equations The link chain and main symbols of our model are shown in Figs. 1A and B, respectively. Adjacent links are coupled by a frictionless joint. The whole body is allowed to move in three-dimestional (3D) space. The motions are controlled by the moments Mi \u00bc MXi MYi MZi\u00bd t applied at the ith joint except the ZMJ. All components of the moment at the ZMJ are always kept to be zero. The procedure to simulate the human normal and abnormal walking is as follows: (A) Reliable results obtained experimentally in normal walking, defined usually as average human characteristics (Bresler and Frankel, 1950; Cunningham, 1958; Eberhart et al., 1968; Murray et al., 1964; Murray, 1967; Saunders, et al., 1953; Yamashita and Katoh, 1976), are incorporated into the model to mimic human walking. These Nomenclature English characters Fi 3D vector of linear force acting on ith link at ith joint F* i 3D vector of net linear force acting on ith link F\u0302i 6D spatial vector of force acting on ith link at ith joint; consists of linear force and moment (\u00bc Fti M t i t ) F\u0302 * i 6D spatial vector of net force acting on ith link; consists of force and moment at posterior and anterior joints where gravitatinal force is omitted by the assumption of imaginary acceleration g (\u00bc F\u0302i F\u0302i 1 \u00bc F*t i M* t i h it ) g gravitational constant Ji ith joint l leg length Mi 3D vector of moment acting on i th link at ith joint, (\u00bc \u00bdMXi MYi MZi t), M0 \u00bc 0 at the ZMP M* i 3D vector of net moment acting onith link mi mass of ith link O X0Y0Z0 inertial coordinate system: X0 in the direction of walking, Z0 in the direction of gravity (positive downward), and Y0in the direction so that the three axes X0Y0Z0 form a right-handed system. O XiYiZi coordinate system fixed in ith link. Rotations of the ith link describe any orientation in terms of a rotation ai about X-axis, then a rotation bi about a new Y-axis, and finally a rotation gi about a new Z-axis. The final rotated coordinate system is fixed in the link. pi 3D vector of position of ith joint SX step length T walking cycle time t time TD double-support period which is a percentage of walking cycle time V walking speed _XE forward velocity of center of gravity of model YE lateral displacement of center of gravity of model Greek characters ei 3D position vector from (i 1)th joint to center of gravity of ith link ni 3D position vector from (i 1)th joint to ith joint Subscripts E center of gravity of walking model i link or joint number l left leg r right leg X ; Y ; Z component in inertial coordinate sys- tem O X0Y0Z0 general symbol for the subscripts i and X ; Y ; Z Superscript t transposed form Unit - nondimensional expression - Length quantities are normalized by leg length l. Inertial terms are divided by l2 P mi. Nondimensional time t ffiffiffiffiffiffi g=l p is introduced. Notation \u20188\u2019 means differential with respect to nondimensional time. consist of motions of the lower torso and of the ZMP in the normal leg; a foot angle relative to the ground during the stance phase; a ratio to distribute floor reaction forces to respective feet during the double-support phase. (B) Angular displacements at the knee and hip joints during the stance phase are calculated kinematically from the predetermined motions in (A). (C) Angular displacements at the leg joints during the swing phase are determined from the linearized equation of motion of the leg, to satisfy the conditions of continuity of the angular displacement and velocity at the toe-off and the heel contact. (D) The equation of motion of the upper torso is obtained by the dynamic balance at the ZMJ, then solved to satisfy a periodic condition. (E) By comparing the simulation and the experimental results for normal walking at some speeds, agreements between them are examined. When the agreement of some variables is insufficient, step (A) is adjusted to improve the disagreement. The force, moment, and ZMP are briefly explained in the following to help in understanding the physical meanings of mathematical expressions: (1) Force and moment The motions of links below the lower torso are completely determined from steps (A) through (C). That is, F\u0302* i (i\u00bc 1; . . . ; 4) can be expressed quantitatively. F\u03020 \u00bc Ft0 M t 0 t is zero during the swing phase. Unknown quantities are F\u03024 and F\u03020 during the stance phase. At the ZMJ (assumed at joint j), the force Fj is transmitted to the adjacent links while the moment Mj is zero. (2) Zero moment point In the abnormal walking having the ZMJ, there are two positions in the diseased leg to satisfy the zero moment condition, at the ZMJ and the ZMP, because the motions of lower extremities are assumed to be the same between the abnormal and normal walking. The following three equations are derived in a case of the ZMJ at the jth joint of the right leg. The position vector of the ZMJ: pj \u00bc p0 \u00fe n1 \u00fe Xj i\u00bc2 ni; \u00f01\u00de where the unknown vectors are p0 and n1. The constrained equation of the moment at the ZMJ and the dynamic equation of the upper torso are expressed, respectively, as Xj i\u00bc1 M* i pj p0 Fj \u00bcMj \u00bc 0; \u00f02\u00de F\u0302 * 5 \u00bc F\u03024 \u00bc F0l Fjr p0l F0l pjr Fjr Mjr \" # X4 i\u00bcj\u00fe1 F\u0302 * i X3 i\u00bc1 F\u0302 * il \u00f03\u00de When an equation obviously refers to the right leg system, the subscript r is omitted for simplicity. Each term appearing in P in Eq. (3) is a net force acting on the link that is determined by steps (A) and (B). The unknown quantity F\u0302 * 5 , which is the net force acting on the upper torso and consists of inertial force and couple, is expressed as a function of the variable representing the motion of the upper torso. The unknown quantities are obtained by the following four steps: (a) to solve the equation of the motion of the upper torso to satisfy a periodic condition for one walking cycle starting from the heel contact of the right leg, (b) to calculate the Fj from Eq. (3), (c) to calculate the p0 from Eq. (2), (d) to calculate the n1 from Eq. (1)." + ] + }, + { + "image_filename": "designv11_65_0000968_acc.1997.611036-Figure4-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000968_acc.1997.611036-Figure4-1.png", + "caption": "Figure 4: Inverted pendulum model scheme.", + "texts": [ + " A physical IP has served as a standard laboratory example for the SIMPLEX architecture described in the introduction. Here we present results of the neural network estimation of the stability region based on data from a simulation model using both supervised and reinforcement learning. The IP model equations are: 1 1 2 2 J tx+-mlcosa&+B, .x - -mlsinatu2 = F 1 1 . 1 -mcosax+-mld!--mgsina = 0 2 3 2 with the following constraints: IF1 I Fmaz , 1x1 I xmaz , 121 I umaz The parameters take the values: Jt 0.6650 Kg g 9.8 m/s2 m 0.21 Kg Fmax 2 N 1 0.61 m xma, 0.45 m B, 0.1 Kg/s vmax 1 m/s Figure 4 shows a model scheme for a standard IP. Su- pervised and reinforcement learning is applied. Table 3 summarized some parameters calculated for the network and some results obtained on the simulator. Figures 5 and 6 show some cuts of the region found for the model compared to the best linear region based on constraints for individual states variables, calculated by maximizing the volume of the hypercube defined by the constraints. solid: neural network estimation. dashed: best linear constraints. solid: neural network estimation" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003375_icmech.2004.1364455-Figure2-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003375_icmech.2004.1364455-Figure2-1.png", + "caption": "Fig. 2. Energy optima! controlled motion.", + "texts": [ + " EXAMPLES The approach has been applied to a variety of nonredundant and redundant planar and spatial SMs. For the sake of clarity a planar manipulator kconsidered here. Results for cubic splines are reported in order to elucidate the effect of non-smooth jerks. The NLP is solvcd with a standard package for constrained o p timization [S],[Zl]. No effort was yet made on advanced and tailorcd optimization methods, that shall be expected to reduce the necessary computation time for solving the NLP. Figure 2 depicts the top view of a redundant planar 5R manipulator. The shown object linings are the isosurfaces for 6, = 0 corresponding t,o shape finictions. The five links are considered as aluminium rods l m in length with constant rectangular cross-sections. The task in figure 2 is to cont,rol the EE on BE from the shown initial to the goal configuration in the presence of two squared obstacles. The duration is T = 20. A spline ansata with M = 10 is used. Prescribed is the initial configuration q (0) and the terminal EEconfiguration E (T). So (c') in (NLP) contains the five initial comt.raints qn (0) = 4: and the terminal constraint d ( E ( T ) ,E T ) = 0. The metric scaling a l p = 1 is set in (4). All controls have equal weights, U,, = 1. Considered are collisions of manipulator a i d obstacles according to the measure graph r" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000391_j100187a036-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000391_j100187a036-Figure1-1.png", + "caption": "Figure 1. A sketch of the model particles showing the dimensions of the parallelipipeds and a few of the six possible orientations.", + "texts": [ + " Our findings are primarily qualitative although we provide numerical trends in sections IV and V. An accurate quantitative treatment of I-N interfacial thermodynamics based on a realistic molecular Hamiltonian appears beyond reach at the present time. Some years ago, Barboy and GelbarP showed that the discrepancy between the calculated and measured strength of the I-N phase transition may lie in the assumption of uniaxiality in the hard-particle models. In particular, they considered rectangular cross-section parallelipipeds (see Figure 1 ) and restricted the orientations of the particles to the three Cartesian directions (an idea introduced earlier by Zwanzig? who treated only infiitely long parallelipipeds). They showed that even a moderate amount of particle biaxiality (always present in real systems) can decrease the density jump at the I-N transition by 1 order of magnitude. More recent ~ o r k ~ O J ~ - ~ ' on biaxial particles with unrestricted (1) Onsager, L. Ann. N.Y. Acad. Sci. 1949, 51, 627. (2) Zwanzig, R. J . Chem. Phys", + " We minimize eq 2 with respect to each density, for the case of a spatially homogeneous nematic ~ [ B A Q ( P I , P ~ , * * * , P ~ ) I / a P i = 0 (8) and require BAQmin = 0. The set of values p i , p2, ..., pq satisfying these equations are the densities of the nematic at coexistence. For the bulk calculation, the SDA form of the weighted densities becomes P / = (CPmV/m)/(CV/j) ( 9 ) m j where VIj is the pair excluded volume (covolume) between parallelipipeds with orientations I and j . 111. The Model: Hard Parallelipipeds with Restricted Orientations The model we investigate is that of hard parallelipipeds with dimensions L, B, and W (see Figure 1). We restrict the symmetry axes ,Of the blocks to lie along the three Cartesian directions i, j , or k . The interaction energy ulm(r - r\u2018) is infinite if two blocks overlap and zero if they do not. We denote the orientations using two indices, the first indicating the direction of the long (L) axis, the second the direction of the intermediate ( B ) axis. With this system of labels, the six possible unique pairs are as follows: (1) ij, (2) ik, ( 3 ) j i , (4) j k , ( 5 ) ki, and (6) kj . From the application of the constraint of uniaxiality, nematic order may be completely described by two parameters: x = the total fraction of blocks with lon (L) axes perpendicular to the nematic director (taken here as 8, and f = the fraction, among those whose long axes are perpendicular to the director, which are \u201cflat\u201d (i" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003102_cdc.1994.410974-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003102_cdc.1994.410974-Figure1-1.png", + "caption": "Figure 1: Switching controller structure.", + "texts": [ + " Furthermore, as we shall illustrate in our examples, the controller u1(z) may result in an unstable system if it is \"left on.\" That is, convergence to the origin is achieved for every z(to] E U1 only by switching from u,(z) to uo(z) once z(t ) E 240. We can extend this design procedure by constructing a third controller uZ(z1 which forces a set of trajectories that originate in U2 into 241, increasing the overall region of attraction to include U z ; the control is switched from u~(z) to u l ( z ) when z( t ) E U1. The paradigm is illustrated in Figure 1. We continue this design procedure until there exists no computable control law U N + ~ ( Z ) to force trajectories commencing from initial conditions outside UN into In Section 2, we present our main theoretical contributions. We focus on the design and construction of uI(z) and on the estimation of the regions of attraction Uo. The N > 1 case is not considered due to space limitations. Section 3 illustrates the controller's performance using the acrobot example. The acrobot is a non-trivial mechanical system which does not satisfy, to the best of the authors' knowledge, the sufficient conditions required for feedback linearization or for semiglobal stabilization [SI, [7)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001696_s0263-8223(01)00168-4-Figure9-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001696_s0263-8223(01)00168-4-Figure9-1.png", + "caption": "Fig. 9. Schematic diagram for the determination of: (a) Ky ; (b) Kxy ; (c) Kxz in surface-loading condition.", + "texts": [], + "surrounding_texts": [ + "Eighteen specimens were manufactured by circumferential winding of plain weave E-glass woven cloth impregnated with epoxy resin on mandrels. The plain weave E-glass woven cloth has the same fibre geometry and mechanical properties in the wrap and fill directions. Loss on ignition test [15] was performed for the determination of fibre volume fraction. Tensile tests were performed on rectangular strips of unreinforced epoxy resin according to the standards set by ASTM [16], to determine the engineering properties of the resin. The elastic modulus and Poisson\u2019s ratio of the epoxy were found as 2.93 GPa and 0.38, respectively. The elastic modulus and Poisson\u2019s ratio of the fibres were taken to be 75.9 GPa and 0.22, respectively [17]. A modified \u2018rule of mixtures\u2019 [17] was used as a mathematical model to predict the engineering properties of the composite springs. Schematic diagrams of the experimental setups for the determination of spring rates Ky ;Kxy and Kxz are shown in Figs. 8(a)\u2013(c), respectively, for line-loading condition. Two semi-circular metal bars were attached diametrically to the specimen for ease of loading. The two semi-circular metal bars were connected externally by bolts and nuts on their extended ends to prevent movement and rotation. In surfaceloading condition, as shown in Figs. 9(a)\u2013(c), both the upper and lower flat contact surfaces of the composite spring were sandwiched, bonded and bolted rigidly by two 15 mm thick aluminium plates such that strain energy is mainly stored in the flexible arms when under loading. All specimens were loaded uniaxially in compression by a digital force gauge for determination of Ky and loaded by dead weights for determination of Kxy and Kxz while the displacements were measured by a dial gauge. A total number of 18 0/90 woven E-glass/epoxy composite springs with flat contact surfaces were manufactured. These specimens were fabricated by winding 5-ply, 7-ply and 9-ply of E-glass cloth impregnated with Ciba\u2013Geigy epoxy resin on custom- 3 7 6 P .C . T se et a l. / C o m p o site S tru ctu res 5 5 ( 2 0 0 2 ) 3 6 7 \u2013 3 8 6 made mandrels of 38, 57 and 85.5 mm nominal outside radii and all have the same width of 50 mm and the same semi-included angle of 37 . The geometry and properties of these specimens are listed in Table 1." + ] + }, + { + "image_filename": "designv11_65_0001435_s0167-8922(02)80057-8-Figure9-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001435_s0167-8922(02)80057-8-Figure9-1.png", + "caption": "Fig. 9 A sample of hysteresis loop", + "texts": [ + "(o~) (1) the inertia force for a mass m can be found by the following correlation = 2 = __ I f mr[ m N o mlal mlJf I (2) solving the above equations simultaneously and eliminating the term \"60t\" yields the following equation _ X 2 F 2 2FXcos(O) + ~ : sin2(0) FTo (4) 3. HYSTERSIS LOOP For a mass-spring system with a single degree of elliptical geometry. Therefore, if the two signals are 520 plotted vs each other the results appears as an elliptical shape on the screen of Oscilloscope. Since the force and displacement is obtained in a cycle, therefore, the area occupied within the ellipse represents the damping and the slope of the force and displacement represent the stiffness. A sample of typical test data in the form of an ellipse is shown in Fig. 9. Depending on the value of 0, the following results might take place a. If 0=90 ~ then Eq.(4) reduces to F 2 X 2 - - + - - = 1 ( 5 ) F 2 Xo 2 which is equation of an ellipse symmetric about X and F axes. Furthermore, if Fo=Xo the result is a circle. b. If 0=0 ~ or 0 = 180 ~ then F=+(Fo/Xo)X. (the + sign is for 0=0 ~ where it reveals an equation of a line If the F/X signals are equal in magnitude, the patterns shown in Fig.10 can be obtained. 4. HYSTERSIS APPROACH The equation of motion for a harmonic forced vibration with dry friction can be represented as m ; + I~F" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000173_s0377-0257(97)00038-4-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000173_s0377-0257(97)00038-4-Figure1-1.png", + "caption": "Fig. 1. Diagram showing the definition of the angles, ~b and ~, together with the polar coordinates (p, 0, z).", + "texts": [ + "(k ) = i (~i1(2k2~ j -q- kje2) + ~jl(2k2t3 i + kie2 ) -Jr_ (~i2kll~ j q_ (~j2klei _~_ kl -~2 Using the results given in Eq. (15), the cycle-averaged rotation-rate of the fibre may be written a s where 3 dkp, (18) and = 3i j l (k \"p)(I - pp )\" (19) co) 2( )1 t~=i l - -~g r 1~--2z5 -~(k2l -kk) \" ok , (20) so that/~ is given by the following three-dimensional integral over Fourier space. 3c~\u00b0E2 1 +---d2~ ~ k4 ( t -pp) . (k2 t - kk). B.k. (21) /i - 8re 3 log r To evaluate this integral it is convenient to change to cylindrical polar coordinates (p, O, z) in Fourier space, about the axis of the fibre (see Fig. 1). The orientation of the fibre is now given in terms of two polar angles ~b and 0, where ~b is the angle between the fibre and the gradient direction (2-axis) and 0 is the angle between the flow direction (1-axis) and the projection of the fibre in the 1-3 plane. With respect to the original coordinate system q~ =Pl cos ~b cos 0 -P2 sin ~b +t53 cos ~b sin 0, (22) = - P l sin ~b sin ~k +,03 sin ~b cos ~b. (23) The integral in Eq. (21) is logaritmically divergent as p---, o% indicating that for high-aspect ratio particles the leading order contribution is of order log(r) and comes from large values of p" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003483_imece2004-59492-Figure10-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003483_imece2004-59492-Figure10-1.png", + "caption": "Fig. 10: Mesh of the fem model", + "texts": [ + " In the following the damping effect is neglected; indeed, the evaluation of c(t) is beyond the purpose of the present work. wnloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 02/07/2016 Term Stiffness evaluation A 2D finite elements analysis is applied to calculate stiffness. MSC-MARC 2003 is used to perform the analysis. In the analysis the whole gears are considered, neglecting teeth which are not involved in the meshing. In the present case only three teeth per gear are modeled, without loss of accuracy. More than 2400 quadrilateral elements (MARC Quad mesh) are used for each gear (fig. 10). The mesh of the teeth is more refined than the mesh of the wheels bodies, because the deformation is almost completely localized to the teeth. The driven wheel shaft is considered locked, i.e. the inner circumference of this wheel has zero rotation. This means that the driven wheel can be deformed around its shaft. The shaft elasticity is not considered here. The driver wheel can rotate around its shaft, i.e. its internal circumference has only rigid rotation. A torque of C=470.71878 Nm is applied to the driver wheel through its shaft The wheel material is an alloy steel with a Young modulus of E=2" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002350_robot.1998.680983-Figure2-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002350_robot.1998.680983-Figure2-1.png", + "caption": "Figure 2: Phases of casting manipulation", + "texts": [ + " In spite of its small size and simple mechanism, the manipulator can have a large work space by making effective use of its dynamics compared with the conventional manipulators. The manipulator uses substantially less energy to move and does not suffer moving tumbles. Therefore it is possible to use this manipulator to collect objects such as garbage floating on the sea, or in agriculture for collecting various kinds of Casting manipulation consists of the following five phases. These phases are illustrated in Figure 2. (1) Swing phase: To swing the rigid link until the necessary motion of the gripper to reach the target is generated. (2) Throwing and releasing phase: To throw the gnpper to the target by releasing the string at the suitable time. (3) Posture control phase: To control the posture of the gripper while it is flying to the target in order to catch the target. (4) Catching phase: To let the gripper grasp the target with the impact force. (5 ) Reeling up phase: To reel up the string with the swing motion of the rigid link, in order to collect the target" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000024_s0167-8922(08)70489-9-Figure7-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000024_s0167-8922(08)70489-9-Figure7-1.png", + "caption": "Figure 7 . Isometric view of oil film pressure on the conrod bearing of a high performance engine", + "texts": [ + " As shown in Figure 5, the severe wear marks appear at the bearing edge, starting from the split line of the cap bearing. A careful examination of the predicted results also shows that the pressure peak on the leading side of the cap bearing is higher that on the trailing side. For the oil film, a thinner film is associated with the leading peak. This is of interest since the wear marks are located on the leading side of the bearing. (Engine peed: 16000 rev/min; 359 degree crank angle) (360 degree crank angle, 7000 revlmin) In Figure 7, the pressure profile of a high performance engine connecting rod bearing is presented. The diagram is at 359 degree crank angle position. This is the time when the bearing experiences its maximum peak oil film pressure, and nearly the maximum inertial load. The pressure disaibution is quite different from the one produced at lower engine speed (Figure 6). The peak pressure regions have shifted right to the bearing split line, 4.2. Sapphire seizure test The Sapphire bearing test rig is widely used by bearing manufacturers for bearing material development and production quality assurance" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002214_robot.1986.1087733-Figure9-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002214_robot.1986.1087733-Figure9-1.png", + "caption": "FIGURE 9 SENSOR AND OBJECT COORDINATE FRAMES", + "texts": [ + " These values a r e t h e n p a s s e d t o s u b r o u t i n e CENTER which c a l c u l a t e s t h e x and y components of t h e c e n t e r of grav i ty , xcg and ycg using the equat ions: Once t h e c e n t e r of grav i ty is ca l cu la t ed the v a l u e s a r e p a s s e d t o program SENSOR and a binary image of t h e o b j e c t i s p lo t t ed on the r igh t s ide of t h e s c r e e n by subrout ine BINARY. The c e n t e r of g r a v i t y i s t h e n p l o t t e d i n t h e c o r r e s p o n d i n g p i x e l l o c a t i o n on the b inary image ( f ig s . 10 and 11) . R e p l a c i n g t h e i n t e g r a l w i t h a summation f o r t h e d i s c r e t e image, the genera l equat ions become: I = E x ' A Ixl = z y ' A Ix ly l = 2 2 Y' where A i s t h e a r e a of an element. a l l t h e elements have the same area. From Figure 9 t h e r e l a t i o n s h i p b Cx'y'A (10) I n t h i s c a s e etween the known c o o r d i n a t e s x , y , which a r e measured with r e s p e c t t o t h e s e n s o r a x e s and the coord ina tes x' , y ' i s seen to be: X I = x - x Y ' = Y - Ycg cg S u b s t i t u t i n g t h e a b o v e r e l a t i o n s i n t o t h e equations (10) gives: These equations are analogous to equation (1). E q u a t i o n (8 ) may now be used t o c a l c u l a t e t h e two v a l u e s o f 8 by s u b s t i t u t i n g i n t h e r e l a t i o n s f o r I X t , I I and & l y t g i v e n i n equat ions (11, (12) and ( 1 3 1 , -2X(x-xc 1 (Y-Y,~)A tan28 = (14) z(y-ycg) A - ~ ( x - x >2A 2 cg Two B A S I C s u b r o u t i n e s , SQUARE and ANGLE, were d e v e l o p e d a n d a r e u s e d t o d e t e r m i n e t h e o r i e n t a t i o n a n g l e O p " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000585_rob.4620121004-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000585_rob.4620121004-Figure1-1.png", + "caption": "Figure 1. The 6-6 Stewart platform.", + "texts": [ + " 0 1995 l o k ~ i Wiley 6 Sons, Inc. Journal of Robotic Systems 12(l O), 661 -676 (1995) 0 1995 by John Wiley & Sons, Inc. CCC 0741 -2223/9510100661-16 662 Journal of Robotic Systems-2995 1. INTRODUCTION Ever since parallel mechanisms were ranked as a promising alternative to serial chains for the realization of stiff, accurate manipulators, the direct kinematics of fully parallel manipulators has attracted the interest of researchers. As is known, fully parallel manipulators feature two rigid bodies (base and platform, see Fig. 1) connected by six variablelength limbs having extremity ball-and-socket kinematic pairs. By controlling the length of some or all the limbs, the platform is endowed with up to 6 degrees of freedom with respect to the base. The direct kinematics-also referred to as direct position analysis (DPA)-is aimed at finding, for a given choice of limb lengths, all possible locations of the platform with respect to the base or, equivalently, all possible assembly configurations of the manipulator. The need to determine every configuration generally suggests the adoption of analytical procedures leading to a final polynomial equation in one unknown. The roots of such an equation directly correspond to the sought-after assembly configurations. The direct kinematics of fully parallel manipulators is generally a demanding task, to the extent that no analytical-form solution has yet been found for the more general leg arrangement. Due to the number of distinct connection points on base and platform, such an arrangement (see Fig. 1) is usually referred to as the 6-6 Stewart platform. By adoption of a numerical technique known as the continuation method, it has already been possible to assess at 40 the number of its assembly configurations in the complex field.' Nevertheless, interest is still keen for a polynomial solution of the 6-6 Stewart platform. To gain the necessary knowledge for tackling the analytical-form DPA solution of the more general case, some researchers have diverted their attention to kinematic arrangements characterized by some peculiarities with respect to the 6-6 Stewart platform" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002435_bf03220888-Figure5-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002435_bf03220888-Figure5-1.png", + "caption": "Figure 5. A schematic drawing of an electron-beam melting (EBM) furnace showing major components, including the cold-wall copper refining hearth. This furnace can be used with bar or loose feed.", + "texts": [ + " Therefore, these systems have been limited to laboratory use only, Electron-beam melting furnaces in clude one or more electron-beam guns (up to 1.2 MW per gun), a cold-wall hearth or hearths, a method for with drawing an ingot, a vacuum pumping system capable of 10-6 torr, and a cham ber to house these components, New electron-beam furnaces have automatic gun beam control which allows sweep ing of the melt area to provide good melting and sound ingots as they are slowly retracted, as shown in Figure 5,33 Electron-beam melting is a method for making primary electrodes and fin ished slabs and ingots from scrap and sponge. The development of cold-wall 1990 March \u2022 JOM Ingot \"'-Rotating Non-Consumable ,..-----Electrode Cold\u00b7Wall Crucible Figure 4. A schematic diagram of an arc-melt, tilt-pour centrifugal atomizer. hearth melting has provided a means for eliminating inclusion-causing defects. Recen t programs have produced in-spec chemistries in titanium alloys with EBM.34 Axel Johnson Metals Corpora tion recently reported on a system of multiple hearths and shielding" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000601_920406-Figure3-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000601_920406-Figure3-1.png", + "caption": "FIGURE 3", + "texts": [ + " Special Test Method 1 (5): A test panel consisting of sheet metal of body panel thickness and visco-elastic material is excited mechanically by a stiff frame and shaker configuration (Figure 2). The performance is expressed in terms of loss factor. Special Test Method 2: The damping material is bonded to a 300 mm x 30 rnm x 0.8 mm thick steel test panel and supported along the nodal line. The test panel is then excited at the resonance frequency. The damping performance is computed using the decay rate technique (Figure 3). Special Test Method 3: This is a slightly different version of the Special Test Method 2. The damping material bonded to a 178 mm x 80 mrn x 0.8 rnm thick panel is freely sup orted on the nodal line and the damping P per orrnance is expressed in terms of decay rate at its resonance frequency (Figure 4). The above discussion identifies two major conclusions: The test methods follow either the decay rate technique or the bandwidth technique, and the results are expressed in terms of decay rate or loss factor, respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000095_0957-4158(94)e0025-l-Figure3-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000095_0957-4158(94)e0025-l-Figure3-1.png", + "caption": "Fig. 3. Changing ratio of diameters during winding.", + "texts": [ + " One might imagine from the simple geometry of the cone that the required compensator motion would be easily mechanically derived. In practice this is by no means the case. Firstly, because the helix angle at which the yarn is wound varies in a complex manner, partly because of the practical limitations of manufacturing the traverse cam which distributes the yarn along the cone, and secondly because the compensation required changes as the winding operation progresses and the package increases in size (Fig. 3). These problems have not prevented purely mechanical solutions from being engineered but they are complex and costly and generally contain sliding contact surfaces (e.g. cams) which need to be carefully shielded from the fibres and dust of the spinning mill. The mechanics of a mechatronic positively driven tension compensator can be much simpler. Figure 4 shows such a device [9]. The simple two-bollard principle of the passive spring compensator is retained, but the bollard disc is now positively driven by a small stepping motor, the motion of which is controlled by a single-chip microprocessor of the 8051 family" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003038_s0094-5765(01)00002-9-Figure2-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003038_s0094-5765(01)00002-9-Figure2-1.png", + "caption": "Fig. 2. Variable geometry manipulator showing: (a) single module with a pair of slewing and deployable links; (b) several modules connected to form a snake-like geometry.", + "texts": [ + " The velocity transforms decompose the system mass matrix into a product of matrices. The inversion of this new form of the mass matrix is computationally far less intensive. As most arithmetic operations in eqn (2) arise from the inversion of the mass matrix, the resulting algorithm is of O(N ) and hence considerably more eVcient. With this as background, Part I of the paper develops a rather general, three dimensional, order N Lagrangian formulation for a class of novel variable geometry manipulators with modules of slewing and deployable links (Fig. 2). The Oexible manipulator is supported by an elastic platform in an arbitrary orbit. The snakelike manipulator has several advantages: It reduces coupling eIects resulting in relatively simpler equations of motion and inverse kinematics; decreases the number of singularities; and facilitates obstacle avoidance (Fig. 3). Dynamics and control of such Multi-module Deployable Manipulator System (MDMS), free to traverse an orbiting elastic platform and carrying a payload, represent a challenging task" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001221_002071797222993-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001221_002071797222993-Figure1-1.png", + "caption": "Figure 1. Vectored thrust.", + "texts": [ + " It is assumed that the thrust can be vectored in some manner such as vectoring engine nozzles similar to the F-16 MATV or paddles that de\u00af ect the engine exhaust such as the HARV F-18 aircraft. The manner in which the thrust is vectored is neglected for the simple model that is proposed. In actuality, the manner in which the thrust is vectored is important to model accurately the propulsion and aerodynamic interaction e ects that are beyond the scope of this study. Thrust vectoring is modelled using fundamental notions from dynamics depicted in Fig. 1. The thrust components along the x and z body-axes for a thrust vector angle d ptv are given by XT = T cos d ptv ZT = T sin d ptv } (6) where T is the thrust. The pitching moment due to thrust vectoring is the product of Flight control for mixed-amplitude commands 1211 D ow nl oa de d by [ N ip is si ng U ni ve rs ity ] at 0 4: 39 0 3 O ct ob er 2 01 4 the thrust component along the z body-axis and the moment arm athr de\u00ae ned by the distance from the CG to the thrust application point. MT = - ZTathr (7) A complete description of the aerodynamic forces XA and ZA and moment MA and the engine thrust T has been provided by Stevens and Lewis (1992)", + " Control actuators Although the throttle is an input to the aircraft, it is not treated as a control input for design since it primarily controls the aircraft trajectory not attitude. The aileron and rudder are the lateral and directional control inputs and are not discussed either. The remaining controls are the elevator and thrust vectoring. The sign convention is that a positive elevator angle provides a negative or nose-down pitching moment. Similarly, a positive thrust vector angle provides a negative pitching moment as shown in Fig. 1. The elevator and thrust vector actuator dynamics are modelled as \u00ae rst-order \u00ae lters: d e d cmd e = 1 \u00bfes + 1 d ptv d cmd ptv = 1 \u00bfptvs + 1 \u00bfe = 0.0495 \u00bfptv = 0.07 \u00fc\u00ef\u00ef\u00ef\u00ef\u00ef\u00ef\u00ef\u00ef \u00fd\u00ef\u00ef\u00ef\u00ef\u00ef\u00ef\u00ef\u00ef\u00fe (11) Flight control for mixed-amplitude commands 1213 D ow nl oa de d by [ N ip is si ng U ni ve rs ity ] at 0 4: 39 0 3 O ct ob er 2 01 4 The elevator and thrust vector angles have the following displacement and rate limits: |d e| < 25\u00ca | \u00c7d e| < 60\u00ca sec- 1 |d ptv| < 17\u00ca | \u00c7d ptv| < 60\u00ca sec- 1 \u00fc\u00ef\u00ef\u00ef\u00ef\u00ef\u00ef \u00fd\u00ef\u00ef\u00ef\u00ef\u00ef\u00ef\u00fe (12) In this section, a control strategy is developed for the F-16 TV model" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001267_s0925-4005(96)01994-6-FigureI-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001267_s0925-4005(96)01994-6-FigureI-1.png", + "caption": "Fig. I. Optical arrangement for the measurements with the flow-t~'ough cell:", + "texts": [ + " Apparatus Fluorescence excitation and emission spectra as well as response curves of the sensing membranes were measured on an Aminco SPF 500 spectrofluorometer equipped with a 250 W tungsten halogen lamp as a light source along with a red-sensitive detector, and linked to a HP 9815A desk calculator. Response curves were recorded by placing the membranes in a flow-through cell to form one wall of the c~ll. Excitation light hits the sensor membrane from outside (after passing the glass wall of the flow cell and the polyester support), and fluorescence is detected at an angle of 55 \u00b0 relative to the incident light beam as shown in Fig. i. Buffers and buffered sample solutions were p~mped through the cell at a flow rate of 1.5 ml rain-s. Excitation and emission wavelengths were set to 550 and 590 nm, respectively, and bandpasses to 5 nm. All experiments were performed at 22 + 2\u00b012. Solutions were obtained by dissolving 2.5 mg PVC or PVC copolymer, 5 mg of the respective plasticizer, 0.57 mg TDMACI and 0,44 mg RBOE in 1.5 ml of tetrahydrofuran (or ethyl acetate in the case of PVC copolymer). A dust-free 12 mm\u00d7 50 mm 175 p~m polyester foil (Mylar, type GA- 10, from DuPont) was placed in a desiccator containing the solvent (THF or ethyl acetate)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001472_s0167-8922(01)80115-2-Figure5-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001472_s0167-8922(01)80115-2-Figure5-1.png", + "caption": "Fig. 5\" Insertion of an Ni-CrNi thermal sensor into the inner-ring flange of a tapered roller bearing.", + "texts": [ + " Test bearings Tapered roller bearings of size 31312 were used; these were bearings specially manufactured for the FE 8 testing machine, and the inner ring retaining flange had a smaller diameter than is usual with series bearings. This allows easy separation of the inner ring and the set of rollers with cage. All the test bearings were from the same production batch. In order to determine the temperature in the immediate vicinity of the contact between inner-ring flange and roller ends, a hole was made in the flange of the inner ring which extended up to approx. 0.2 mm below the raceway. A temperature sensor (NiCrNi) was glued into this hole (Fig. 5). For comparison, Fig. 7 shows the results of a long-time test with grease B under the same experimental conditions. A telemetric measurement system transmits the thermoelectric voltage from the rotating shaft. 3. Results of long-time tests The basically different behaviour of greases A and B, which had already been observed in practice, was demonstrated in a variety of long-time tests. The complete results are reported in [ 13]. Fig. 6 is an example of the result of a long-time test with grease A at Fax = 50 kN, n = 75 min -~ and T = 10~ at the outer ring of the bearing" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002425_cbo9780511529627.012-Figure9.2-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002425_cbo9780511529627.012-Figure9.2-1.png", + "caption": "Figure 9.2. A machine part is located in space by a sequence of frames consisting of axes S ; and their common normals A,;.", + "texts": [ + " The mathematical relation that defines the position of each machine part M in the frame F is called its kinematics equations. If we model the local geometry of higher-pair joints by using rotary and sliding joints, then the position of every component of a machine can be obtained from a sequence of lines representing the axes Sy of equivalent re volute or prismatic joints. Between successive joint axes, we have the common normal lines Aijy which together with the joint axes forms a serial chain, as shown in Figure 9.2. This construction allows the specification of the location of the part relative to the base of the machine by the matrix equation [D] = [Z(0i, pi)][X(ai2, , p2)] \u2022 _i> w , flm_i,w)][Z(0m, pm)], (9.4) Cambridge Books Online \u00a9 Cambridge University Press, 2009https://www.cambridge.org/core/terms. https://doi.org/10.1017/CB 9780511529627.012 Downloaded from https://www.cambridge.org/core. UCL, Institute of Education, on 17 Jul 2018 at 11:52:53, subject to the Cambridge Core terms of use, available at known as the kinematics equations of the chain (Paul, 1981; Craig, 1989)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001007_jsvi.1997.0998-Figure13-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001007_jsvi.1997.0998-Figure13-1.png", + "caption": "Figure 13. A model of two-phase induction motor.", + "texts": [ + " Belt systems are usually driven by a three-phase induction motor with a squirrel cage rotor. Therefore, to analyze the transient responses of belt systems, the transient output torque of the induction motor must be examined. As is well known, a three-phase induction motor can be transformed to an equivalent two-phase induction motor (see, for example, reference [4]). Therefore, in this paper, the transient output torque of an induction motor is examined by using the two-phase induction motor model shown in Figure 13. Moreover, by using a d\u2013q transformation, a rotor coil can be transformed to an equivalent stator coil (see, for example, references [5, 6]). Therefore, a three-phase induction motor is transformed to an equivalent two-phase induction motor whose two windings in the rotor and stator are fixed in the d- and q-axes. In the equivalent two-phase induction motor, the relationship between voltages and currents is expressed as V=Ri+Li + p 2 vmGi, (16) where R1 0 0 0 L1 0 Lm 0 0 R1 0 0 0 L1 0 Lm R=G G G K k 0 0 R2 0 G G G L l , L=G G G K k Lm 0 L2 0 G G G L l , 0 0 0 R2 0 Lm 0 L2 0 0 0 0 0 0 0 0 G=G G G K k 0 Lm 0 L2 G G G L l\u2212Lm 0 \u2212L2 0 i= {i1d , i1q , i2d , i2q}T, V= {V1d , V1q , V2d , V2q}T, L1 =Lm +L'1 , L2 =Lm +L'2 " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000471_bf01178518-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000471_bf01178518-Figure1-1.png", + "caption": "Fig. 1. Notation for displacements and stresses at the contact interface layer", + "texts": [ + " A micromechanical approach to the analysis of shakedown in a surface layer with specified asperities was considered by Johnson and Shercliff [12]. Such asperity models could also provide quantitative assessment of parameters used in the present model. Some general properties of slip rules discussed for elastic, friction contact in [2] will now be incorporated into the model formulation. Assume a contact layer of thickness to adjacent to the contact plane to represent local effects occurring during slip of two bodies. Figure 1 illustrates the notation used for displacements and stresses at the contact. In the contact plane the coincident points M and N are assumed to belong to two bodies. After loading the relative displacements of M and N are ul, u2 and v, where ul and Z. Mrdz and A. Jarz~bowski u2 are the tangential components in the orthogonal reference frame and v is the displacement normal to the contact plane (compacting normal displacement is assumed as positive). The engineering strain components within the contact layer are U 1 U 2 V ~I : --, ~)2 : --, 8 n : --, (1) to to to where the small strain definitions are used" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002250_978-3-662-04068-3_8-FigureB.2-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002250_978-3-662-04068-3_8-FigureB.2-1.png", + "caption": "Fig. B.2. The action could successfully be demonstrated by using this system. The PVC arm showed flapping action with the rate of 2 Hz at I x lOsYm- 1 in air. Used gel was the same as that shown in Fig.B.l", + "texts": [], + "surrounding_texts": [ + "8.1 Introduction Polymer gel actuator has been extensively studied from the standpoints of new intelli gent materials. When we define the intelligent material as one that has three functions in a body - sensing, processing, and response or action - polymer gels are adequate as intelligent materials, since they respond by taking various stimuli in the surroundings into account. Polymer gel actuator can be considered, in this sense, as one of the intel ligent materials or artificial muscle, which can actuate by sensing various triggers or stimuli to generate strain, taking various factors into account. There has been vast var iations of reports on swelling-and-deswelling features induced by changing ionic strength, pH value, solvent composition, and temperature [1-5]. In polyelectrolyte gels, a strain can also be induced by an electric field [6-10]. The electrically induced strain has been considered to be caused principally by electrostatic repulsion or attrac tion among the ionic species in the gels. The other type of approach is a magnetostric tive action of a ferrofluid immobilizing hydrogel, in which the superparamagnetic properties of the ferrofluid has been utilized and the dispersion mode of the ferro mag netic particles has a serious effect on the strain generation [11-15]. Here we will introduce our results on the electrostrictive polymer gels which are swollen with non-ionic organic solvents and which can be actuated by applying an electric field. The gels showed very quick motion compared to the conventional poly electrolyte gels with application of a direct current electric field. We will show the fea ture of the action and the structure change of the gel induced by the electric field, dis cuss the mechanism of the electrostrictive motion of the gels, and make some sugges tions to the future development of this type of material. 8.2 Electrically Induced Strain in PVA-DMSO Gel [16] 8.2.1 Electrostrictive Motion of PVA-DMSO Gel Poly(vinyl alcohol) gel swollen with dimethylsulfoxide (DMSO) can be actuated by ap plying an electric field as shown in Fig. 8.1. The gel contains 98 wt% of DMSO and" + ] + }, + { + "image_filename": "designv11_65_0003249_cdc.1989.70168-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003249_cdc.1989.70168-Figure1-1.png", + "caption": "Figure 1. The structure of the musculoskeletal system of a cat hindlimb in the sagittal plane. Heavy lines denote skeletal segments. Light lines denote tendons. Intermediate lines denote muscles. Muscles are grouped according to attachment geometry. The circle at the knee indicates that the patella is modelled as a pulley. The circle at the ankle denotes a pulley that models the ankle crossing of several dorsi-flexors of the toes.", + "texts": [ + "1 T h e Dynamics of t h e Musculoskeletal Mechanics Mechanical structures similar to the articulated leg in our model have been studied extensively in both robotics research and biomechanics. Therefore the derivation of the dynamics is omitted. The major differences are in the choice of coordinate systems and the generation of joint torques. We use intersegment joint angles ( ( p h , p k , pa)' = p) as the generalized coordinates. The joint torques are generated by a set of 10 musculotendon actuators. This can be seen from figure 1 where the mechanical structure of the svstem is rewesented. The dynamical equations of the system for a standing posture are of the form: where J(.): the moment of inertia matrix; M ( . ) : the vector of centrifugal and Coriolis forces; N ( .): the coefficients for the gravitational force; 519 T ~ g: the gravitational const ant; (%)': the matrix of moment-arms for musculotendon actuators. The musculotendon lengths L, = (L,1, L p z , . . . , Lplo)' are determined from the limb configuration and the attachment geometry" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001271_1.2833499-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001271_1.2833499-Figure1-1.png", + "caption": "Fig. 1 Three-dimensional contact fatigue model", + "texts": [ + " In order to investigate how the cracks interact with adjacent cracks, a three-dimensional fracture mechanics model is developed in this paper, where an elastic half-space con taining multiple cracks is subjected to the rolling/sliding con tact. The rolling contact fatigue is simulated by a cyclic Hertzian contact loading moving across the surface of the half-space. A three-dimensional fracture analysis is applied to determine the three modes of stress intensity factors around the crack fronts, where the mixed mode crack growth under contact loading is simulated based on the modified Paris law. Contact Fatigue Model A three-dimensional contact fatigue model is considered as shown in Fig. 1, where an elastic half-space containing n multi ple cracks is subjected to a Hertzian line contact loading moving across its surface. Although Fig. 1 shows a series of coplanar Journal of Tribology JULY 1997, Vol. 1 1 9 / 3 8 5 Copyright \u00a9 1997 by ASME Downloaded From: http://tribology.asmedigitalcollection.asme.org/ on 01/28/2016 Terms of Use: http://www.asme.org/about-asme/terms-of-use cracks, the formulation and numerical method described in this paper are for the general multiple cracking problem where the cracks may have arbitrary shapes and locations. Two coordinate systems, x for the global coordinates and x ' for the crack coordi nates, are used" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003012_1350650011543637-Figure2-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003012_1350650011543637-Figure2-1.png", + "caption": "Fig. 2 Sector-shaped tilting-pad thrust bearing", + "texts": [], + "surrounding_texts": [ + "The equations that are needed to determine the pressure distribution for hydrodynamic bearings, such as those analysed in this report, are the classical Reynolds equation in combination with the film thickness equation. The film thickness equations in their dimensionless forms are h0(x0) \u02c6 1 \u2021 k \u00a1 kx0 (1a) h0(r0, j) \u02c6 1 \u2021 k r0 sin(\u00f5 \u00a1 j) sin(\u00f5) (1b) h0(j) \u02c6 1 \u00a1 \u00e5 cos(j) (1c) for the rectangular tilting-pad thrust bearing, the sectorshaped tilting-pad thrust bearing and the circular journal bearing respectively. The dimensionless Reynolds equation for hydrodynamic lubrication is @ @x0 h3 0 @ p0 @x0 \u00b4 \u2021 @ @ y0 h3 0 @ p0 @ y0 \u00b4 \u02c6 6 @h0 @x0 (2a) @ @ r0 r0h3 0 @ p0 @ r0 \u00b4 \u2021 1 r0 @ @j h3 0 @ p0 @j \u00b4 \u02c6 6r0 @h0 @j (2b) @ @j h3 0 @ p @j \u00b4 \u2021 @ @ y0 h3 0 @ p @ y0 \u00b4 \u02c6 6 @h0 @j (2c) for each of the three types of bearings previously mentioned. The boundary conditions are given by zero pressure at all edges for all three bearings. Since the first two thrust bearings only have a converging film thickness variation, no cavitation occurs in these two bearings. For the journal bearing, the boundary conditions for the start of Proc Instn Mech Engrs Vol 215 Part J J01001 # IMechE 2001" + ] + }, + { + "image_filename": "designv11_65_0003221_50011-7-Figure11.4-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003221_50011-7-Figure11.4-1.png", + "caption": "Figure 11.4. A three-cable system", + "texts": [ + " Ideally, the measurement system should be accurate, inexpensive and should be operated automatically. The goal is to minimize the calibration time and the robot unavailability. At this time, such devices are not yet available. Nevertheless, we present in this section sonie principles that have given place to industrial realization. 11.8.1. Three-cable system Such a system is basically composed of three high resolution optical encoders PI, P2, P3. Low mass cables are fixed to one of their ends on the encoder shafts whereas the other ends are fixed on the endpoint M of the robot (Figure 11.4). The encoder readings give the cable lengths, which represent the radii pi, p2, P3 of three spheres whose centers are on the encoder shafts. The intersection of the spheres determines the coordinates of M. This low cost device provides automatically the coordinates of the endpoint M. As a commercial example of such a system, we can mention the 3D CompuGauge from Dynalog whose accuracy is about 0.1 mm for a cubic measuring space of 1.5 m of side. 11.8.2. TheodolUes A theodolite is a telescope where the two angles giving the orientation of the line of sight can be measured precisely" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002743_bi002891t-Figure5-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002743_bi002891t-Figure5-1.png", + "caption": "FIGURE 5: Model depicting possible interactions between PLB and the Ca-ATPase. PLB is shown as a monomeric unit interacting with each Ca-ATPase; two Ca-ATPase-PLB heterodimers interact as a functional unit. Activation of the Ca-ATPase by PKA is suggested to require the phosphorylation of two PLBs within a dimeric complex containing two Ca-ATPase polypeptide chains. Phosphorylation of PLBs results in conformational changes to both PLB and the Ca-ATPase, including the spatial rearrangement of Ca-ATPase polypeptide chains with respect to one another, as previously shown to be associated with enzyme activation by PKA (27, 59). Conformational changes of PLB, largely undefined as yet, exclude its dissociation from the Ca-ATPase (61, 64, 67). Larger oligomers of PLB (not shown) are not directly involved in the functional regulation of the Ca-ATPase. The overall shape of the Ca-ATPase is highly asymmetrical, and was derived from the 14 \u00c5 resolution image obtained using image-enhanced cryo-electron microscopy (65).", + "texts": [ + " We find that upon increasing the fraction of FITC-PLB in reconstituted membranes, there is a second-order loss in the ability for PKA to activate the Ca-ATPase (Figure 4), indicating that activation of the Ca-ATPase requires the activation of multiple PLB molecules within a quaternary complex containing the Ca-ATPase. These results suggest that functional linkages between PLB polypeptide chains are important to Ca-ATPase activation, either by direct linkage or through interacting Ca-ATPase molecules (Figure 5). Relationship to PreVious Results. PLB represents a major target of the \u00e2-adrenergic cascade in the heart, and has been shown to function as a major regulator of the Ca-ATPase in cardiac SR membranes (1, 46-48). The enhanced calcium sensitivity of the Ca-ATPase concomitant with PLB phosphorylation results in increased myocardial contractility. The nature of the structural interaction between PLB and the Ca-ATPase has attracted considerable attention, as the disruption of this interaction may represent an important target for drug development that could enhance the calcium handling properties of the failing heart (49)", + " The second-order decrease in the extent of activation of the Ca-ATPase co-reconstituted in the presence of variable fractions of FITC-PLB observed in this study indicates a requirement for two PLB molecules in the PKAinduced activation of the Ca-ATPase (Figure 4). As an illustration of the interpretation of this kind of data, the effects on activation by random modification with FITC of 50% of the PLB molecules can be considered. For example, random modification of 50% of PLB in a dimeric arrangement with a functional unit (either monomer or oligomer) of the Ca-ATPase would result in 75% loss of PKA-induced activation of the Ca-ATPase, i.e., second-order inactivation, as observed here (Figure 5). In contrast, random modification of 50% of PLBs that are arranged as one PLB per functional unit of the Ca-ATPase would result in 50% loss of activation, i.e., first-order inactivation. For an alternate arrangement in which pentameric PLBs are associated with the Ca-ATPase in such a manner that two PLB subunits directly interact with the Ca-ATPase, random modification of 50% of the PLBs would be expected to result in a 40% loss of activation. Thus, the latter two models do not fit the inactivation data (Figure 4). While these data do not address the functional unit of the Ca-ATPase, previous cross-linking and radiation inactivation experiments have indicated a dimeric Ca-ATPase (56-59). Therefore, a possible model for these interactions involves a functional complex consisting of two heterodimers of PLB associated with the Ca-ATPase (Figure 5). The free pool of PLB may exist predominantly as oligomers that under some conditions are capable of exchange with PLB bound to the Ca-ATPase (19, 22). That the monomeric form of PLB is able to fully regulate Ca-ATPase function suggests a monomeric intermediate between free oligomers and bound PLB peptides (19, 20, 23). This model also reflects previous experimental observations. Based on site-directed mutagenesis, sequences in both transmembrane and cytosolic portions of PLB and the Ca-ATPase are critical to the inhibition of Ca-ATPase function, and are presumably involved in intraprotein interactions (4, 5, 20, 60, 66)", + " Thus, while FITC-PLB undergoes normal inhibitory interactions with the Ca-ATPase, the second-order loss in the ability to activate the Ca-ATPase by PKA upon reconstitution of varying fractions of FITCPLB indicates that the activation of multiple PLB molecules is necessary for activation of the Ca-ATPase. Since sitedirected mutations that stabilize the monomer of PLB are able to fully activate the Ca-ATPase, these results suggest that a single PLB molecule binds to each Ca-ATPase to form a functional complex consisting of two heterodimers (Figure 5). Thus, in the presence of saturating PLB (as used in this study), PKA-dependent activation of both molecules of PLB is necessary to promote conformational changes within individual Ca-ATPase polypeptide chains necessary for enzyme activation. This mechanism of regulation has the potential to fine-tune the activation of the Ca-ATPase through the \u00e2-adrenergic cascade, such that a threshold level of PKA activation is necessary to result in enhanced Ca-ATPase activity. 1. Kirchberger, M. A., Tada, M" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003067_iros.1998.724635-Figure4-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003067_iros.1998.724635-Figure4-1.png", + "caption": "Fig. 4: The projected two object system", + "texts": [ + "(ll), dimDB = (Cj\u201d=, 3cj +3r) x 6m and dimDF = (Cj\u201d=, 3 c j +3r) x Cyx, sj , where sj shows the number of joints of finger j Note that, while we dealt with a 3D model, for a 2D model, the skew-symmetric matrix equivalent to the vector product ax is redefined as ( a x ) = [-av a,], and dimDB = (cj\u201d=, 2cj + 2r) x 3m. 4 Kinematics for Enveloping In this section, we consider whether two objects can be lifted up by a simple pushing motion(Fig.1). As shown in Fig.3, the common tangential plane of two objects are defined as n. The plane which is normal to II and tangent to the gravity vector is defined as I?. We consider the motion of the objects projected on I?. The rolling condition is assumed to be satisfied at each contact point. The kinematic relationship between the objects projected on r is shown in Fig.4 where the suffix y denotes a vector on the two dimensional plane r. For simplicity, we assume that an object contacts with one finger at one point or that contact points between an object and fingers are overlapped when they are projected on r. Now we provide the following defini- tion. jDefinitioni W h e n two o jects rotate in opposite direction such that both center of gravity may be close to the palm, we call such a phase palm-reaching phase For the objects being in palm-reaching phase, object 1 and object 2 have to rotate counter-clockwise and clockwise, respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003034_cdc.1994.411508-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003034_cdc.1994.411508-Figure1-1.png", + "caption": "Fig. 1 - Definition of coordinate systems and errors", + "texts": [ + " For certain passenger cars schemes like proportionality of the rear steering angle to that of the front wheel, in the same or opposite direction, have been considered by car manufacturers. The contribution of this paper is twofold. First, based on the kinematics of motion the control problem is formulated. Then, the stabilization and error correction in ath following is investigated for a number of control schemes. %he results are applicable to mobile robots equipped with double steering, in view of their low operating speeds, and to any four wheel steering vehicle when operating at low speeds. Figure 1 shows the schematic of the telescopic model of a double steering vehicle and its configuration with respect to a desired path. The momentary coordinate system XPY is attached to point P of the desired path, which is the nearest point on the path to int C, the mass cenbe of the vehicle. PX is tangent to the p a t r a t this moment and PY is perpendicular to PX. Similarly, at point C the coordinate system UCW is attached to 0-7803-1 968-0/94$4.0001994 IEEE the vehicle, CU being the longitudinal axis and CW the lateral axis" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002252_robot.1997.614367-Figure2-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002252_robot.1997.614367-Figure2-1.png", + "caption": "Figure 2. Task frame-sensor frame definitions for a stereo camera system and a manipulator force sensor.", + "texts": [ + " Our object schema definition includes a geometric environment model of an object augmented by sensor mappings that describe how the object will be perceived by the actual system sensors. Forward projective sensor mappings and the associated Jacobian matrix for each sensor that exists in the system provide this representation. Figure 1 shows a diagram of an object schema, including the current internal pose of the geometric model; its desired pose (for controllable schemas) as determined by a supervisor; and sensor mappings used to direct variable sensor parameters as well as update the current internal pose of the schema based on actual sensor feedback. Figure 2 shows sensor mappings for vision and force sensors. The pseudoinverse of the Jacobian mapping used for determining sensor resolvability is also used to \u201cservo\u201d the geometric model in order to reduce errors between the current internal visual representation of the object and the actual visual representation. The concept of sensor resolvability, which we proposed in [ 111, provides a measure of the ability of both force and vision sensors to resolve positions and orientations in task space. This provides a technique for assimilating the data from the two sensors", + " Vision Resolvability ply perspective projection mappings and are of the form 6 ~ s = J(q)SX, (1) For a single camera, forward sensor mappings are sim- fYC S,ZC Ys = - where xs and ys are the projected image coordinates of a point on the object in the internal representation located with respect to the camera frame at (Xc,Yc,Zc), f is the focal length of the camera lens, and sx and sy are the pixel dimensions of the CCD array. For the experimental results to be presented, an orthogonal stereo pair is used. Figure 2 shows the coordinate frame definitions for this type of camera-lens configuration. If the axes are aligned as shown in the figure, the Jacobian mapping from task space to sensor space for a single feature can be written as In (3), we assume the camera-lens parameters are identical for both cameras. The other terms in (3) correspond to Figure 2. 2.3. Force Resolvability For force sensor resolvability, the Jacobian mapping of the force sensor is represented by where Sxs is the infinitesimal displacement vector in force sensor space and is measured from strain gauges mounted on the force sensor body; J,(t) is the Jacobian mapping and is time-variable; and SX, is the infinitesimal displacement vector in task space. Figure 2 shows a typical wrist force sensor mounted at a manipulator end-effector and the associated coordinate frame definitions. Force sensing is based on Hooke's law and is a highly linear process assuming induced strains remain within the elastic range of the material of the force sensor body. Through a quasi-static system stiffness analysis described in [ll], the Jacobian mapping can be written as (5) where C, is the force sensor calibration matrix; JST is the mapping shown in Figure 2; J,,,,(0) is the manipulator Jacobian matrix and varies with 0 , the vector of joint positions; and K, is the joint stiffness matrix. 6xs = J&r)SX, (4) T -T J,(O = CSJsTJM (e)K,JZ(0) T . Resolvability In order to perform a comparison of the resolvability of force and vision feedback, the variance of sensor noise must be considered in terms of the resolvability of the sensor. This is given by oT = J+(k)os (6) where oT is the vector of positional variance in task space; Jc(k) is the pseudoinverse of the sensor Jacobian; and os is the vector representing measurement variance in sensor space" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000471_bf01178518-Figure4-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000471_bf01178518-Figure4-1.png", + "caption": "Fig. 4. Various concepts of back-stress relaxation: a stress reversal point sliding on stress reversal surface, b shrinking of the stress reversal surface, e radial mapping center moves towards the point C", + "texts": [ + " Figure 2 c shows the evolution of active loading and yield surfaces for superelliptical forms of these surfaces, where 1 < n < 2 for a < eL and n > 2 for a > eL, and eL denotes the stress value at the connecting line between two superellipses. Consider now the unloading process, with the stress path penetrat ing inside the loading surface. It can be assumed that the active loading surface F2 -- 0 develops along the unloading path so that it passes through the actual stress point P and remains tangential to the pr ior Phenomenological model of contact slip 65 loading surface F~ = 0 associated with the stress unloading point, Fig. 4 a. In the multisurface formulation the active unloading surface F2 = 0 was tangential to the prior loading surface F1 = 0 at the stress reversal point R 1. Now, however, we shall assume an additional relaxation of back stress during the unloading process, so the tangency point is assumed to translate towards the point C on the o--axis. In Fig. 4a, there are two unloading paths shown, R1P ~, R2p 2 and the respective position R 1' and R 2' of tangency points sliding along the ellipse F~ = 0. A similar translation of conical loading surfaces towards the vertex of the limit surface was observed by Jarz~bowski and Mr6z [2] in the case of two elastic sliding spheres. In developing the constitutive model for sand (where it was assumed that C = 0), the same writers [5] assumed similar translation of the tangency points of consecutive loading and unloading surfaces towards the center 0. An alternative description of back stress relaxation was presented by Mrdz et al. [7] by assuming that during the unloading program the yield surface remains tangential to the stress reversal surface which remains tangential to the loading surface at C and its center translates towards C during unloading. Figure 4b presents two unloading paths R~P ~ and R 2 P 2 with two associated stress reversal surfaces Fr 1 = 0 and F, 2 = 0. The positions of the active unloading surfaces F2 ~ = 0 or F22 = 0 are now uniquely specified by the positions of stress points Pt or p2 and the stress reversal surfaces. The tangency points now translate towards C along the lines C R ~ and C R 2. Figure 4 c presents a similar concept of evolution of back stress by assuming that the center of mapping S of all nested loading surfaces lying within the domain F~ < 0 translates along the radial line C R 1 (or C R 2 - both cases are indicated in Fig. 4c) towards C. The center of mapping specifies the positions of all loading surfaces of given diameter. In fact, if the position of S ~ is known, then the centers of all surfaces lie on the segment S let. Similarly for the center position at S 2, the centers of all potential loading surfaces lie on the segment S Z c q . The idea of specifying evolution rules for mapping centers was introduced by Hashiguchi [8] in his subyield surface model formulated for both metals and geomaterials. Once the positions of stress points P~ or p2 and of mapping centers S* or S 2 are known, the active unloading surfaces F21 = 0 and F22 -- 0 are uniquely specified. In this paper, we shall follow the description used by Mr6z et al. [7] for the case of cyclic loading of clays and sands and assume that the size of the stress reversal surface F, = 0, cf. Fig. 4 b, evolves according to the relation d r l = w i d e , S - wgld~l (14) where r l represents the diameter of the stress reversal surface, wl and w2 are material parameters of the contact layer, and de,*, d3,* are normal irreversible strain and the tangential slip increments, respectively, specified by (1) and (2). Referring to Fig. 4 b and assuming that the tangency point R' translates along the line C R ' , we can write the evolution rules for the center of the stress reversal surface (cq ~, e~) and for the tangency point R', namely d e j = d r a , de1 ~ = 0 (15) and for the point R', R ~ + C R ~ d R ~ - - - d r 1 , d R ~ = - - d r 1 , (16) r i r l where R ~ and R e denote the coordinates of R'. Fo r the un load ing process, the active un load ing surface F2 = 0 remains tangent ia l to the stress reversal surface F, = 0 at R', so we have = - - + - 1 = 0 , (17) where c~/, c~2 ~ specify the center of the un load ing surface F 2 = 0, and r 2 denotes its size parameter " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002371_iros.1991.174511-Figure3-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002371_iros.1991.174511-Figure3-1.png", + "caption": "Figure 3: Applied forces at the reference point and reaction forces at the contact point for the contact of Figure 1 .", + "texts": [ + " Just as we have modelled the normal reaction force, so too can we model the frictional reaction force. We can think of friction as acting tangenoally to the physical edge of contact. Let t, denote the unit tangent to the edge of contact. Then t, must be of the Friction acts along this tangent through the point of contact. For a unit frictional reaction force, the induced torque about the center of mass is therefore vq, with vq = t, A D r form t, = *(ny, -nx). Observe that vq = ~ ( 1 1 , rx + ny ry ) . L,et us now write down the equations of motion. Figure 3 depicts a force-body diagram for the contact of Figure 1. Lnt FA = (Fx3 F y . F s ) be a generalized applied force. In other words, the applied Cartesian force is (F,.F,,) and the applied torque is 7 = ,)Fy. This force is applied at the center of mass, that is, at the reference point. Let the normal reaction force at the point of contact have signed magnitude f n r and let the frictional reaction force have signed magnitudef,. We measure f along the outward normal no, and f along the tangent vector t," + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001646_elps.1150030505-Figure7-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001646_elps.1150030505-Figure7-1.png", + "caption": "Figure 7. Isotachopherogram obtained at constant potential difference. Electrolytes contain impurities (A and B) with mobilities intermediate between leading and terminating ion.", + "texts": [ + "D.) are 0.326 2 0.007 and -3.034 + 0.003 for slope and intercept, respectively. A correlation coefficient (r) of 0.997 is obtained. The data show that once the terminating electrolyte is fully dissociated, the AM becomes independent of pH. In these conditions, an AM of 41.5.10-4. sec-' is calculated for fully dissociated acetic acid. Impurities present in the electrolyte solutions, with mobilities intermediate between the leading and terminating ion, give rise to isotachopherograms as shown in Fig. 7. The AM of the terminating ion, however, is influenced only if the mobilities of the impurities are similar to the mobility of the terminating ion. The determination of point I, Fig. 7, is hampered when there is a minor difference in mobility between leading and terminating ion. Point 11, Fig. 7, can be determined without problems. 3.4 Reproducibility of the measurement of the apparent mobility Reproducibility of the measurement of the AM was investigated by the percentage distribution of the coefficient of variation of 338 experiments performed in duplicate. The 268 H. Carchon and E. Eggermont Electrophoresis 1982, 3, 263-274 data are not normally distributed. The coefficient of variation (%) ranged from 0 to 9.1, with a median value of 1.2. As the 97th percentile equals 4.2, it can be concluded that the measurement shows satisfactory reproducibility" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002166_20020721-6-es-1901.01108-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002166_20020721-6-es-1901.01108-Figure1-1.png", + "caption": "Fig. 1: 2-sliding mode", + "texts": [ + " Thus, sliding modes 0 may be classified by the number r of the first successive total derivative (r) which is not a continuous function of the state space variables or does not exist due to some reason like trajectory nonuniqueness. That number is called sliding order (see (Levant, 1993; Fridman and Levant, 1996) for the formal definitions). Hence, the r-th order sliding mode is determined by the equalities = = = ... = (r-1) = 0 (1) which impose an r-dimensional condition on the state of the dynamic system (Fig. 1). The standard sliding mode used in the most variable structure systems (VSS) is of the first order ( is discontinuous). Consider a dynamic system of the form x = a(t,x) + b(t,x)u, = (t, x), (2) Here x R n , a, b, are unknown smooth functions, u R, n is also uncertain. The relative degree r of the system is assumed to be constant and known. It is supposed that 0 < Km LbLa r-1 KM, | La r | L (3) for some Km, KM, L > 0. Since LbLa r-1 = u (r) , La r = (r) |u=0 , conditions (3) are reformulated in terms of input-output relations", + " It is easy to build such a sequence of ellipsoids Bi, Bi% Bi+1, using transformation (12), that their union covers the whole space and all trajectories of (9) - (11) starting in Bi+1 enter Bi, i = 0, 1, ..., in finite time and stay there forever. Thus, B0 is a finite-time attracting set. The convergence time is easily estimated. Now taking any > 0 and applying transformation (12) with # = ( / 0) 1/r achieve the desired asymptotics of the attracting set. & Theorems 1, 3 are proved in a similar way. Consider a variable-length pendulum control problem. There is no friction. All motions are restricted to some vertical plane. A load of known mass m is moving along the pendulum rod (Fig. 1). Its distance from O equals R(t) and is not measured. An engine transmits a torque w which is considered as control. The task is to track some function xc given in real time by the angular coordinate x of the rod. The system is described by the equation ' '( = - 2 R R) x* - g R 1 sin x + 2 1 mR w, (13) where g = 9.81 is the gravitational constant, m = 1 was taken. Let 0< Rm + R + RM, R, , R,, , cx- and cx-be bounded, . = x-xc be available. The initial conditions are x(0) = x/ (0) = 0. The relative degree of the system is 2, but La 2 " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001694_03093247v253147-Figure4-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001694_03093247v253147-Figure4-1.png", + "caption": "Fig. 4. Composite nut approximating the theoretical elastic properties of Fig. 3", + "texts": [ + " y(x) = x; y'(x) = 1; y\"(x) = 0 (13) Final combination of equations (1 3) with equation (6) yields B + Cx A + Bx B + Cx A + Bx m' + -m+y-- - 0 where h 2 2a D tan B tan (B - 4)(1 + v) A = - + lrD3 tan 8 tan (8 - d) T 4a A, ~ V L D ~ tan B 2aAn X D L ~ c = -- a An B = A y = A A, By indicating with mo the ratio between bolt and nut Young's moduli at the unloaded face of the nut, the general solution of equation (14) is Figure 3 shows the axial trend of m(x), given by equation (15), for a standard I S 0 M30 screw coupled with a regular nut in the assumption of the same elastic modulus at the free face of the nut (m, = 1). If bolt and nut Young's moduli followed exactly this distribution, the thread load would be uniform along the engagement, as required by equations (1 3). A realistic approximation to this ideal aim could be represented by a homogeneous steel bolt engaging with a 149 at WEST VIRGINA UNIV on June 23, 2015sdj.sagepub.comDownloaded from layered nut, the layers being progressively softer moving from the free face to the loaded face of the nut (Fig. 4). In this way, a thread load concentration factor lower than that produced by the engagement of the bolt with a single nut as soft as the lowest one and as high as the whole stack is to be expected. 5 CONCLUSIONS A popular equation describing the thread load distribution in standard nut-bolt connections has been extended to incorporate the case of threaded elements with axially variable modulus of elasticity. The model has been first applied to inspect the engagement between elasticallyhomogeneous components" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002190_70.850647-Figure8-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002190_70.850647-Figure8-1.png", + "caption": "Fig. 8. (a) Isometric and (b) top views of the simulation.", + "texts": [ + " 7(a) and (b) shows the normalized trajectories for the path in the collision case and in the obstacle-avoidance case. Notice that for the evasion case, curves 2, 3, and 5 take off faster than in the collision case, and curve 6 takes off slower than in the collision case, while curve 1 remains almost the same for both cases. In the case of Fig. 7(b), all the parameters were modulated according to the requirements to evade the obstacle. A robot arm simulator, which was developed within this project [18], performed the PPO process successfully. Fig. 8(a) and (b) shows the simulation of the paths in each case and in different views. Note that the continuous line shows the path in the case of evasion and the discontinuous line shows the path in the case of collision in all these figures. A method to generate trajectories, of hyperbolic type, for manipulators of any number of degrees of freedom has been successfully developed. With a given set of Cartesian coordinates that define the initial and the final position of a PPO, a normalized curve is scaled throughout the entire range, for all the joint variables" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000051_s0091-679x(08)61736-7-Figure22-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000051_s0091-679x(08)61736-7-Figure22-1.png", + "caption": "Fig. 22 the first few seconds of a dynamic reaction. Schematic of \u201cstop-flow\u201d device which allows kinetic measurements to be made during", + "texts": [ + " Recently, we have developed a system which markedly reduces the temporal reference frame so that this can be varied within the interval 1 to 20 sec (Watson et m l . , 1988; Watson, 1992). The technique employs computer-controlled precision drive syringe pumps, one for substrate, the other for cells, together with variable length tubing between a mixing chamber and analysis point, selected by an array of zero dead-space pinch valves. The different tube lengths at a given pump flow rate give different times between mixing and analysis. A schematic of the system is shown in Fig. 22. Calibration was effected using two sets of microbeads with different fluorescence intensities (Polysciences, Inc., Warrington, PA). Pumps 1 and 2 were filled with beads of the higher and lower intensity, respectively. The concentrations were adjusted to give a flow rate of 250 beads per second with both pumps running at 100 p1 min-\u2019. The concentration of beads in pump 2 (lower intensity) was about 1.4 times greater than that in pump 1. Both pumps were activated before data collection in order to fill the input pipes to the mixing chamber, and they were then stopped", + "6, where T is time, F is the average flow rate of the two pumps, and L is tube length. The expression 0.528 - (0.283 x fi), derived from Fig. 25, is equivalent to the slope, rn, in the equation Y = mX + C , and L + 24.73 is equivalent to X, derived from Fig. 24. The constant C , 0.6 sec, can be dropped when both pumps are started before data collection. The method was tested with EMT6 mouse mammary tumor cells in exponential growth adjusted to a concentration of 2 x lo5 cells mi-' and introduced into pump 1 (see Fig. 22). FDA was made up at a concentration of 2 p M and loaded into pump 2. Both pumps were started and the high tension voltage of the green detector was adjusted to record the fluorescein emission histogram with a mean in about channel 600 at pump flow rates of 100 p1 min-' for a tube 31. Enzyme Kinetics 501 length of 40 cm. Recordings were then made at each of the four tube lengths at the same flow rate. From the data shown in Fig. 26 it can be seen that the fluorescence distribution is progressively shifted to higher values with increasing tube length" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003788_20060517-3-fr-2903.00123-Figure8-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003788_20060517-3-fr-2903.00123-Figure8-1.png", + "caption": "Fig. 8. Kinematic chain of the UVMS in the XY plane", + "texts": [ + " 6, as follows \u03b1 = \u2211 R i=1 \u03b1 \u2032 i \u00b5 i \u2211 R i=1 \u00b5 i , (7) where \u03b1\u2032 i are constants and \u00b5 i are definied as in the section 3. Equation 7 is the output of the algorithm, i.e. a hysteresis behaviour as described before. We consider an UVMS with six degrees of freedom, i.e. a manipulator with three dofs mounted in the vehicle as showed in (Antonelli, 2003). This vehicle movement is equivalent to the virtual kinematic chain xy\u03c6 as the vehicle dofs in the XY plane. Moreover, to allow the Davies method application, another virtual kinematic chain x e y e \u03c6 e is employed to represent the end-effector movement as showed in Fig. 8. In this case, the primary vector is composed by the end-effector and vehicles magnitudes, \u03a8 p = [ x\u0307 e y\u0307 e \u03c6\u0307 e x\u0307 y\u0307 \u03c6\u0307 ] , while the secondary vector is the manipulator dofs, \u03a8 s = [ q\u03071 q\u03072 q\u03073 ] . This choice of primary and secundary magnitudes results in the follow determinant of the secundary matrix, det[N s ] =a1a2sin(\u03b82), where a1 and a2 are the first and the second links lenghts of the manipulator and \u03b82 is the angle of the second joint with respect to the first link. It can be easily noted that the maximum value of det[N s ] happens when sin(\u03b82) = 1" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002927_s0967-0661(02)00171-5-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002927_s0967-0661(02)00171-5-Figure1-1.png", + "caption": "Fig. 1. Construction of high-speed magnetic actuator.", + "texts": [ + " Difficulties in designing the controller stem from magnetic force properties, such as severe instability and nonlinearity, which prevent push\u2013pull type actuation. To overcome such difficulties with sufficient robustness, an adaptive control scheme has been developed, which varies the control force depending on the initial condition at which the controller is activated. r 2002 Elsevier Science Ltd. All rights reserved. Keywords: Robust control; Adaptive control; Magnetic actuator; Sliding mode control; Magnetic levitation; Engine control This paper describes a motion control method for a linear actuator like that shown in Fig. 1. This type of actuator can be used, for example, for in-take/exhaust valve actuation of internal combustion engines (Yan et al., 2000) because it is designed to enable periodic piston motion of an actuated object at high speed with travel of 7\u20139mm within 3\u20134ms through the use of springs. The armature is stopped and held by an electromagnet for a certain duration (Fig. 2) at the end of its travel. The moving parts of the actuator are suspended by two springs that facilitate periodic motion at a certain resonance frequency" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001730_s0039-9140(01)00577-x-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001730_s0039-9140(01)00577-x-Figure1-1.png", + "caption": "Fig. 1. Scheme of the flow-cell: (A) 2\u00d71\u00d71.5 cm3 Teflon block; (B) carrier or sample solution inlet; (C) polypropylene pipet tip in which the carbon fibre microelectrode is inserted; (D) Ag/AgCl reference electrode; (E) steel flow outlet tubeauxiliary electrode.", + "texts": [ + " Cylindrical microelectrodes prepared from single carbon fibers (Union Carbide Corp., Danbury, CT, 8 m o.d.) and pretreated as described previously [28] were used as working electrodes. The reference electrode was a BAS MF-2063 Ag/AgCl electrode. The flow injection arrangement consisted of a Gilson Minipuls-3 peristaltic pump and an Omnifit 1106 valve with variable injection volumes. The flow-cell consisted of a home made Teflon block provided with a 2-mm diameter flow channel where the microelectrode is inserted (Fig. 1). This design allows the flow solution passing through the whole active surface of the microelectrode, thus assuring the smallest possible dead volume. The Ag/AgCl reference electrode was also inserted into the flow cell, and the steel flow outlet tube was used as the auxiliary electrode. 2.2. Reagents and solutions Stock 1.0\u00d710\u22123 mol l\u22121 methylthiouracil (MTU) (Aldrich, 95%) solutions were prepared in methanol (SDS, HPLC, 0.02% water max). Working solutions were prepared from these by suitable dilution with methanol", + "05 mol l\u22121 TBAP flowing at a flow rate of 1.5 ml min\u22121. Methanol was selected as the flowing solvent to be used because of the high solubility of the above mentioned antithyroid drugs in such solvent, and it has been used as the extractant of such drugs from animal feed samples [21]. Furthermore, methanol shows additional advantages for its use as the carrier solution with amperometric detection, such as a wide accessible potential range and a cleaning effect of the electrode surface. Using the flow-cell depicted in Fig. 1, true hydrodynamic voltammograms were recorded for MTU, TU and PTU by continuously flowing a 5.0\u00d710\u22125 mol l\u22121 solution of each drug in methanol, containing 0.05 mol l\u22121 TBAP as the tion contribution to the current could be possible if MTU was charged in the organic solvent, absorption spectra of methanolic MTU solutions containing different amounts of NaOH were registered in the 190\u2013450 nm wavelength range. These spectra (Fig. 3) were compared with data reported in the literature [22]. According with these data, the band observed at 276 nm corresponded to the neutral form of MTU", + " The design of this flow-cell, in which the solution flow passes perpendicularly through the microelectrode, permitted only the use of 4- mm length microelectrodes. Under the same experimental conditions, i.e. employing 4-mm length CFMEs, MTU amperometric responses in methanol showed both a higher ip and a lower residence time and peak width when working with the new cell design. This can be attributed to a more efficient contact of the analyte solution with the electrode surface in this case. The peak-to-peak background noise was evaluated working with the flow-cell depicted in Fig. 1, and with CFMEs of different length in the 4\u201310 mm range. Although similar studies with disk microelectrodes showed an increase in the noise level as the electrode radius was increased [12], in our case the background noise remained practically independent on the CFME length. Therefore, the signal-to-noise (S/N) ratio increased with the microelectrode length, and, for example, an MTU concentration of 3.3\u00d710\u22128 mol l\u22121 would correspond to a S/N ratio of 2:1 with a 8-mm length CFME. 3.2. Repeatability of the amperometric measurements One of the most important practical features of the use of CFMEs as amperometric detectors under flowing conditions is their ability of yielding reproducible measurements with no need of electrode pretreatment" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002490_s1474-6670(17)60988-1-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002490_s1474-6670(17)60988-1-Figure1-1.png", + "caption": "Fig. 1. A manipulator with two links.", + "texts": [ + " Further we present some results for kinematic optimal cont rol of two types of Soviet industrial ro bots. These optimal regimes were realized experimentally. The obtained theoretical and experimental results show that optima regimes require 20 - 60% less time that \"natural\" non-optimal ones. The paper is based on the results obtaine\u00b7 in the Department of Mechanics of Control led Systems, Institute for Problems of Mechanics, USSR Academy of Sciences. OPTIMAL MOTIONS OF AN ANTHROR:lMORPHIC MANIPULATOR A manipulator with two rigid links is con sidered (see Fig. 1). Links have equal lengths and are connected with each other by means of the joint O2 \u2022 The joint 01 connects one of the links with a base; a gripper with a load is at the end of th other link. The axes of both cylindrical joints are orthogonal to the plane of the manipilator. The manipilator is controlle, by two torques applied to the axes of the joints. The mass of the manipulator is as sumed negligible compared to the mass of the load. The linear dimension of the lo ad is much less that the length of the links, therefore the load is regarded as a mass point. The motion of the load is described by differential equations (Bo lotnik,and Kaplunov, 1982): ntt = Rx + Ml y I .p 2, ntj\" = Ry - Ml xl P 2 R = r (Ml -2M2 )1 [2 2(4L2_ P 2)] 1/2 \" (1) 2 2 2 .p =x +y r = sign If ' Here m is a mass of the load, x, y are Cartesian coordinates of the load; L is a lenth of the links of the manipulator; ~ is an angle between the links; Ml , M2 are'control torques applied to the joints 01 , O2 respectively, see Fig. 1. The geometry of the system implies two pos sible configuratio'ns of the manipulator for any position of the load. These con figurations differ from each other by the sign of the angle tp \u2022 The value r = 1 (see Eq. (1\u00bb corresponds to the configu ration of the type f+ shown by the solid line in Fig. 1, and r = - 1 corres ponds to the configuration of the type r(dash line). The type of the configura~ion is constant during the motion if the load does not ~each the joint 01 or the circle ~ + y2 = 4L2 whicfi is the bound of the operation area of the manipulator. Problem 1. Let at the initial time moment t = 0 the load rest in the state x(O) = xo ' y(O) = Yo' i(O) = y(O) = 0, (2) It is required to determine control tor ques Ml = M~(t), M? = M~(t) and the type of the t!onfigur~tion (1\"+ or r _) which provide time-o~timal transfer of the load from the state (2) to the given state x( T) = Xl' yeT) = Yl' x(T) = yeT) = 0, 0) 2 2 2 o < Xl + Yl <" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002210_s0022-5096(99)00047-2-Figure7-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002210_s0022-5096(99)00047-2-Figure7-1.png", + "caption": "Fig. 7. The method how the energy of the singularity is measured. Path1 and Path2 are the `crumpling' of a perfectly \u00afat plate and the reloading of the already crumpled plate. The area between the thick line (path2) and the thin line (path1) is the energy of the singularity.", + "texts": [ + " But, if the plate is bent until the internal face feels compression and the external face feels stretching, the plate will have deformed plastically and will not recover its shape. In force measurements, the load necessary to bend the plate to the same point, in subsequent tests will be lower because the plate has been `weakened'. In order to measure the singularity energy, we \u00aerst measure the force required to reach a deformation z , then release the plate, and measure the force required to reload the plate to the deformation z . The area between the loading lines, shown in Fig. 7, is the singularity energy, corresponding to the energy dissipated while creating the scar region. If the plate is loaded a third time, the force follows the loading line corresponding to the second loading. If z is the displacement, then the singularity energy is given by: Esing ' z 0 F1dz path1 \u00ff z 0 F2dz path2 6 Where path1 and path2, correspond to the \u00aerst and second loads respectively. F1 and F2 are the force on a non-deformed plate and the force on an already deformed plate respectively. For subsequent loads after the initial loading, the force always follows path2 as long as the deformation does not exceed z " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001944_abab:102-103:1-6:471-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001944_abab:102-103:1-6:471-Figure1-1.png", + "caption": "Fig. 1. A schematic diagram of the dissolved oxygen electrode set up with the sandwich enzyme membrane system.", + "texts": [ + " In order to measure the immobilized enzyme activity in the presence and absence of various concentrations of GdmCl, an amperometric prin- Applied Biochemistry and Biotechnology Vol. 102\u2013103, 2002 ciple based Clark-type dissolved oxygen electrode along with a dissolved oxygen meter (EDT, UK) was used. A composite membrane system comprising the enzyme layer was held in a sandwiched form between a Teflon membrane and a cellophane membrane, secured tightly with an O ring on the electrode surface (see Fig. 1). The activity of the immobilized enzyme was measured by immersing the enzyme-sensing element in a glass sample cell of 25 mL containing 5 mL buffer, kept agitated continuously with air bubbled through a portable air pump in order to eliminate dissolved oxygen limitations. After initial bubbling of air for saturation, the dissolved oxygen meter was set at 100%. A known, relatively high concentration of glucose solution (chosen here as 10%) was now injected and the decrease in percentage dissolved oxygen at the end of 3 min (time taken to reach steady state) was monitored" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003159_12.470666-Figure7-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003159_12.470666-Figure7-1.png", + "caption": "Fig. 7 Bearings for automatic watches Fig. 8 Axes and crowns of mechanical watches welded with the SHADOW technique.", + "texts": [ + " Only the feasibility of the welding was to be proven. The line energy applied for the different applications measures from 0.6 to 1.38 J/mm. The energy needed to weld the two parts strongly depends on the diameter and the volume of the parts. The heat capacity and the heat losses into the surrounding material influence the required laser power. SHADOW technique creates smooth weld seams with high strength as shown in the following pictures: courtesy of In an automatic watch the oscillating mass is mounted on a precision bearing as shown in Fig. 7. The inner ring is adjusted precisely to guarantee a movement free of clearance. After the adjustment process, the two rings are fixed to each other. The conventional process of crimping was now replaced by laser beam micro welding with SHADOW. The diameter of the weld seam is 1.7 mm. The welding speed is 16 m/min. The strength of the joint is increased compared to bearings which are conventionally crimped. Most of the applications in watch industry are axis/wheel combinations. Since the parts are often stamped out of sheet metal the axis has to be added and joined" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001384_978-94-009-1718-7_18-Figure2-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001384_978-94-009-1718-7_18-Figure2-1.png", + "caption": "Figure 2. The planar manipulator and the location of segments with near-maximum path manipulability.", + "texts": [ + " Therefore thej th joint cQ-ordinate can be calculated from the values of the corresponding (m/k)-bit using the following mapping: qj = q~ + [Bitvalue 12m/k 1 q~ (1) where Bitvalue is the decimal number corresponding to string of thej th joint. The efficiency of the proposed method for estimating the location of a straight line path with the higher path manipulability has been demonstrated through simulated experiments using a planar 3-DOF manipulator and a PUMA robot. It is assumed that the tip of the hand of the planar 3-DOF manipulator shown in Fig. 2 has to move along the straight line segment AB, with the axis of its end-effector perpendicular to the segment. It is a prerequisite to locate the segment AB in the location with the higher path manipulability. 185 The proposed algorithm has been tested in a number of experiments with straight line segments of various lengths. Fig. 3 illustrates the convergence of the GA with the following control parameters:Population size = 50, Maximum number of generations = 50, Probability of crossover = 0.4, Probability of mutation = 0.0 IS. The best-of generation value and the generation average are presented for one run of the GA solution. As it is shown in this figure the convergence criterion proposed by De long (1975) according to which the (average fitness/best fitness) ratio has to be greater than 95% is satisfied in this case.The locations of the estimated near-maximum manipulability in some of the executed runs are shown in Fig. 2 with dashed line. The algorithm estimated the near-maximum path manipulability at the location of the straight line path where the starting point has the following co-ordinates (3.7, 2.3) and the angle between the segment and the x axis is -1.43 rad. This path was followed by the tip of the hand using the Resolved Motion Rate Control (RMRC) introduced by Whitney (1972). The manipulability and the normalised Manipulator Velocity Ratio (MVR) introduced by Dubey and Luh (1986) for each step of RMRC algorithm are plotted in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001291_oca.4660120407-FigureI-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001291_oca.4660120407-FigureI-1.png", + "caption": "Figure I . Minimum distance between collidable link and obstacle", + "texts": [], + "surrounding_texts": [ + "which is away from the obstacle. Baillieul' introduced the concept of extended Jacobian to utilize redundancy and optimize instantaneously some distance function to avoid obstacles. Yoshikawa* defined a collision-free posture in joint space and utilized redundancy to approach this posture while tracking the path. Nakamura and Hanafusa' presented a global optimization scheme by using joint space and applying Pontryagin's theorems. This approach results in an extensive calculation of an n x n configurationally varying matrix at each time step." + ] + }, + { + "image_filename": "designv11_65_0003531_j.mechmachtheory.2004.07.003-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003531_j.mechmachtheory.2004.07.003-Figure1-1.png", + "caption": "Fig. 1. Structure of the VGPM.", + "texts": [ + " The goal of this paper is to present a method for solving the inverse kinematics problem of the VGPM. For the given outlet area and the deflection angle of the VGB, the inverse kinematics problem of the VGPM can be divided into two steps: determining the poses of the top platform and calculating the inverse kinematics problem of the Stewart platform. The second step is straightforward, and it will not be presented in this paper. This work only focuses on the first step, i.e. for the given outlet area and the deflection angle of the VGB determining the poses of the top platform. As shown in Fig. 1, the VGPM consists of two parts, the driving Stewart platform mechanism and a number of spatial RSRR kinematic chains (AiBiCiDi). The RSRR kinematic chain connects with the base frame and the top platform through the revolute pair at point Ai and the revolute pair at point Di, respectively. The rigid plate with a spatially shaped inner enveloping surface X is fixed to the SR links (BiCi) in the RSRR chain. By a number of surfaces X, the variable geometry body (VGB) will be enveloped. The variation of the pose of the top platform will make the RSRR kinematic chains change their spatial positions", + " Thus one has f2 \u00bc S2aSG 2 Sx C2aSG 2 Sy C2aSG 2 Sz \u00bc 0 \u00f04\u00de where C2aS \u00bc cos2aS, S 2aS \u00bc sin2aS: There are six parameters to identify the pose (position and orientation) of the top platform of the Stewart platform. In order to improve the capacity of the VGPM to bear the side-load and to prevent the interferences between surfaces X, the center of the top platform should be on the zaxis. In this situation, the center coordinate of the top platform is E (0,0, l0), where l0 is the distance from the top platform center E to the base frame center O, as shown in Fig. 1. Moreover, in order to prevent collisions between surfaces X, the rotational degree of freedom of the top platform about z-axis is restricted. In this case, the two parameters h and n, as shown in Fig. 1, can be used to represent the orientation of the platform. Here h is the angle between the normal vector ~n of the top platform and z-axis, and n is the angle between the projection of the normal vector ~n on the xOy plane and x-axis. Since the VGB is symmetrical about its deflection plane, we know that n = b. Therefore, to the inverse kinematic problem of the VGPM, the three unknown parameters describing the pose of the top platform are fl0; h; ng \u00f05\u00de Let di denote the rotation angle of link AiBi at the revolute joint Ai, xi denote the rotation angle of link BiCi at the revolute joint Ci, and mi denote the rotation angle of link CiDi at the revolute joint Di" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003764_rob.20045-Figure2-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003764_rob.20045-Figure2-1.png", + "caption": "Figure 2. Planar 2-link manipulator.", + "texts": [ + " A large value for wo ensures that there is no collision with obstacles, whereas a small value yields faster convergence of the optimization. The tolerances o and c are values of Jo 0 and Jc 0 , respectively, used to accept a minimumoverload trajectory as an obstacle-free minimum-time motion. Thus, the convergence criteria for obstaclefree minimum-time motions are Jo 0 o , Jc 0 c . (39) These convergence criteria are tested at the end of every minimum-overload search. In the numerical simulations, 10 7 and 10 5 are used as o and c , respectively. The first example is a simple planar 2-link arm shown in Figure 2, where two joints are revolute pairs around their z-axes. The masses, lengths and crosssections of the first and second links are 25 and 15 kg, 0.8 and 0.6 m, and (0.15, 0.15) and (0.12, 0.12) m, respectively. Gravity is acting in Y0 direction. c 530 90 T Nm, c 6 6 T rad/s. The dimensions (lx ,ly ,lz) and the geometric center of one hexahedral obstacle are (0.4, 0.6, 0.4) and (1.2, 0, 0) m in base co- ordinates. The manipulator moves from ( 30\u00b0 , 30\u00b0) to (30\u00b0 ,30\u00b0) in joint space. The velocities and the accelerations at two end points are zero" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002155_mssp.2001.1461-Figure3-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002155_mssp.2001.1461-Figure3-1.png", + "caption": "Figure 3. Elastic micro-contact model. The global contact patch consists of a certain number of statistically distributed on average similar contacts. The classical Hertzian theory is assumed to hold at each single microcontact (see details). ceff is now an effective half-contact width, which, in general, is larger than the corresponding measure of a total smooth contact according to the Hertzian theory.", + "texts": [ + " Therefore, in the present investigation the input roughness profile results from a normally worn steel surface, picked-up with a stylus radius of 5 mm, only filtered with a four-pole-high-pass filter with a cut-off, khp, defined above. The used input profile compared with the original one, can be seen in Fig. 1, represented in the wavenumber range. The present analysis is based on the well-known basic work of [11, 15\u201318, see also 12, 14]. In these references the Hertzian static contact theory of smooth elastic surfaces [19], Fig. 2, is extended to the statistically distributed rough elastic contact under certain conditions, see Fig. 3. It is not the purpose of the paper either to repeat the basic theories of these authors or to prove basic mathematical relationships, this is well documented in the mentioned references. Only those basic definitions should be repeated which are important in understanding the author. The preconditions in all the approaches are, see again Fig. 3: (i) On an average contacting asperities are identical, they have spherical capped tips with a mean radius, ri, independent of their heights. (ii) Asperities are mechanically independent, that means no interactions between them. (iii) Single mean asperities deform elastically corresponding to the Hertzian contact theory in the following form: single contact area: Ai \u00bc pridi \u00f0m2\u00de \u00f02\u00de and single elastic force Fi \u00bc 4 3 E 0r 1=2 i d3=2i \u00f0N\u00de \u00f03\u00de where E0 is the resulting elasticity (Young\u2019s) modulus 1 E0 \u00bc 1 m21 E1 \u00fe 1 m22 E2 \u00f0m2=N\u00de \u00f04\u00de consist of the single moduli E1,2 and the single Poisson ratios m1;2, respectively, the subscripts 1,2 refer to the two corresponding contacting bodies, di is the elastic compression of one single mean contact", + " Now, under the assumption that z(x) is a Gaussian random variable, the three necessary input parameter can be calculated in the following way [24]: Z \u00bc m4 m2 32:65 \u00f08\u00de ss \u00bc \u00f01 0:8968=a\u00de1=2m1=2 0 \u00f09\u00de ri \u00bc 0:375 p m4 1=2 : \u00f010\u00de The expressions include the fact that the model originally is based only on the summits, whereby available roughness data and, therefore, the moments normally are generated from a complete profile. The parameter, a, sometimes called \u2018bandwidth parameter\u2019, is defined as a \u00bc m0m4=m 2 2: \u00f011\u00de This parameter indicates that, in general, the summits have a lower variance than the surface profile as a whole. According to Fig. 3, if the separation of the nominal surface at the position of a particular asperity is u, there will be a contact at that asperity if the height, z, is larger than u. The probability of this event is determined by the probability density function, F\u00f0z\u00de, of the distribution of asperity heights. The expected value of a summit contact area, %Ai, is thus %Ai \u00bc Z 1 u pri\u00f0z u\u00deF\u00f0z\u00de dz \u00f012\u00de the expected force, %Fi, is %Fi \u00bc Z 1 u 4 3 E0r 1=2 i \u00f0z u\u00de3=2F\u00f0z\u00de dz \u00f013\u00de with z u \u00bc d (compare equations (2) and (3))" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002671_iros.1997.649067-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002671_iros.1997.649067-Figure1-1.png", + "caption": "Figure 1: Regrasp primitives (a) Rotation (b) Pivoting", + "texts": [], + "surrounding_texts": [ + "1 Introduction In order to bring the held object into a specific desired orientation without being released or intermediate placing and repicking, orientation planning on the part of the hand is required. One strategy to perform reorientation is through execution of primitives. Primitives, in original adapted under inspiration from human hand, can be defined as routine procedure movements (generated by fingers), marked under a name. Use of primitives simplifies the complex three-dimensional reorientation tasks by disintegrating it into a sequence of simple and usually two dimensional motions.\nPreviously, studies on primitives based reorientation has' widely been conducted. Some of the identified (both of regrasping and non-regrasping type) primitives are sliding and rolling [3, 71, rotation [4], pivoting 181, finger-tracking [5, 6]? fumbling 191. Others are found io [lo, 11][13]-[14]. Okada [l] and Fearing [2] worked on the twirling reorientation of a bar by\nthree-fingered hand. Rus [5] presented reorientation of a polygon by sliding one finger (finger tracking) along the edge of a polygon. Gupta[6] proposed a planner enhancing the strategy adapted by Rus. Qinkle 171 discussed non-primitive dexterous manipulation by hand. Paetsch and Wichert[S] adapted fumbling strategy to solve peg in a hole insertion task in view to promote use of such useful strategies for different tasks. Michelman and Allen[lO] built a set of primitive functions to perform fingertip and precision manipulation tasks. Speeter[ll] has defined primitives for Utah/MIT hand based on joint motions.\nIn OUT proposed planner two primitives rotatzon[4] and pivoting[8] are adapted.l This work differs from the previous related work in the sense that the proposed planner takes into consideration the geometrical shape of the object. This work, does not formulate reorientation plan for each primitive, which has already been discussed in [4] and [8], but is concerned with generating the goal-bound sequence of primitives.\nRotation is a primitive[4], in which the object is rotated in a plane more or less parallel to the plane of out-stretched fingers' tips. Fig.l(a) illustrates the rotation with sense of rotation and its axis shown by arrows. During rotation, when fingers reach the boundary limits of their work spaces, they are alternatively lifted and repositioned back in the opposite direction of rotation. This enables the fingers to rotate the object at large angles despite of their limited range of work spaces.\nIn pivoting primitive[8], the object is pivoted, as illustrated in Fig.l(b). Two of the fhgers form the axis of pivoting by grasping the object at its opposite sides. Another finger then pivots the object by pushing on its side. By pivoting too, the object can be rotated at large angles by repositioning the finger(s).\nRotation and pivoting primitives are adapted in this work because of their certain advantageous feature5 as described below.\nl1n italic form, rotation represents the name of the primitive; while in non-italic form the word rotation simply means the w+ating of an object.\nProc. IROS 97 0-7803-4119-8/97/$1001997 IEEE", + "(a). Simplicity. Both of these are simple and two dimensional primitives. However, their sequential combination can achieve three dimensional orientation. (b). Adapatability. Most of the real world objects do have geometric features suitable to adapt the execution of rotation and pivoting. Rotation requires a set of peripheral sides at which the hand must be able to grasp the object and pivoting requires two opposite sides large enough to hold the axis for pivoting. Most of the objects readily offer these features.\nThis work is not limited to polygonal objects only. Practically these primitives are also applicable to objects with round contours zs the conditions defined above remains valid. (In round contoured objects the set of sides for Rotation become infinite, and the opposite sides for pivoting becomes infinitesimally small.)\nThis paper is organized as follows. Section 2 describes the planner with its four main steps in subsections. An example is simultaneously presented alongwith the description of the planner. Section 2 is followed by Discussion and Conclusion.\n2 The Planner Given the object, the planner generates a sequence of rotation and pivoting primitives in order to attain the desired goal orientation from certain given initial grasping orientation. We make the following assump t,iOIls.\nFour-fingered hand is preferred mer three-fingered hand particularly for rotation primitive, as at least three fingers ensure an stable grasp, when the fourth one is lifted for repositioning.\nOrientation of the object with reference to the hand reference frame is known.\nAll fingertips are capable of reaching the same plane in space, which is parallel to the xy-plane of the hand reference frame {H}zyz (shown in Fig.2).\nMost of existing robotic hands satisfsi assumption (3). The following is the outline of the planner. (1)- The first step is to locate available axes on the object. Available axis implies an axis on the object about which execution of rotation and/or pivoting primitives is feasible. Available axes me defined with respect to the body reference frame (not hand reference frame). (2)- The 2nd step is to construct the states network using the data of available axes. The states network is composed of the orientation states of the object defined w.r.t. {M}zyz. The essential characteristics of the orientation states is that at each state at least one of the rotation and pivoting primitives for among all the available rotation and pivoting axes on the object is feasible. (3)- The final step is that, given the initial and goal orientation states, searching the routes between the two states. The routes are compwed of the sequence of primitives.\nConstruction of states network do not require initial and goal orientations, but only the data of given object, therefore steps (1) and (2) can be executed off-line.\nThis paper i s mainly concerned with steps (2) and (3) but not with step (l), that is, this paper does not give complete conditions for locating available rotation and pivoting axes.\nAs mentioned in Section 1, rotation requires a set of peripheral side facm and pivoting requires two opposite side faces. The rotation axis is located perpendicular to the normals of the side faces if they are parallel to each other, and the pzvoting axis is located perpendicular to the two opposite faces if they are exactly parallel. However, these requirements are not so rigid as far a", + "287\nthe hand is able to grasp the object at side faces not exactly parallel. Moreover, any edge or portion of the object must not hamper the fingers movements during the execution of primitives, and the distance between the grasped sides must be within the dimensions of fingers' workspace.\nFor the planner, initially we merely declare the rotation and pivoting axes as available without checking the geometrical feasibility. Our previous papers [4] [8] partially check the geometrical feasibility for locating rotation and pivoting axes. However, it is not practical to state the complete conditions. A practical way to check the geometrical feasibility of the declared available axes is to check them during on-line execution of the primitives. We follow this approach, and discard an available axis whenever found geometrically infeasible for exection. Thus locating more axes than the real feasible ones does not cause any significant problem.\n2.1 Construction of states network for the given object\nAvailable axes become feasible (orientationally) for execution when the criteria provided by Assumptions (4) and (5) are satisfied. As described earlier, the states network database is composed of all the orientation states of the object at which at lemt one of the available rotation and pivoting axes is feasible for execution of primitive about it.\nLet { R j } , j = 1 ...... U and {P'}, j = 1 ...... II be the available rotation and pivoting axes. Then w = U 4- v represents the total number of available axes on the object. Let {Si}, i = l . . . . . .n be theorientationstateset of the object when at each Si at least one axis in {EJ} or {Pj} is feasible. Value of n depends upon the value of W. The presented algorithm discovers the value of n, while constructing the set {Si). Each orientation state s k is represented by a 3 x 3 rotation matrix defined relative to hand reference frame { H j Z V 2 .\nTo start constructing the set {Si), the object is simulated to set at a state SI at which at least one pivoting axis ap (aP E {Pj} at SI) is feasible (i.e.> lying in the xy-plane). Therefore SI becomes first subset state to initialize {Si}. For searching more states, the planner simulates to rotate the object alternatively about all feasible pivoting axes available at SI, and mark an orientation state as a new one whenever an axis a k ( a k E { R j ) , { P j } ) is raised to feasibility (represented by a',). The question arises that why only the feasible pivoting axes at an state s k are considered for searching its offspring states and why not the feasible rotation axes. This will be explained later in Proposition 2.\nThe above process is represented by the following equation of rotation matrix,\n[@(a,, e)]ak = ;ak # a p (1)\na k E (Rj 1 ; then for feasibility ai must be parallel\nIf Uk E (6) ; then for feasibility a& must lie in ?q-\nAxes ap, ak and u i are shown in Fig.3.\nto the *z-axis of {H)zy2 (assumption (4 ) ) .\nplane of {H}zv2 (assumption (5)).\nWhen ab can be raised to feasibility, then the angle 6 is derived from the above equation. Fig.3(a) & (b) is given to illustrate that how angle 6 is calculated for a k being a rotation and pivoting axis respectively.\nAfter calculating 6, if available, the object is rotated at 0 about axis up to reach some other orientation state s k . If SI, does not pre-exist in (Sz}, then it is added in IS,}.\nThe following proposition is useful to check an available rotation axis if it could be raised to feasibility by rotating it about currently feasible pivoting axis.\nProposition 1: A necessary and sufscient condition in order to make a vector a k E ( R j } parallel to the fz-axis by rotating it about a pivoting axis a,, is that it is orthogonal with ap.\nProof: A x i s ap being a pivoting axis at feasibility is lying in q-plane, thus is orthogonal to f z-axis or a&, (or in turn to a h ) . 0\nNoticing that a feasible rotation axis is always parallel to the z-axis of {H}Zyz, a proposition is made in this regard.\nProposition 2: An axis a k E { R J ) or { P J } , ( a k not parallel to z-axis and not lying in xy-plane); if rotated about z-axis, cannot be made parallel to saxis and cannot be made to lie in xy-plane. Hence cannot be raised to feasibility.\nProof: Immediate. 0 The above proposition suggests that if the object is rotated about a feasible mtation axis (being parallel to z-axis), the feasibility/infeasibility state of all other axes will remain nneffected, i.e., the feasible axes will remain feasible and infexsible axes will remain infeasible. That is why the feasible rotation axes at an state are not considered for searching new ofkipring states. This avoids creation of infinite number of states of orientations at which the object can be rotated only about the feasible rotation axis ( R k ) (= z-axis). Instead the" + ] + }, + { + "image_filename": "designv11_65_0003038_s0094-5765(01)00002-9-Figure4-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003038_s0094-5765(01)00002-9-Figure4-1.png", + "caption": "Fig. 4. SimpliAed system showing a space platform supporting one-module manipulator.", + "texts": [ + " The slewing maneuver at any joint can also be spatial. (e) The damping is accounted for through Rayleigh\u2019s dissipation function. (f) The governing equations account for gravity gradient eIects, shift in center of mass and change in inertia due to maneuvers. 2.1.1. Reference frames The mobile manipulator with an arbitrary number of modules (bodies) is supported by a platform orbiting around Earth. To help appreciate notation for such a general system, consider a simple case of the manipulator with one module (Fig. 4). Note, body 1 refers to the orbiting platform and body 2 represents module 1 of the manipulator, supported by the mobile base. The inertial frame F0 is located at the center of Earth. The x0; y0-axes establish the orbital plane and z0 represents the orbit normal. The position and attitude of each body are described by the body-Axed frames F1 and F2. The frame F1 has its origin at the center of mass of the platform while F2 is attached to body 2 at the joint with the mobile base. As the platform and manipulator modules are considered to be beam-type structures, the x1 is taken along the beam axis, y1 is perpendicular to x1 in the orbital plane, and z1 completes the orthogonal triad according to the right hand rule", + " With these introductory comments, consider the chain-type manipulator with N bodies as shown in Fig. 5. As before, body 1 represents the platform, while the remaining bodies (2 to N ) correspond to the manipulator modules. Thus, the second body represents the Arst module of the manipulator, while the body N corresponds to the (N th\u22121) module. Note, the lengths of bodies 2 to N can vary with time. Moreover, each body is free to rotate and translate with respect to its neighbors. As in the case of one module system (Fig. 4), the xi axis is along the length li of the body i; yi is perpendicular to xi in the orbital plane; while zi completes the orthogonal triad. It may be emphasized that for both the platform and manipulator modules, the stiIness, damping and inertia properties can vary along xi. Di represents the position vector to the frame Fi, attached to the body i at the joint between bodies i \u2212 1 and i, with respect to the frame F0. As explained earlier, li\u22121 + di + fi would represent position of the base, free to traverse body i \u2212 1, if it were present" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001694_03093247v253147-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001694_03093247v253147-Figure1-1.png", + "caption": "Fig. 1 . Basic bolt (to left) and nut (to right)deformations affecting the thread load distribution", + "texts": [], + "surrounding_texts": [ + "EFFECT OF NUT COMPLIANCE ON SCREW THREAD LOAD DISTRIBUTION\nE. DRAGON1 Department ofMechanics, University oflologna, Italy.\nA well-known theoretical equation for the distribution of contact load along threads in standard nut-bolt connections is generalized to incorporate the case of axially variable Young\u2019s moduli of the threaded elements. The influence of the material selection for homogeneous bolt and nut on the thread load concentration factor is initially addressed. The theoretical possibility of equalising the thread load distribution over the whole engagement by a convenient axial variation of the nut Young\u2019s modulus is then examined.\n1 INTRODUCIlON Threaded connections are probably the most common components in mechanical constructions. According to an early investigation (l)t, fatigue failures of bolts occur, for about 15 per cent of cases, at the fillet between the shank and the bolt head, for about 20 per cent of cases at the thread run-out, and for the remaining 65 per cent of cases, at the cross section where the bolt first engages with the nut. The stress concentration factors associated with the first two kinds of failure are easily reduced by increasing the fillet radius at the base of the head and conveniently decreasing the shank diameter at the beginning of thread, respectively. Regarding the third type of failure, the problem is to smooth out the stress distribution at the root of mating threads which, in ordinary screw-nut connections, can exhibit a peak as high as five times the mean core stress at the loaded face of the nut (21, (3).\nSince the pioneering works of Jaquet (4) and Maduschka (5), a huge number of proposals have appeared in the technical literature. The basic aim is to relieve the first threads of the bolt from the high load needed to recover the pitch difference between the fully stretched bolt and the fully compressed nut. This design objective is mainly sought by either providing variable axial play between unloaded screw and nut threads (6)- (8), or acting on the external shape of the nut (6), (9Hll).\nThe matching of an ordinary bolt with a softer standard nut has also been shown to produce beneficial effects on the fatigue life of the connection (12), (13). This is a consequence of the more uniform load distribution along threads which can be attained. For example, in the engagement between a steel screw and a titanium nut, a reduction in the thread load concentration factor of about 20 per cent with respect to equimodulus connections has been theoretically estimated (14).\nIn this paper, the possibility of levelling even further the load distribution through a suitable axial variation of the nut Young\u2019s modulus is investigated. The MS. of this paper was received at the Institution on 6 October I989 and\nt Rejerences are given in the Appendix.\nJOURNAL OF STRAIN ANALYSIS VOL 25 NO 3 1990 0 IMechE 1990\naccepted for publication on 9 March 1990\n1.1 Notation a A, B, c\nAn\nD Eb, E n\nh L m\nm0 P, 4 P PO S\nSl W\nw m W X\nV\n4\nThread pitch Non-dimensional constants Mean cross-sectional areas of bolt and nut, respectively Effective (mean) diameter of thread Young\u2019s moduli of bolt and nut, respectively Thread deflection factor Length of nut Ratio of bolt Young\u2019s modulus to nut Young\u2019s modulus Value of rn at free face of nut Non-dimensional constants Bolt axial load Full bolt axial load Distance from free face of nut along thread helix at diameter D Value of s at loaded face of nut Axial load on thread per unit length of thread helix Mean value of w Thread load concentration factor Non-dimensional thread length from free face of nut Non-dimensional bolt axial load Non-dimensional constant Half of thread angle Ratio of nut to bolt cross section Relative axial movements between roots of bolt and nut threads Poisson\u2019s ratio Friction angle\n2 THEORETICAL THREAD LOAD DISTRIBUTION\nThe differential equation governing the thread load distribution in screwed connections stems from the compatibility condition of bolt and nut axial deformations (7), (15)\n141\n0309-3247/90/07004147 $02.00 + .05 at WEST VIRGINA UNIV on June 23, 2015sdj.sagepub.comDownloaded from", + "(1) d ds 60 = nD - (6, + 6 2 + 6 3 )\nEquation (1) shows that the pitch difference, 6, (Fig. l(a)), that would take place if the screw were free to stretch and the nut free to shorten, must be compensated by the gradient of three deformation mechanisms: the thread deflection as a cantilever beam, 6, (Fig. l(b)), the axial separation of threads created by bolt shrinkage and nut bulge due to radial pressure between threads, 6, (Fig. l(c)), and the axial recession arising from the radial movements associated with the Poisson\u2019s effect, 6 , (Fig. l(d)). By referring to the symbol list, the above terms can be expressed analytically as follows\n6 2 = D tan B tan (B - 4) {-(\u2019+-) nDz 1\n2a EbAb\n2 6, =\n(3)\n(4)\nwhere a common value, v, of the Poisson\u2019s ratio for bolt and nut materials and the same geometric thread deflection factor, h/2, for both male and female threads have been assumed. Expressions of h as a function of thread geometry and friction angle 4 are provided in (7) and (15), based on the displacement field in a wedge due to concentrated forces and couple at its apex. A check of these theoretical values against the results of a finite element analysis of nut and bolt thread interaction is provided in (16).\nUpon insertion of equations (2H5) into equation (l), the following equation in terms of the dimensionless parameters\nEll m(X) = - E\u201c\nis obtained\nD tanp tan (B - 4x1 + v) (m - 1) {:(I + m) + 2a\nnD3 tan p tan (B - 4)\n+[{5+ 2a\nno3 tan /3 tan (B - 4) 4aAn +\n~ V L D ~ tan f i ( lb I~)]\n1 t 4a (a, + :)Iy.l\nh DtanBtan(B-4Xl + v )\n1.\u2018 -+- y\u2019\n+ 2a\n+ { 2aAn nvLDZ tanp m \u2018 - y ( ~ + ~ ) } y = O nDL2 (6)\nin which the primes imply differentiation with respect to\nEquation (6) can be tackled in two primary ways: (i) once the geometry and the material elastic properties have been stated, the resulting axial load distribution, fix), is derived; from this, the thread load distribution is finally obtained by differentiation, (ii) a convenient axial load distribution is assumed and the equation is solved for the unknown function m(x), characterising elastically bolt and nut materials able to yield the inferred distribution.\nBoth the above direct approach (i) and the inverse one (ii) are considered in the following.\n3 EFFECT OF NUT MATERIAL The simplest situation occurs when both screw and nut are homogeneous so that m(x) = m, = constant. In this case, equation (6) reduces to\nX.\ny\u201d + 2py\u2019 - (42 - p2)y = 0 (7) where\n2{ha(l + m,) .. + D tan 8 tan (B - dX1 + vXmn - 1))\n148 JOURNAL OF STRAIN ANALYSIS VOL 25 NO 3 1990 0 IMechE 1990\nat WEST VIRGINA UNIV on June 23, 2015sdj.sagepub.comDownloaded from", + "(9)\nEquations (8)-(10) are the extension of equations (26), presented in (7), to cover the case of different Young's moduli between bolt and nut.\nWith the aid of the boundary conditions\ny(x = 0) = 0; y(x = 1) = 1 (1 1) equation (7) is easily solved. After differentiation, the thread load concentration factor is found to be\nAs a numerical example, in Fig. 2 a graphical representation of equation (12) is offered for an I S 0 M30 screw engaging with a regular nut. In the calculation, a frictional coefficient of 0.2 was assumed. The diagram clearly indicates the beneficial effect associated with an increase of m,, that is, with the adoption of nuts increasingly softer than the screw. From a practical standpoint, the advantage that can be gained by switching from a standard steel bolt-nut connection (m, = 1) to a steel bolt engaging with an aluminium nut (m, = 3) is a 15 per cent decrease of the maximum thread load. Results of fatigue tests on similar connections have been reported by Wiegand in (12).\nInterestingly, it must be noted that the thread concentration factor can by no means be lowered beyond a limit greater than unity (dashed line), even if the screw were fitted with an infinitely softer nut (m, + a). In other words, a uniform thread load distribution (W = 1) cannot be reached by simply changing the nut (and/or screw) Young's modulus as a whole.\nJOURNAL OF STRAIN ANALYSIS VOL 25 NO 3 1990 6 IMechE 1990\n4 NEARLY-OPTIMUM NUT COMPLIANCE An alternative way to make use of equation (6) is to look at y(x) as a known function and solve the resulting equation in terms of m(x). The solution of this inverse problem provides indications about how to distribute, axially, screw and nut elasticity in order to obtain the desired load distribution. Clearly, a major point in this respect is to find a convenient form of the object function\nFrom the screw strength point of view, the most desirable situation originates in the maximum stresses at the base of bolt threads reaching the same value over the whole engagement. Now, the total thread root stress results from the concurrence of two distinct contributions (17): (i) the cantilever action of the flank load w, and (ii) the notch effect under the bolt axial load P. Since the latter contribution is unavoidably maximum at the bearing face of the nut, where the bolt supports the full axial load, P o , a uniform total stress can be theoretically obtained only when the bending thread reaction is kept smaller at this location than at the free face of the nut. For the sake of simplicity, however, the following development relates to the reference assumption of uniform thread load. This particular request, coupled with the boundary conditions (1 l), results in\nfix).\ny(x) = x; y'(x) = 1; y\"(x) = 0 (13) Final combination of equations (1 3) with equation (6) yields\nB + Cx A + Bx\nB + Cx A + Bx m' + -m+y-- - 0\nwhere\nh 2 2a\nD tan B tan (B - 4)(1 + v) A = - +\nlrD3 tan 8 tan (8 - d) T\n4a A,\n~ V L D ~ tan B 2aAn\nX D L ~ c = -- a An\nB =\nA y = A A,\nBy indicating with mo the ratio between bolt and nut Young's moduli at the unloaded face of the nut, the general solution of equation (14) is\nFigure 3 shows the axial trend of m(x), given by equation (15), for a standard I S 0 M30 screw coupled with a regular nut in the assumption of the same elastic modulus at the free face of the nut (m, = 1). If bolt and nut Young's moduli followed exactly this distribution, the thread load would be uniform along the engagement, as required by equations (1 3).\nA realistic approximation to this ideal aim could be represented by a homogeneous steel bolt engaging with a\n149\nat WEST VIRGINA UNIV on June 23, 2015sdj.sagepub.comDownloaded from" + ] + }, + { + "image_filename": "designv11_65_0002008_12.474436-Figure5-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002008_12.474436-Figure5-1.png", + "caption": "Figure 5. Some HUR-Badger Platform Functions. (Note: Model does not show tracks.)", + "texts": [ + " In response to these limitations a second HUR robot was designed with the rigors of urban operations and a realistic time frame in mind. The vehicle is called the HUR-Badger and is shown in Figures 5 and 6. The Badger consists of two tracked units connected to a common body using rotational joints. The tracked units are sized such that they can be rotated through each other. The individual tracks of each unit do not rotate relative to each other and each unit is skid steered using di erential speed control of individual tracks. Tables 1 and 2 summarize some of the useful con gurations and functions of the Badger of Figure 5 and 6. The Badger is shown capable of snow-shoeing on soft terrain or loose rubble, conventional tracked locomotion in tight, short and long con gurations, caster turning on hard surfaces, a ratchet maneuver up stairs and steep rubble and a wedge con guration with counter rotating tracks that pull it under wire and and allow it to bury into rubble. The badger is further capable of crossing gaps, climbing linear obstacles, climbing up vertical ducts and pipes and running in low pro le and high ground clearance con gurations" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003946_aict.2005.53-Figure2-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003946_aict.2005.53-Figure2-1.png", + "caption": "Figure 2. The updating of MQR routing tables", + "texts": [ + "00 \u00a9 2005 IEEE of available ports of node y and \u03b4(s, a) denotes the resulting state after applying action a to state s. The general idea of the MQR algorithm is presented in Fig. 1. The simultaneous updating of routing tables for nodes x and y speeds up the convergence. However, the exchanging of the additional information (4) from node y to x, causes that the load increases. Therefore, it need to be considered by a value of the factor \u03b7MQR, which is added to the transmission cost [11]. The main rule of packet propagation and the routing tables updating is shown in Fig. 2. Note, that autonomous agents cooperate exchanging their immediate reward (from x to y) and the expected reward (from y to x), as a result the effectiveness of the learning process is greatly increased [14]. Therefore, the local optimization (Q(x)) is included in the global optimization strategy. Such an approach increases a tolerance of faults or network changes (Fig. 3). Analyzing MQR algorithm we can distinguish two learning aspects. The first is called episodic, and in this case, it is the reaching the destination nodes by propagating packets, which is guaranteed by a negative value of the immediate reward (4); thus MQR is deadlock free" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001344_s0040-4020(97)00330-x-Figure2-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001344_s0040-4020(97)00330-x-Figure2-1.png", + "caption": "Figure 2", + "texts": [ + " Aromatic fluorine containing radical anions 9843 These data show that the energy differences between the two basis sets is negligible. Introduction of the electron leads to a change in the C - F bond length from 1.33Ato 1,36~. The major change is elongation of the C - C ortho meta bond flom 1.38Ato 1.45A.The reason for this is that the electron densityis Iodized and shared nearly equally between the ortho and meta carbons which is in excellent agreement with the ESR spectrum of the radical anion.8 The electron densitiesof the frontierorbitak are illustrated in Figure 2. It is importarrt to note that both the HOMO ad the LUMO orbhals are n orbitais whereas the LUMO+l is a a\u201c orbital. It was fust pointed out by Clarke and Coulsonz7that for dissociation of a radical anion, such as the fluorobenzeneradical anio~ to occur the electron must shitl born the rt\u201d orbital to a o\u201c orbital. There does not appear to be any driving force for such an electron shift. Clarke and Coulson27also argued that a third body was required for this shifl to occur. The importance of this will be discussed" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003243_ias.1998.732417-Figure2-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003243_ias.1998.732417-Figure2-1.png", + "caption": "Fig. 2. Prototype of the LIM servo system.", + "texts": [ + " In FE region instead of the voltage equation, the inner primary winding can be embodied on the form of the search coil that calculates the primary flux linkage directly. 4) In simulation, the transform angle of the control system and the calculated flux angle of FE region, thrust, current are analyzed. 5) To prove the propriety of the proposed method, a DSP installed experimental devices are equipped and the experiment is performed. 11. CALCULATION OF ASYMMETRJC CONSTANANTS A. Analysis Model The two dimensional model of LIM is shown in Fig. 1 , and the specifications are as follows in TABLE I. Fig.2 shows the prototype of the LIM servo system. In this paper, 1 ) The phase asymmetrical d-q circuit constants are 0-7803-4943-1/98/$10.00 0 1998 IEEE 799 For a linear induction motor (LIM), the constants of each phase are different due to the motor structure. Thus, it is difficult to perform the accurate vector control of LIM by the established rotation machine theory. The phase asymmetrical d-q circuit constants are calculated through the lock test, the equivalent no-loaded test, in which the non-magnetic conductor plate is removed from LIM, and the circuit equations" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002671_iros.1997.649067-Figure5-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002671_iros.1997.649067-Figure5-1.png", + "caption": "Figure 5: Wall Bracket with selected axes", + "texts": [ + " ,ui exists), then Rotate object by 19 (state S,, reached) If S,, $4 {Si>, update {Si} by adding S,,, Next U,+ Next up Next s k Terminate Afier obtaining all the existing states {Si} connected in the form of a tree structure, the states\u2019 data o b tained in tree form is transformed into a convenient form of network diagram, the states network because various states do appear more than once in the tree structure. Fig.4 illustrates such transformation for an imaginary example. In states network each state appears once only. States are connected by links depending on their inter-reachability. Double-sided arrows indicate that both the inter-linked states are mutually reachable fiom each-other. [Example]: Considering an object as shown in Fig.5. Though more locatable, but only three axes are selected for the sake of simplicity. These are one rotation; RI = [0 0 and two pivoting axes; PI = [0 0 1]*, P2 = [0 1 01\u2019. The states searching process is shown in Fig.6. Each node in rectangular form represents an orientation state. The angle of rotation to reach an offspring state from its parent state is written on each branch. All feasible axes at a state are written inside the rectangular. The state ID number appears in a circle besides each rectangular" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001935_0021-9797(84)90478-8-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001935_0021-9797(84)90478-8-Figure1-1.png", + "caption": "FIG. 1. Forces due to gravity between two adjacent settling particles of different sizes. For symbols see text.", + "texts": [ + " - (H + 2a) ~ -~ j [41 -32A FA -- 3as3( s 2 _ 4) 2 \u2022 [5] The derivative d f l / d H was calculated by taking the derivative of the approximation for 3 in case of small ra (see p. 155, Ref. (10). For the derivation of Fc we follow Crommelin (11), using a more general calculation which is also applicable to nonsperical primary particles and improving the expression for the drag force on a settling doublet. The total gravitational force between two particles of any shape and differing in size, in the position shown in the middle of Fig. 1, is the sum of the differences between the gravitational forces Fg and the difference between the friction forces Ff. Fc = Fg,1 - Fg,2 + Ff,1 - Ff,2. [6] In Eq. [6], the subscript 1 refers to the larger particle. When that is on top of the smaller one, F~ is a joining force (FG > 0). Contrary to the repulsive and attractive forces, FG does not depend on the interparticle distance as will be shown later. On a single settling particle in a fluid two forces are working, the gravitational force Fg 1 Fg = ~ 7rd3vAOg [7] and the friction force Fr Fr = - 37rnddv [8] where dv is the volume diameter (= diameter Journal of Colloid and Interface Science", + " It has to be emphasized that contrary to the DLVO interaction forces F~ is not dependent on the interparticle distance as long as the particles settle together and the interparticle distance is constant. But F~ does depend on the difference between the dimensions of the primary particles in a doublet, as shown in Fig. 2. F~ also depends on the shape of the primary particles. For cubes, the value o f F c , according to Eq. [13], is lower than for spheres with equal volume diameter, and depends on particle diameter and difference in particle di- ameter. Until now we dealt with the situation as depicted in the middle of Fig. 1. In the experiments we used suspensions with known, relatively narrow, particle size distributions. This allowed us to find the order of magnitude of F6. The angle (a) between the line connecting the centers of the two particles and the gravitational field influences the value of FG. The particle position on the right hand side of Fig. 1 results in a value of F6 that is smaller by a factor of cos (a). F6 reaches a m ax imum value for a -- 0 \u00b0. We conclude that the total gravitational interaction force FG is dependent on particle size and shape, the difference in size and on the angle a. Calculation of the forces acting upon particles in a doublet in a coarse suspension makes it possible to use a criterion for the formation of a doublet. As a criterion we set IFGI < I(FR + FA)max] [15] where (FR + FA)max is the m a x i m u m value of the net DLVO interaction force according to Eqs" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003918_j.jmatprotec.2005.02.163-Figure11-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003918_j.jmatprotec.2005.02.163-Figure11-1.png", + "caption": "Fig. 11. Flap bar.", + "texts": [ + " n the case where the operator must make a manual change, he plug can be removed in this way and at the end of the esting process; the pneumatic is used to open everything. As entioned above, this solution is unacceptable because of the isruption of elements that are above the surface. The second eason for neglecting this solution is that it is not very easy o manually operate both bars quickly to remove the plug. Since the operator uses both hands to remove the plug nd that it is necessary to brace at the testing site opens the ossibility to propose a solution of a flapping bar, as shown n Fig. 11. Similar to the linear bar, a spring could operate n which way the force reaches the plug or the fixing element. gain, the limited space is a rigid requirement. As mentioned, the fixation problem can be reduced to fix a olid body with one edge. If no parts are above the surface, it is bvious that the whole plug has to be inserted into the testing evice or even lower. The element that fixes the plug must e in the space between surface and the edge. The existing lements deal with this problem by a linear bar as described bove" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002100_pime_proc_1986_200_139_02-Figure4-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002100_pime_proc_1986_200_139_02-Figure4-1.png", + "caption": "Fig. 4 Contact coordinate system", + "texts": [ + " The element i, j has the dimensions 6 y x hx, where a 6x = - m (( 1 - x2/a2)b2) 6y = n However, (20) a m x = ( i - 4) 6x = ( i - 4) - therefore, (21) ([I - { ( i - $)/m)2]b2)\u201d2 n 6 y = Proc Instn Mech Engrs Vol 200 No C5 @ IMechE 1986 at Purdue University on March 13, 2015pic.sagepub.comDownloaded from The coordinates of the centroid of the element i,j are ((i - 4) sx, 0\u2019 - +) 6y) thus the radius r is given by ri j = [ { ( i - +) Sx}\u2019 + {(j - 4) 6y}]\u2019/\u2019 and the angle 4 by + i j = tan-\u2019{( 0: - $1 6Y } i - $) Sx For the coordinate system shown in Fig. 4, uspin = rijms uy = uspin cos $Jij Velocity of ball surface in rolling direction: = ub + uy Rolling velocity : Vr = ui + Urb Sliding velocity : us = ui - v Slide-roll ratio : rb Force on element in the y direction : SF, = pLp Sy Sx where p = - - 27tab 3P ( I - - - - ;:)1\u20192 Force in the x direction : S F , = -SF , tan 4ij = pj tan cPij 6y 6x Spin moment due to force in the y direction: 6M, = x SF, = rij cos q5ijpp Sy Sx Spin moment due to force in the x direction: SM, = -y SF, = rij sin + i j p p tan cPij 6y 6x Total spin moment: SMsPi, = SM, + 6My therefore, 6MSpin = rijpp(sin + i j tan 4ij + cos +ij) Sy Sx Equations (31) and (36) can be expressed as integrals: F , = I:a Sffi dY dx (38) (39) Since r, p, p and 4 vary throughout the ellipse, equations (23), (24), (32) and (50) must be substituted into (38) and (39)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002529_00037-1-Figure5-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002529_00037-1-Figure5-1.png", + "caption": "Figure 5 Wyoming modified IITRI compression test fixture: Size comparison with standard IITRI and Wyoming modified celanese compression test fixtures.", + "texts": [ + " It will be noted that the standard Celanese fixture is designed to accommodate a specimen only 6.3mm wide and 4mm thick. The much higher forces required to fail the larger specimen require a sturdier fixture. The IITRI test fixture was added to ASTMD 3410 in 1987, as Method B. It will be noted that this was a full 10 years after it was first introduced into the open literature. This is not untypical for new test methods and new test fixtures. To reduce both weight and cost, a smaller version of the IITRI fixture is available (Adams and Odom, 1991; WTF, 2000). Shown in Figure 5, this fixture can accommodate a 12.7mm wide specimen up to 7mm thick. It weighs only 10.5 kg, about one-quarter that of a standard IITRI fixture. Its cost is similar to that of the Celanese fixture. There are at least two so-called modifications of the Celanese fixture also. The Wyomingmodified Celanese fixture, shown in Figures 5 and 6, retains the efficient circular shape of the holders, but the alignment sleeve has been replaced by posts and linear bushings, like the IITRI fixture. Also, the wedge grips are tapered circular cylinders rather than cones" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000509_elan.1140040606-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000509_elan.1140040606-Figure1-1.png", + "caption": "FIGURE 1: lay; B: polyethylene ring; C; acrylic piece; D: silicone ring; E; stainless steel needle; F; 3/8 in. Swagelok nut; G: PTFE tape; H; stainless steel tube; 1, K: silicon tube; J; PTFE tube; L: plastic piece; M: Tygon tube; N: spring; 0, Q, T solenoid; P: aluminum support; R: aluminum arm; S, V: magnetic plate; U, W: screw; X: aluminum plate.", + "texts": [ + " Comparison with the model 303A SMDE from EG&G Princeton Applied Research and the 663 VA Stand from Metrohm, incorporating the MME, reveals other favorable characteristics of the proposed automatic mercury electrode, henceforth referred to as AME. Among them, potentiostatic i versus t drop extrusion curves closer to theory and easy adaptation to any cell, including flow ones, as described in this paper. The complete AME is relatively simple, inexpensive and can be reproduced as described. The electronic circuit allows autonomous or computer control of drop size and of the drop knocker. A cross section of the AME is shown in Figure 1. The glass capillary (A) is enwrapped with PTFE tape (GI, fitted into the acrylic piece (C), and held in place with a polyethylene ring (B). A stainless stell nut, 3/8 in. Swagelok, (F) is screwed over it. A lateral hole is made in the nut, through which a short stainless steel tube (HI is adapted and held in place with epoxy glue. A 150 cm long PTFE tube, about 1 mm i.d., connects this entrance to the mercury reservoir. (A disposable 20 mL polypropylene syringe containing some ml of mercury served for this purpose" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003492_elan.200403081-Figure2-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003492_elan.200403081-Figure2-1.png", + "caption": "Fig. 2. Schematic of the flow system.", + "texts": [ + " All batch measurements were carried out at room temperature in a 20 mL \u2013 Metrohm cell (Switzerland) with a three electrode configuration. A working carbon paste electrode, with a surface diameter of 3 mm, is made of a hollow Teflon tube in which a stainless steel rod is inserted and acts as the electrical contact between the carbon paste packed in one of the sides and the potentiostat. It is used together with a platinum wire counter electrode and a KCl saturated Ag/AgCl reference electrode. An Asincro magnetic stirrer from JP Selecta S.A (Spain) was used in the electrochemical pretreatment and the accumulation steps. Figure 2 shows schematically the FIA system used in the detection of leucoindigo. The 12 cylinder Perimax Spetec peristaltic pump (Spetec GmbH, Germany) allows the 0.1 M Tris-HCl pH 7.2 (2 M NaCl) stream to flow through the system. Desired solutions are injected by means of a sixport rotary valve, Model 1106 (Omnifit Ltd., UK) equipped with a 50 mL loop. Detector consists of a homemade thinlayer flow cell with a carbon paste electrode with a surface diameter of 3 mm. Flow cell consists of two methacrylate blocks and a PVC spacer fixed together with four screws", + " The first electrodic process is higher than the second one and it is not interfered with the reduction wave caused by oxygen in solution (compared with Figure 3A obtained by cyclic voltammetry). Using ACVas the electrochemical technique reversible processes of adsorbed species are greatly enhanced and very well discriminated from diffusion controlled processes. Based on this fact, the potential scan can be applied in the same accumulation solution (0.1 M NaOH, 0.8 g/L Na2S2O4). The alternating current voltammogram registered is shown in Figure 10 (curve B). In this case there is no interference due to the presence of sodium dithionite in solution (compared with Figure 2 obtained by cyclic voltammetry). Therefore, the washing step between adsorption Electroanalysis 2005, 17, No. 2 2005 WILEY-VCH Verlag GmbH & Co. KGaA, Weinheim of leucoindigo and recording signal stages was not necessary. Leucoindigo is detected in the same accumulation solution. In this basic media (NaOH 0.1 M) leucoindigo exhibits the same two electrodic processes at potentials around 0.7 V and 0.05 V when a potential scan is applied between 1.0 V and \u00fe 0.5 V. Thus, for the recording of the analytical signal, alternating current voltammetry scans were performed from 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002778_jpdc.2001.1820-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002778_jpdc.2001.1820-Figure1-1.png", + "caption": "FIG. 1. Single-body mass spatial parameters.", + "texts": [ + " Translational and angular velocities \u00f0v;w\u00de, translational and angular forces \u00f0f ;N \u00de at any point on a body in R3, are defined in terms of spatial velocity V , spatial acceleration \u2019V , and spatial force F in R6 as V w v \" # ; \u2019V \u2019w \u2019v \" # and F N f \" # : \u00f03\u00de In multibody dynamics, spatial quantities must be propagated and projected onto points or unique frames in order to be operated on. For this purpose, operators for translation and rotation are also defined in spatial operator forms as #Pi; i\u00fe1 U *pi; i\u00fe1 0 U \" # \u00f04\u00de and Ri; i\u00fe1 ri; i\u00fe1 0 0 ri; i\u00fe1 \" # ; \u00f05\u00de where U 2 R3 3 is the identity operator, pi; i\u00fe1 2 R3 is any vector joining two points (e.g., from point 2 to point 1 in Fig. 1), ri; i\u00fe1 2 R3 3 corresponds to the generalized rotation matrix that takes any point in coordinate frame i\u00fe 1 and projects it onto frame i, and *pi; i\u00fe1 is the skew symmetric matrix corresponding to the vector cross product operator also having the following additional properties (for any point in R3): *pi; jpk; l \u00bc pi; j pk; l; *pT i; j \u00bc *pi; j: \u00f06\u00de The 6 6 spatial inertia of a body is obtained by combining its mass and first and second moments of mass with respect to the point of interest in the body. Assuming that the body has a mass mi and moment of inertia Ji; cm about the body\u2019s center of mass, cm, the spatial inertia operator is then defined as Ii; cm Ji; cm 0 0 miU \" # 2 R6 6: \u00f07\u00de From Fig. 1 let Oi and cm, be two points on a rigid body i, and sOi be the vector from point Oi to cm. Furthermore, let cm be the center of mass of the rigid body. Let vi; cm and wi; cm be the translational and angular velocities of body i at cm, and Fi; cm and Ni; cm the forces and moment at and about cm, respectively. Then the forces and velocities at the point cm and Oi are related to each other as follows: Vi; cm \u00bc #S T Oi; cmVOi; \u00f08\u00de FOi \u00bc #SOi; cmFi; cm; \u00f09\u00de where #SOi; cm has the same form expressed by Eq", + " (8) with respect to time the expression for the spatial acceleration at and about point cm is obtained, \u2019V i; cm \u00bc #S T Oi; cm \u2019VOi \u00fe bOi; \u00f010\u00de where bOi \u00bc \u2019#S T Oi; cmVOi represents the velocity \u00f0V \u00de-dependent centrifugal acceleration components at point Oi. From the definition of the spatial moment, the spatial forces acting on the body\u2014 about cm\u2014correspond to Fi; cm \u2019Li; cm \u00bc Ii; cm \u2019V i; cm \u00fe \u2019I i; cmVi; cm; \u00f011\u00de where the second term is the velocity \u00f0V \u00de-dependent gyroscopic spatial force at cm (call it Pi; cm \u00bc \u2019I i; cmVi; cm). In order to complete our description and generalize the EOM for serially connected multibody chains, these equations are developed about points other than the center of mass. Assuming from Fig. 1 that point Oi corresponds to the point about which body i is rotating, then from Eq. (9) and Eq. (11) the spatial forces at/about Oi are FOi \u00bc #SOi; cmIi; cm #S T Oi; cm \u2019VOi \u00fe #SOi; cmIi; cmbOi \u00fe #SOi; cmPi; cm: \u00f012\u00de The first term is clearly identified as the application of the parallel axis theorem for spatial inertia, that allows the computation of the inertia about Oi, IOi \u00bc #SOi; cmIi; cm #S T Oi; cm: \u00f013\u00de The gyroscopic forces at point Oi are recognized from the second and third terms in Eq" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000533_bf02192245-Figure4-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000533_bf02192245-Figure4-1.png", + "caption": "Fig. 4.", + "texts": [ + " Note that, in this case, the position at time = 1 has fallen outside the range shown in the figure. Making the stepsizes smaller can get rid of the oscillations as seen in Fig. 3 (right). Proposition 2.2 states that there is a lower bound on the iteration time so that a given error level can be achieved. Increasing the iteration times Parallel gradient descent behavior with respect to iterations at each stage: (L) parallel gradient descent with r~ =0.1, Vz=0.1, ?,t= 15; (R) parallel gradient descent with r~=0.5, r2=0.5, 7'=1. can also compensate for the small stepsize as seen in Fig. 4 (left). However, there is a practical limit to how large the iteration times can be chosen as it will increase the magnitude of the oscillations. Reducing the iteration times may get rid of the oscillations as seen in Fig. 4 (right), but can lead to slow convergence. Finally, let us compare the path of parallel gradient descent with that of parallel decision making. The parallel decision making is expressed by the following iterative algorithm: p~ e argmin F,(p~- 1 , - 1 t- 1 t- 1). , . . . , P i - I , P i , P i + I . . . . . P N Pi In our simulated example, we assume that the sequence generated by this iterative algorithm is as shown by the thick solid path in Fig. 5 (right). For this particular example, it is observed that the evolution of game under an inaccurate search using the parallel gradient descent is very similar to that of parallel decision-making model" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001978_tgrs.1982.4307531-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001978_tgrs.1982.4307531-Figure1-1.png", + "caption": "Fig. 1. Current loop perpendicular to the dividing the two media surface.", + "texts": [ + " The displacement current for the following analysis is neglected. The transient state of the eddy current distribution inside the material and the distribution of the magnetic flux density in the air are examined. For the above analysis the diffusion differential equation is used. The Laplace transform is applied for the calculation of the transient state. II. DEVELOPMENT OF EQUATIONS A. Perpendicular Circular Loop A circular current loop is located in the half-space, while a conducting medium occupies the other half-space (Fig. 1). The loop is placed perpendicular to the dividing surface at a distance a. The magnetic vector potential (MVP) fulfills the following differential equations: V2A1 = 0, in region 1 (1) VA2A =HoAuA2 in region 2. (2) Manuscript received February 9, 1981; revised August 18, 1981. The authors are with the School of Electrical Engineering, Faculty of Technology, University of Thessaloniki, Thessaloniki, Greece. The current density within the conducting medium i2(X,~)z, t) = i2,x(x,yY,Z, t) -jO + i2$y(xIy,Iz,I t)YO (3)+ i2,Z (Xy,Y Z, t) io fulfills the differential equation v2T =u aai212~U t The Laplace transform L [2(x, y, z, t)] = I2 (x, y, z, s) is applied for the three components of the current density i2" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003197_memsys.2000.838488-Figure8-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003197_memsys.2000.838488-Figure8-1.png", + "caption": "Figure 8: Multilayer piezoelectric actuator fabricated by planer processes", + "texts": [ + " However, because Ni has etch resistance for KOH, complete cutoff of the Ni electrode by LAE was difficult in KOH solution as shown in Fig. 7(a). In this study, FeC13 solution was used to etch PZT and Ni, instead of general KOH solution. The FeC13 solution is one of the etchant for Ni, so good result is achieved as shown in Fig. 7(b). But if the FeC13 concentration is too high, non-assisted part will be etched away. So 2-5 wt% FeC13 solutions were used for this reason. The etching reaction of Ni is thermally assisted by the laser light. Displacement characteristic The fabricated actuators are shown in Fig. 8. These actuators have a multilayer structure in which 100pm thick PZT layers and 3 0 ~ thick Ni layers are stacked. The height of all actuators is 500pm. Each actuator shown in Fig. 8 has (a) 23 PZT layers, 3mm width, 3mm length, (b) 60 PZT layers, 2mm width, 8mm length, and (c) 120 PZT layers, 2mm width, 16\" length, respectively. These actuators were polarized by applying DC 200V at 100 \"C, for 30min., then expand displacement characteristics versus DC drive voltage were measured using the focus point detection mode of the scanning laser microscope. The result is shown in Fig. 9. The displacement at 100 V was about 2.5pm (23 layers) and 7 . 3 ~ (120 layers), respectively. These values agree with the calculated values from the piezoelectric properties of the material. This means that piezoelectric properties of the used material are preserved in the fabrication processes. Bending displacement problem When DC drive voltage was applied for the displacement measurement, undesirable bending displacement was also observed. Such displacement was especially large in the long actuator. The displacement of Fig. 8(c) actuator tip was about 50pm at 200 V. This value was 4 times larger than the expand displacement. It is assumed that a non-uniform force distribution in thickness causes the bending displacement. For example, 0 The wedge shape profile of the LAE grooves causes 0 The slight inclination of inner electrode induced by the dicing process causes expand force nonuniformity. This problem is not resolved yet, so there is room for further investigation. restriction force non-uniformity. Discussions for practical use In our method, a long, narrow actuator that has a large number of active layers can be easily fabricated" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002425_cbo9780511529627.012-Figure9.24-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002425_cbo9780511529627.012-Figure9.24-1.png", + "caption": "Figure 9.24. Contacts between planar features: (a) circular arc-line segment, (b) two circular arcs. Shading indicates the part interior.", + "texts": [ + " The derivation of the contact constraints proceeds by formulating the geometric conditions for the features to be in contact, and then substituting the configuration parameters into these expressions to obtain algebraic functions. We illustrate contact constraint derivation for general planar pairs whose shapes are formed by arc and line segments. There are three types of contact constraints, corresponding to the types of features in contact and their motions: moving arc-fixed line, moving line-fixed arc, and moving arc-fixed arc. Contacts involving points are identical to those for arcs of radius zero. Line-line contacts are subsumed by linepoint contacts. Figure 9.24(a) shows an arc-line contact. The contact condition is that the distance between the center o of the arc and the line Im equals the arc radius r: (oA \u2014 I) x (m \u2014 /) = dr, where the multiplication sign denotes the vector cross product, d is the length of the line segment, and its interior lies to the left when traversed from Horn. Figure 9.24(b) shows an arc-arc contact. The contact condition is that the distance between the centers equals the sum of the radii: (oA - p) \u2022 {pB = (r+s)\\ where r and s are positive for convex arcs and negative for concave arcs. Cambridge Books Online \u00a9 Cambridge University Press, 2009https://www.cambridge.org/core/terms. https://doi.org/10.1017/CB 9780511529627.012 Downloaded from https://www.cambridge.org/core. UCL, Institute of Education, on 17 Jul 2018 at 11:52:53, subject to the Cambridge Core terms of use, available at We obtain the contact constraint functions from these equations by expressing the vectors in coordinate form" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002993_1.1737377-Figure5-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002993_1.1737377-Figure5-1.png", + "caption": "Fig. 5 X mechanism in a singular configuration", + "texts": [ + " Composing C from d53 independent vectors from the involutive closures D\u0304i and D\u0304 j and with the projector ~9! the SVD yields PC 5UTSV and the matrix U\u03045S 20.25 0.433013 0.75 0.433013 0.0 0.433013 0.25 0.433013 20.75 0.0D . The remaining two kinematic constraints are obtained from ~11!. An advantageous circumstance when using the involutive closure basis is that the matrix C is well conditioned and the SVD does not suffer from numerical errors. X-Mechanism. While in the preceding examples the corresponding motion group was clear by inspection it is not so obvious for the X mechanism in Fig. 5. Assume for simplicity that the side lengths are a5b5A2. Evaluating Ci , j for all four joints yields that always d5dim Ci , j53. So J2,3 was taken as cut joint and the constraint matrix is L5S 0 cos q1 21 sin q3 sin q2 cos q3 0 cos q3 2sin q1 sin q3 0 sin q3 sin q1 cos q3 0 0 2c1 1 0 0 0 D . Assembling C and with P as above the SVD yields MAY 2004, Vol. 126 \u00d5 493 17 Terms of Use: http://www.asme.org/about-asme/terms-of-use Downloaded F U\u03045 1 A2 S 21 0 0 0 0 0 21 21 0 0 D and U\u0304PL5 1 A2 S 0 2cos q1 1 2sin q3 2sin q1 cos q3 0 D , which gives two independent constraints for the three generalized velocities ( q\u0307a)5( q\u03071, q\u03072, q\u03073). Hence the mechanism DOF is one. Also the assembling configuration in Fig. 5 is a singular configuration but very likely chosen as initial configuration for the X-mechanism. There are only two independent constraints in this configuration and again a simple rank determination is not admissible to eliminate redundant constraints. It can be shown analytically that C2,3 is equivalent to so(3). 6R Mechanism. An example for which the dimension of the constraint space really depends on the chosen cut joint is the 6R mechanism in Fig. 6 constructed from two planar 3R chains where the respective planes of motion are orthogonal" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001681_s0022-0728(85)80005-x-Figure6-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001681_s0022-0728(85)80005-x-Figure6-1.png", + "caption": "Fig. 6. Electrochemical cell for the thin-layer potential sweep voltammetric experiments.", + "texts": [ + " In this specific case, those molecules of AB whose reduction has been promoted by PCo are directly oxidizable at the electrode. Study in a thin-layer cell A few ~1 of a CH2C12 or CF3CO2H solution of the organic compound (P, AB, QC1 or QNH2) were deposited onto the glass plate forming one of the walls of the thin-layer cell [14] and were Mr-dried. The cell was then filled with the buffered solution. To ensure that the carbon disk was not in contact with the crystals, a very 66 thin sheet of paper was placed between the electrode and the spacer (cf. Fig. 6). The potential was scanned repeatedly at 10 mV s -1 between potentials including reduction of the mediator, until a steady-state voltammogram was reached. The conditions for the various experiments are given in Table 2. Without the paper sheet, the steady-state voltammogram is reached more rapidly. The enhancement of the apparent surface coverages in mediator solution (Fm) compared to the apparent surface coverages (F) in solution free of redox species shows the influence of the mediator on the global process" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002425_cbo9780511529627.012-Figure9.13-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002425_cbo9780511529627.012-Figure9.13-1.png", + "caption": "Figure 9.13. The CC open chain robot.", + "texts": [], + "surrounding_texts": [ + "Kaufman (1978) was the first to transform this mathematical result into an interactive graphics program for linkage design, called KINSYN. See also Rubel and Kaufman (1977). He used a modified game controller to provide the designer the ability to input a set of task positions. An important feature of this software was the decision to allow the designer to only specify four, not five, positions. Rather than obtain a finite number of RR chains, his software determined the cubic curve of solutions known as the center-point curve. This curve is obtained by setting the minor obtained from the first three equations in Equation (9.44) to zero. Kaufman's software would ask the designer to select two points on this curve in order to define two RR chains that it assembled into the one degree of freedom 4R linkage, or four-bar linkage. Analysis routines evaluate the performance of the design and provide a simulation of its movement. Erdman and Gustafson (1977) introduced LINCAGES, which, like KINSYN, focused on four task positions for the design of a 4R planar linkage. This software introduced a \"guide map\" that displayed the characteristics of every four-bar linkage that could be constructed from points on the center-point curve. This software was extended by Chase et al. (1981) to design an additional 3R chain to form a six-bar linkage. Waldron and Song (1981) introduced the design software RECSYN, which again sought 4R closed chains that guide a body through three or four task positions. Their innovation was an analytical formulation that ensured the linkage would not \"jam\" as it moved between the design positions. In linkage design a jam is equivalent to hitting a singular configuration in a robot, which occurs when the determinant of the Jacobian becomes zero. An important feature of this software was the growing reliance on graphical communication of geometric information regarding the characteristics of the available set of designs. Larochelle et al. (1993) introduced the Sphinx software for the design of spherical 4R linkages, which can be viewed as planar 4R linkages that are bent onto the surface of a sphere. A spherical RR chain is obtained when the link offset is p = 0, as shown in Figure 9.11. The fixed and moving axes of spherical RR chains are defined by the unit vectors G = (x, y, z)T and W = (X, /x, v)T. The workspace of relative rotations [Aik] reachable by this chain is defined by the algebraic equation GT[Alk]W = cosa. (9.45) This equation is evaluated at five specified task orientations to obtain equations that are essentially identical to Equation (9.44) and solved in the same way (McCarthy, 2000). Cambridge Books Online \u00a9 Cambridge University Press, 2009https://www.cambridge.org/core/terms. https://doi.org/10.1017/CB 9780511529627.012 Downloaded from https://www.cambridge.org/core. UCL, Institute of Education, on 17 Jul 2018 at 11:52:53, subject to the Cambridge Core terms of use, available at Following the pattern established by KINSYN and LINCAGES, Sphinx asks the designer to specify four task orientations, and then it generates the center-axis cone, which is the spherical equivalent of the center-point curve. The software also computes a \"typemap,\" which classifies by means of color coding the movement of every 4R chain that can be constructed from pairs of axes on this cone. The typemap also included filters that eliminated designs with known defects. A later version of this software called SphinxPC also included planar 4R linkage design (Ruth and McCarthy, 1997). The display and typemap windows of SphinxPC are shown in Figure 9.12. The three-dimensional nature of the interaction needed for spherical 4R linkage design presents severe visualization challenges. The designer finds that specifying a task as a set of spatial orientations is an unfamiliar experience. Furlong, Vance, and Larochelle (1998) used immersive virtual reality in their software IRIS to enhance this interaction. Larochelle (1998) introduced the SPADES software, which provided interactive design for a truly spatial linkage system, the 4C linkage, for the first time. A CC chain is the generalized robot link that allows both rotation about and sliding along each axis. Let G = (G, P x G) r and W = (G, Q x W) r be the Pliicker coordinates locating the fixed and moving axes in space. The workspace of this chain can be defined as the displacements [Du] that satisfy the pair of geometric constraints, (P x G)T[Alk]W + GT[Dlk](Q x W) = -psina. (9.46) These equations constrain the link parameters a and p to be constant for every position of the moving frame. There are 10 design parameters consisting of four parameters for each of the two axes, the link offset p, and twist angle a. By evaluating the two constraint equations at each of five task positions, we obtain 10 equations in 10 unknowns. Five of these are identical to those used for the synthesis of spherical RR chains and can be solved to determine the directions G and W. The remaining five equations are linear in the components of P x G and Q x W and are easily solved. SPADES generates the center-axis congruence, which is the set of spatial CC chains that reach four spatial task positions. It then assembles pairs of these chains Cambridge Books Online \u00a9 Cambridge University Press, 2009https://www.cambridge.org/core/terms. https://doi.org/10.1017/CB 9780511529627.012 Downloaded from https://www.cambridge.org/core. UCL, Institute of Education, on 17 Jul 2018 at 11:52:53, subject to the Cambridge Core terms of use, available at into two degree of freedom 4C linkages. This software demonstrates the significant visualization challenge that exists in the specification of spatial task frames and evaluation of candidate designs. These linkage design algorithms ensure that the workspace of each linkage includes the specified taskspace. However, in each case the designer is expected to Cambridge Books Online \u00a9 Cambridge University Press, 2009https://www.cambridge.org/core/terms. https://doi.org/10.1017/CB 9780511529627.012 Downloaded from https://www.cambridge.org/core. UCL, Institute of Education, on 17 Jul 2018 at 11:52:53, subject to the Cambridge Core terms of use, available at search performance measures and examine simulations for many candidate designs in order to verify the quality of movement between the individual task positions." + ] + }, + { + "image_filename": "designv11_65_0001903_0301-679x(85)90003-9-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001903_0301-679x(85)90003-9-Figure1-1.png", + "caption": "Fig 1 Diagram of misaligned journal bearing", + "texts": [ + " University Aachen, lnstitut f~r Maschinenelemente und Maschinengestaltung, Schinkelstrasse 8, 5100 Aachen, FRG The dimensionless Reynolds equation is written: a H a ~P ~ H a aP ) /~H 0--0 (--if- ~ ) + ~-~ (-if- ~ = 127r 3---0 (1) and the adiabatic dimensionless energy equation is given by: I.t2 //2 OP t)T /_/,2 aP c~T) (2) ~P 2 ( ~ _ ) 2 ) ) = 4 7 r ( 1 + 4 8 - ~ 2 ( ( ~ ) + For a shaft misaligned relative to its bearing bush, the oil film thickness variation in both axial and circumferential directions is described according to the notations given in Fig 1 such as 9 : H (0, V) = (1 + e (2-) cos (0 - ~ (2-)) (3) TRIBOLOGY international 0301-679X/85/010013-04 $03.00 \u00a9 1985 Butterworth & Co (Publishers) Ltd 13 where e (2) = (e 2 + 2'0 eo cos (0 - fro) + ' 02 ) 1/2 and '0 - - s i n ( 0 - ~ o ) qJ (2-) = ~o + arctan { eo } (1 + ( ~ ) c o s (0 - fro) eo The severity of beating misalignment is defined by a parameter D m where: '0e '0e D i l l -- '0rn {1 - [eo sin (0 - ffo)] 2 _ eo cos0 (\u00a2 - fro) ) (4) Maximum magnitude of misalignment occurs when Dm = 1" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002302_iros.1997.655123-Figure4-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002302_iros.1997.655123-Figure4-1.png", + "caption": "Figure 4: Force decomposition.", + "texts": [], + "surrounding_texts": [ + "948\nimposed by the edges and vertices involved in both contacts.\nProposition 2: Since &, lies on the intersection of the straight lines containing the ruling segments (equation (3)), then Eij is a segment of the curve Laj given by:\n64) B; sin +wj - Dj sin + w ~\nx = sin(.lltwj - $ w i )\nDi COS $wj - Dj COS $wi Y = - sin($wj - $wi) Q = P 4\nifsin($wj -$wi) # 0, otherwise it is a segment of the straight line defined by equation (3) f o r the orientation that satisfies Bi = Dj[g ] .\n3 Determining motion commands 3.1 Force-compliant control In order to be able to automatically execute assembly tasks with robots in the presence of uncertainty, fine-motion planners make use of active compliance strategies based on sensorial feedback: when contact exists between the mobile object and the fixed objects, the robot complies with the reaction force while moving along the projection of the commanded direction into the subspace orthogonal to the reaction force [6]. In this work, the force-compliant control based on the generalized damping model, represented by the following equation, is assumed:\nf(t) = B(&(t) - q t ) ) (5)\nwhere f(t) is the actual reaction force exerted by the fixed objects on the mobile object a t time t , B is the damping matrix, Go(t) the actual velocity of the mobile object, and &(t) the commanded velocity.\nBy using the generalized coordinates introduced in section 2 to describe the configuration of the mobile object, the force f = (Fx, Py) and the torque T acting on the object can be represented by the generalized force vector ?j = (Fx,F,, Fq), with Fq = r/p. Since p is the radius of gyration, the inner product defining orthogonality is the same as the one defining kinetic energy, thus ensuring that orthogonality makes physical sense [Z].\n3.2 Directions of motion Definition 6: The tangent plane ITt associated to a given contact configuration co of a basic contact i , is the plane tangent to .Fi at c,.\nI& is defined by the direction ii normal to the e-face at the contact point given by:\n( 6 ) - 1 n = -(n z , ~ y , n q l P ) an\nwhere\nA, =\nwith (nx,ny) = (cos$w,s in$~) being the normal to the contact edge, and ( T , , T ~ ) the vector from the contact point to the mobile object reference point.", + "949\nFor any contact configuration, the directions of movement that instantaneously maintain the contact are those that belong to the tangent plane. The following directions belonging to the tangent plane l\"It will be of interest (figures 1 and 2) : a) Direction Zr: Direction of pure rotation about the\n(71 t, = 6 - T y , % , P )\nA posit,ive motion along i$ corresponds to a rotation that increases the orientation Cp of the moving ob.iect. b) Direction t,: Direction perpendicular to ;T, being [frp TP, A] a right-handed frame.\nc ) Direction. ( 9 : Direction of pure sliding:\ncontact point: -0\nt', = b y 7 -n,,O) If9\nd) Direction Zq: Direction perpendicular to t\",. A corresponds to a rotation positive motion along\nthat increases 4.\nThe sense of t', is such that the frame [&, { 9 , 61 is a right-handed frame\nDefinition 7: The contact reference b a m e is the orthogonal reference frame [t\",, fp, 61 with its origin at the contact configuration.\nThe contact reference frame allows the vectorial decomposition of an applied force in order to analyze its effect on the movement. of the mobile object [2]. Figures B and 2 show the contact reference frame for type-A and type-B basic contacts, respectively, corresponding to different contact configurations. In each contact configuration, the tarye$ plane IIt is drawn together with the directions t,, t,, t\", and xq. The effect of friction for planar assembly tasks in the tridimensional Configuration Space C has been studied in depth by Erdamnn [2], who introduces the generalized friction cone. Definition 8: The gerzeralized fr ic t ion cone is the range of possible generalized reaction force directions arising from a basic contact in a given contact configuration.\nThe generalized friction cone is a bidimensional cone in C, determined by ii f pdf, 93 being the direction normal to the C-face defined in equation ( 6 ) , df the generalized friction vector, and p the friction coe6cient:\ndf = (nys -nr,wqlld (9)\nThe direction normal to ll, is the direction of pure rotation Tra\nThe effect of an applied force when the mobile object is in a one-point contact with the environment can be analyzed by decomposing that force, making use of the contact reference frame. As a result, a net force in the movement direction and a reaction force are obtained.\nLet ~ J A be the applied generalized force that points into the C-face associated to the basic contact. can be decomposed in the following way\n3 A = gf -t- i t , . (10)\nijf being the component on the plane ILf and iJtr the component along the direction t,, perpendicular to ITf . 4\nProposition 3: The reaction force ~ J R produced in a basic contact is t j ~ = -tjf i f iJf i s i m i d e the generalized f i c t i o n cone, or the negated projection of df along onto the edge of the fr ic t ion cone, otherwise\nProposition 4: T h e ne t force j j j , ~ that defines the movement direction i s the projection of j i ~ along the direction determined by $R into the plane\n121 *\n[2].\nwith uQ = nzrz + nyry. The unitary vectors in the 3.3 Motion commands that maintain directions defined by ii+pGf and il-pdf will be noted by e'+ and $-, respectively. one basic contact\nDefinition 9: The fr ic t ion plane IIf is the plane that contains the generalized friction cone.\nIn order to cope with uncertainty, fine-motions use the geometric constraints to guide the motion of the", + "950\nmobile object towards its goal, i.e. the commanded velocity t$ must allow the motion of the mobile object while maintaining contact with the fixed objects:\nwhere e v\"b: is the component that al%ows the motion towards\nthe goal, moving the mobile object reference point tangentially over the C-face; it is on the desired motion direction over the tangent plane. e i?f is the component that allows the contact maintenance, producing a reaction force; it is in the opposite direction of the edge of the generalized friction cone which is determined by the motion direction.\nThe magnitudes ut and vf depend on the desired velocity and on the desire< reaction force, respectively.\nLet us define (figure 5 ) : e p1 the start configuration ( 2 1 , y1, ~1). e p2 the desired stop configuration ( x 2 , y 2 , q 2 > . e paux the configuration obtained by rotating the\nobject an angle a, = (q2 - q l ) / p around the contact point at configuration pl :\nx,,,, = 2 1 + r , COS a,. - ry sin a,. (12) yanX = y~ + r, sin a,. + ry cos a,. Qaux = Qz\n0 dp the vector from paux to pa:\nProposition 5: For one basic contact :\nwhere as = (.it, &) and a , = (42 - q I ) / p . Proof: The direction of rotation about the contact point is t\", and the amount to be rotated is a,. The direction of pure sliding is t', and the amount to be translated in this direction is as , since the rotation about the contact point also translates the reference point. The proposed & is the composition of both motions. 0\nProposition 6: For one basic contact:\nProod: When the commanded velocity points towards the tangent plane, a reaction force arises and the mobile object moves along an instantaneous direction of motion over the tangent plane. If the object motion has a positive component along the direction zp, then the reaction force will have the direction d - , or e'+ otherwise. The proposed iJf is in the opposite direction. 0\nFigure 6 shows the decomposition of the commanded velocity iJc in its components." + ] + }, + { + "image_filename": "designv11_65_0003236_a:1020985609601-Figure2-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003236_a:1020985609601-Figure2-1.png", + "caption": "Fig. 2. Manifold for flow-injection measurements of the biosensor response: ( 1 ) mixer, ( 2 ) loop injector, and ( 3 ) counter electrode and reference electrode.", + "texts": [ + " The cholinesterase sensor was fixed with a sealing layer in a rectangular-shaped hollow located in the channel junction in the lower part of the cell (Fig. 1). All the junctions were implemented with polytetrafluoroethylene tubes (i.d. 0.5 mm). The circulation of solutions and the dosing of reagents were provided with a BR-1 flow-analyzer unit (AO Khimavtomatika, Moscow). The unit was equipped with a peristaltic pump and a loop for injection of a substrate and inhibitor. The loop volume was 45 \u00b5 L; fill-up time, 30 s; and flow rate, 0.1 mL/min. The manifold for biosensor-response measurements is given in Fig. 2. An anodic-oxidation current of thiocholine measured at +560 mV against an Ag/AgCl electrode served as a measured signal (biosensor response). The measurements were performed with a PA-2 voltammograph (Czechoslovakia). A nickel foil was used as a counter electrode. The foil and a reference electrode were placed into a vessel with a waste solution, which con- 1044 JOURNAL OF ANALYTICAL CHEMISTRY Vol. 57 No. 11 2002 IVANOV et al . tacted hydrolytically with the cholinesterase sensor. A common shape of the signals measured on intermittent injection of butyrylthiocholine iodide in the flow-injection mode is shown in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001138_1.2833218-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001138_1.2833218-Figure1-1.png", + "caption": "Fig. 1 Bearing dimensions for Case 1", + "texts": [ + " The plots shown here are for 2 milliseconds, as longer simulation would require prohibitively large CPU time. Four different cases of bearing applications are presented to examine the vibrations and the effects of various factors on the resulting vibrations. These factors include different sizes of the bearings, and the conditions under which they operate. 2.1 Case 1. Vibration of Rolling Bearings Under NoLoad Condition. From the manufacturers' catalogs, a stan dard ball bearing and a roller bearing are chosen (Fig. 1). Table 1 contains the geometrical and physical properties of the bearing and Table 2 gives the parameters used in the model. These parameters are normalized with respect to mass and radius of the rolling element. The normalized parameters are presented in Table 3. 2.1. J Normalized Form of Contact Stiffness and Its Deriva tive. To obtain the normalized form of stiffness expressions for both hne and point contacts, we follow the same procedure as that used in normalizing the equations of motion" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003429_s0007-8506(07)60688-x-Figure2-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003429_s0007-8506(07)60688-x-Figure2-1.png", + "caption": "Figure 2 : Active deflection compensation", + "texts": [ + " Subsequently, the finished part's deviation from the theoretical tooth profile averaged in respect of the gear's width dz, ma,,(i) is determined using a Visual Basic routine. Thereafter, the die cavity tooth profile is corrected using a displacement field ycw(i) = -dz, This procedure is to be iterated until the desired degree of accuracy of the finished part is reached. The idea underlying the concept of active deflection compensation is to counterbalance the pressure loads on the inner die walls - which cause the die to deflect elastically - by means of a counter pressure generated by an elastomer ring embedded in the lower die (Figure 2a). During the closing of the dies at the beginning of the cold sizing / forming process, a compressive counter stress phYdr is generated in the elastomer ring by the downward movement of the stopper ring attached to the upper die (Figure 2b). Being of equal magnitude, the pressure arising in the workpiece and the counter pressure generated in the elastomer ring compensate each other at the inner die walls and, thus, the elastic deflection of the lower die is inhibited. Consequently, the die geometry does not deviate from its unloaded state when loaded. At the end of the forming process, i.e. during unloading and the opening of the dies, both the pressure in the workpiece and the counter pressure in the elastomer ring decline and the stresses in the die are relieved (Figure 2c). This leads to substantially reduced ejector forces and die wear compared to conventional dies [9]. The FE-models of the lateral extrusion cold sizing process subject to this study are generated based on the theoretical tooth profile of the planetary gear. Due to its cyclic symmetry only a 20\" segment of the planetary gear is regarded in the simulations (Figure 3). In the course of this study, two variants of the cold sizing process are designed and investigated by means of FEA: one involving a conventional die with tooth profile correction and one incorporating the concept of active deflection compensation" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000213_bf00119549-Figure3-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000213_bf00119549-Figure3-1.png", + "caption": "Fig. 3. Front and side views of the Autonomous Benthi~ F.hqplorer [l 1]", + "texts": [ + " The results bear out that the situated reasoning architecture described in this paper can be used for semi-autonomous and au tonomous operations in underwater environments , and that it does allow for the easy integration of extant control-theoretic algorithms and human supervision. These exper iments , however, only point to the capabili ty to avoid obstacles and maintain vehicle stability, posit ion and att i tude while moving f rom way point to way point. If the power of the AI and control- theoret ic marr iage is to be bet ter exploited, more complex robot goals need to be undertaken. Candidate underwater tasks which would require such goals are those of the Autonomous Benthic Explorer (ABE) vehicle. ABE is a vehicle being designed by W H O I (see Figure 3) to pe r fo rm scientific survey of the sea floor over an extended period of time without a support vessel[ l 1]. A pr imary application of ABE will be to survey hydrothermal vent areas, so as to provide data concerning the long-term variability of the vents. ABE will fill a deficiency of manned submersibles and Remote ly Operated Vehicles (ROVs) by being able to remain on station over extended periods independent of a surface vessel, thus capturing the dynamic nature of vent processes. The final sections of this paper detail a simulation of thermal vent activity which we believe can be replicated in the test tank, and describe the new vehicle behaviors and their implementation for use against this simulation as a prelude to test tank exper iments " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002659_bf02984288-Figure2-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002659_bf02984288-Figure2-1.png", + "caption": "Fig. 2", + "texts": [ + " Candidates lbr analysis reference point Introducing the fully recursive algorithm, differences in the formulation according to the reference point used will be discussed. Any multibody system can be spanned to tree like structure by cutting a joint per each independent closed loop. If a body of a pair of bodies in a chain from leaf to the root is nearer to root of the tree, it is called the lower body, and the other is called the upper body. Upper body j of the adjacent two bodies always has a unique lower body and a unique associated joint j. Fig. 2 shows the kinematic configuration of the two adjacent bodies. In the figure, lower body i is denoted by body j -1 to avoid confusing index i and j from vague printout. All body parameters (e.g. centroid, joints, forced points, etc.) are defined at body reference frame ( Oj }. This frame is defined at joint j but fixed on upper body j. Provided that expressions for parameter conversion are well defined prior to the main analysis, no restrictions are placed on where the body reference frame should be defined (e" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002155_mssp.2001.1461-Figure4-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002155_mssp.2001.1461-Figure4-1.png", + "caption": "Figure 4. Illustration of the used nomenclature for the calculation of the separation, u(r), regarding the elastic contact between a cylinder and a rough plane.", + "texts": [ + " In the cases under consideration, the integration is much more complicated, because the separation, u, now is a function of the shape of the nominal surfaces, either this can be a contact of rough spheres [17] or the contact of rough cylinders [18]. Determined by an existing experimental set-up data the following representation is focused on the latter mentioned case. Then the calculations can be carried out in a twodimensional way in terms of cylinder length, ly, and the surface element dA0 can be treated as 1 dr that means a reduction of the surface integration to the dimension, r, of the contact area width, compared in Fig. 4. In a first step one has to define the relations between the separation, u, and the global geometry of the bodies in elastic contact. The dependency of u from the coordinate, r, generally consist of two terms: (i) the pure geometrical shape term of the nominal surfaces: u\u00f00\u00de \u00fe r2=2R \u00f024\u00de with R \u00bc R1R2 R1 \u00fe R2 \u00f025\u00de where u(0) is the minimum separation in the origin, R denotes the resulting radius, and R1,2 are the corresponding single radii of the two bodies, and (ii) the term which considers the elastic bulk deformation, w: w\u00f00\u00de w\u00f0r\u00de: \u00f026\u00de The index (0) also refers to the origin of contact, see Fig. 4, thus, the total separation can be described as follows: u\u00f0r\u00de \u00bc u\u00f00\u00de \u00fe r2=2R \u00bdw\u00f00\u00de w\u00f0r\u00de : \u00f027\u00de For the solution of the problem the following three expressions with respect to the contact width coordinate, r, are necessary: (i) the pressure distribution according to equations (18), (19) and (27) p\u00f0r\u00de \u00bc dF dr \u00bc Z 4 3 E0r 1=2 1 s3=2s I3=2 u\u00f00\u00de \u00fe r2=2R \u00bdw\u00f00\u00de w\u00f0r\u00de ss \u00f028\u00de (ii) the well-known elastic deformation solution of a semi-infinite plane, for instance [12] w\u00f00\u00de w\u00f0r\u00de \u00bc 2 pE0 Z 1 1 p\u00f0x\u00de ln x r x dx \u00f029\u00de which has to be subjected to a pressure distribution corresponding to equation (28), where, x, is the current coordinate, and (iii) according to equation (27), finally one gets u\u00f0r\u00de \u00bc u\u00f00\u00de \u00fe r2=2R 4 pE0 Z 1 0 p\u00f0x\u00de ln x r x dx: \u00f030\u00de After standardisation of the single variables, in [18] an approximate way of solving equation (30) is shown" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003012_1350650011543637-Figure4-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003012_1350650011543637-Figure4-1.png", + "caption": "Fig. 4 Cavitation zones", + "texts": [], + "surrounding_texts": [ + "The design quantities are the performance characteristics used in the design and analysis of a hydrodynamic bearing. The common quantities for all three bearings are the load capacity, the load position, the oil flowrate, the power loss, the temperature rise and the relative power loss. In the case of the journal bearing, the load position corresponds to the attitude angle \u00e2. Regarding the two thrust bearings, an additional design quantity is calculated, the coefficient of friction. All design quantities are listed in the Appendix." + ] + }, + { + "image_filename": "designv11_65_0003034_cdc.1994.411508-Figure2-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003034_cdc.1994.411508-Figure2-1.png", + "caption": "Fig. 2 - Defiiition of points S and M", + "texts": [ + " PX is tangent to the p a t r a t this moment and PY is perpendicular to PX. Similarly, at point C the coordinate system UCW is attached to 0-7803-1 968-0/94$4.0001994 IEEE the vehicle, CU being the longitudinal axis and CW the lateral axis. The distance CP is the position error and the angle between PX and CU is the orientation error. hereafter denoted by and 5. r e y t i v e l y . The sign notation for the errors conform with that o the coordinates. The steering angles for the front and rear wheels are denoted by S, and h, respectively, as depicted in figure 2a, which also shows the vectors of linear velocity for the points A and B, the middle points of the front and rear axels. a and b are the distances of A and B from C, respectively. When the vehicle moves, the front and rear tires roll momentarily along a circle with centre at point S, the intersection of lines normal to V, and V,. The projection of S on BA, denoted by M, is the only point on the longitudinal axis which has a purely forward velocity. whose magnitude V,,, defines the speed of the vehicle. The angular speed 61 at this moment is determined from: w = V, / (MS) (1) where the sign of MS determines whether w is positive or negative. The velocities of the other points on BA have a lateral component equivalent to a rotation by w about point M. as depicted in figure 2b. In this sense, v , = v , + m x m , v , = v , + a x m (2) The coordinates of point S can be determined from geometrical relations, as a tan 6, + b tan 6, CM = , M S = a + b tan 6, - tan 6, tan s, - tan s, (3) Equation (3) implies that the radius for a circular path is minimum when S, and S, assume their maximum absolute values and have different signs. Moreover, it may be seen by inspection that the steering wheel with larger angle (absolute value) is dominant. That is, when abs($) > abs(&), if the front wheel is to the right the vehicle turns right no matter whether the rear wheel is to the right or left" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001696_s0263-8223(01)00168-4-Figure8-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001696_s0263-8223(01)00168-4-Figure8-1.png", + "caption": "Fig. 8. Schematic diagram for the determination of: (a) Ky ; (b) Kxy ; (c) Kxz in line-loading condition.", + "texts": [], + "surrounding_texts": [ + "Eighteen specimens were manufactured by circumferential winding of plain weave E-glass woven cloth impregnated with epoxy resin on mandrels. The plain weave E-glass woven cloth has the same fibre geometry and mechanical properties in the wrap and fill directions. Loss on ignition test [15] was performed for the determination of fibre volume fraction. Tensile tests were performed on rectangular strips of unreinforced epoxy resin according to the standards set by ASTM [16], to determine the engineering properties of the resin. The elastic modulus and Poisson\u2019s ratio of the epoxy were found as 2.93 GPa and 0.38, respectively. The elastic modulus and Poisson\u2019s ratio of the fibres were taken to be 75.9 GPa and 0.22, respectively [17]. A modified \u2018rule of mixtures\u2019 [17] was used as a mathematical model to predict the engineering properties of the composite springs. Schematic diagrams of the experimental setups for the determination of spring rates Ky ;Kxy and Kxz are shown in Figs. 8(a)\u2013(c), respectively, for line-loading condition. Two semi-circular metal bars were attached diametrically to the specimen for ease of loading. The two semi-circular metal bars were connected externally by bolts and nuts on their extended ends to prevent movement and rotation. In surfaceloading condition, as shown in Figs. 9(a)\u2013(c), both the upper and lower flat contact surfaces of the composite spring were sandwiched, bonded and bolted rigidly by two 15 mm thick aluminium plates such that strain energy is mainly stored in the flexible arms when under loading. All specimens were loaded uniaxially in compression by a digital force gauge for determination of Ky and loaded by dead weights for determination of Kxy and Kxz while the displacements were measured by a dial gauge. A total number of 18 0/90 woven E-glass/epoxy composite springs with flat contact surfaces were manufactured. These specimens were fabricated by winding 5-ply, 7-ply and 9-ply of E-glass cloth impregnated with Ciba\u2013Geigy epoxy resin on custom- 3 7 6 P .C . T se et a l. / C o m p o site S tru ctu res 5 5 ( 2 0 0 2 ) 3 6 7 \u2013 3 8 6 made mandrels of 38, 57 and 85.5 mm nominal outside radii and all have the same width of 50 mm and the same semi-included angle of 37 . The geometry and properties of these specimens are listed in Table 1." + ] + }, + { + "image_filename": "designv11_65_0000585_rob.4620121004-Figure4-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000585_rob.4620121004-Figure4-1.png", + "caption": "Figure 4. Nomenclature for the (5-5)B manipulator.", + "texts": [ + " By an original solving procedure, four unknowns are eliminated while all spurious solutions-both those inherited from the adopted constraint equations and those inherent in the elimination procedure-are en iiiasse identified and purged away. As a result, a final polynomial equation of degree 24 is obtained whose 24 roots represent-in the complex field-as many assembly configurations for the (5-5)B fully parallel manipulator. It is worth noting that the number of assembly configurations agrees with the forecast published in the l i t e r a t ~ r e . ~ Finally, a cast study is reported that confirms the new theoretical results. 2. THE AUXILIARY STRUCTURE With reference to Figure 4, input data to the DPA of the (5-5)B fully parallel manipulator are: positions of points A;, AY, A*, A3, and A4, in a reference system 664 Journal of Robotic Systems-1995 Wt, fixed to the base; positions of points B1, Bi, Bi, B3, and B4 in a reference system W,, fixed to the platform; leg lengths L ; , LY, L;, L; , L3, and L4. Due to legs A;BI and A;\u2018B,, point BI is constrained to lie, with respect to the base, on a circle lying on a plane orthogonal to line A;A;\u2019. Such a circle has center A, and radius L1 given by where the convention adopted is to represent a vector as a point difference, and the scalar product of a vector times itself as the vector squared", + " Accordingly, the position of point A2 with respect to W,, can be expressed as: (A2 - B 4 , = L2(a2p cost$ + blp sin&) (6) where 82 is the counterclockwise rotation angle about (&\u2019 - Bi) that superimposes a2 on vector Clearly, performing the DPA of the (5-5)B fully parallel manipulator-once the leg lengths are given-is equivalent to finding the assembly configurations of the auxiliary structure represented in Fig. 5, where two revolute pairs are introduced in lieu of the couples of legs converging at points B1 and A2 of the (5-5)B fully parallel manipulator represented in Figure 4. Links AIBl and A2B2 of the auxiliary structure, respectively connected to platform and base by spherical pairs, are orthogonal to the axis of their extremity revolute pair. (A2 - B 2 ) . 3. THE FIRST SUBSET OF CONSTRAINT EQUATIONS With reference to the auxiliary structure of Figure 5, the guidelines for laying down the constraint equations can be summarized as follows. The position of the platform with respect to the base is first expressed as a function of as reduced a number of parameters as possible", + " (42) and (43), whose left-hand sides can now be considered as polynomials in t l . Such polynomials admit of a greatest common divisor of the first order in tl; by imposing the vanishing of this greatest common divisor, the value f l l , for f l can be determined. Hence also solution 13~1, for 8, can be found. Finally, for 81 = 8lh and 8 2 = 82j1, linear system (34) yields the values ulll, u21,, and u3/, for u l , u2, and u3. Value O11, provides, through Eq. (3), the position of point B1 (see Fig. 4) with respect to reference frame Wb. Moreover (see Eq. (7)), values ull,, uzfl, and u3J: allow the position of point B2 in WI, to be determined too. At this stage of the analysis, all terms on the right-hand side of Eq. (22) are to be considered as known, which implies that positions of points B3 and B4 with respect to reference frame Wb, can be directly determined. Finally, any three points out of B1, Bz, B3, and B4 can be exploited to determine the position in Wb of points B$ and B! belonging to the platform of the (5-5)B manipulator. Because every root of Eq. (67) univocally identifies an assembly configuration, it can be concluded that the DPA of the (5-5)B fully parallel manipulator admits of 24 solutions in the complex field. 676 Journal of Robotic Systems-1995 6. NUMERICAL EXAMPLE With reference to Figure 4, a (5-5)B fully parallel manipulator is considered with points on the base and platform having the following coordinates in reference systems wh and W,, respectively (all lengths are given in arbitrary length unit): A; = (5.1, 0.9, -0.3) B1 = (4.6, -1.9, 3.8) A;\u2019 = (4.2, -1.2, 0.4) A2 = (0.8, 4.5, 0.9) A3 = (1.6, 2.2, -1.4) A4 = (-2.3, -1.7, 0.7) B; = (2.1, 0.0, 5.1) Eli = (-2.4, 1.4, 4.7) B3 = (-1.5, 4.4, 4.6) B4 = (-3.7, -1.9, 5.6) Input is given in terms of leg lengths: L; = 5.9; L3 = 7.1; L4 = 5" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002529_00037-1-Figure4-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002529_00037-1-Figure4-1.png", + "caption": "Figure 4 IITRI compression test fixture (ASTM D 3410).", + "texts": [ + ", Adsit, 1983) that the Celanese fixture can produce compressive properties equal to the other shear-loading fixtures subsequently developed. But it requires more care, and a higher skill level, which is a definite disadvantage. Following the almost immediate criticisms of the Celanese fixture as soon as it was standar- dized by ASTM, Hofer and Rao (1977) of the Illinois Institute of Technology Research Institute (IITRI) introduced what became known as the IITRI compression test fixture, shown in Figure 4. The two major deficiencies of the Celanese fixture were eliminated. Flat wedge grips, like tensile wedge grips but inverted, were used. These could accommodate a range (about 2.5mm) of specimen thickness, while always maintaining full contact with the holders. Also, the \u00aaalignment\u00ba sleeve was replaced by posts and linear ball bushings, which truly do maintain alignment. It is essentially impossible to bind up linear bushings. The deficiency of the IITRI fixture is its extreme weight, typically about 45 kg (WTF, 2000)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003454_1.1757489-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003454_1.1757489-Figure1-1.png", + "caption": "Fig. 1 An example slider", + "texts": [ + " We then choose some contemporary slider designs and show that the three-dimensional effects lead to markedly different results from the 2-dimensional flow analysis @5,6#. Finally, the predicted contamination pattern is compared with experimental results for a particular ABS design. In order to determine the trajectory of a particle within the air-bearing we must first determine the spacing between the slider and disk surfaces as well as the pressure and velocity fields. For a complex slider design with etch steps as shown in Fig. 1, the point-to-point spacing between the slider and the disk surfaces varies abruptly in places, introducing local three-dimensional airflow. Since the spacing is about three to five orders of magnitude less than the slider\u2019s lateral dimensions, we retain the following assumptions: 1. The pressure gradient in the vertical direction is negligible; therefore, the pressure field calculated from the Reynolds equation is still considered to be valid. 004 by ASME OCTOBER 2004, Vol. 126 \u00d5 745 s of Use: http://www", + " The particles are first assumed to be uniformly distributed above the disk surface with velocities close to the air-bearing\u2019s velocity where the particles are located, as determined by Eqs. ~5! and ~6!. The particle\u2019s initial vertical velocity is assumed to be zero. In this section the three-dimensional airflow effects on particle trajectories and contamination are studied. A representative mod- rom: http://tribology.asmedigitalcollection.asme.org/ on 08/30/2017 Term ern negative pressure slider is chosen for this study ~Fig. 1!. The flying height of the slider is 26 nm. A sketch of the slider-disk assembly and coordinates is given in Fig. 2. The pressure profile shown in Fig. 3 is obtained without three-dimensional effects by solving the Reynolds equation for the air-bearing of the slider using the CML code Quick 4. To study particle flow in the air-bearing, we first calculated the spacing function between the slider and the disk. The slider-disk spacing map is shown in Fig. 4, where it can be observed that particles may enter the recessed region of the air-bearing through the leading edge of the slider" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001529_iemdc.2001.939301-Figure5-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001529_iemdc.2001.939301-Figure5-1.png", + "caption": "Fig. 5. Flux lines at an instant of time.", + "texts": [], + "surrounding_texts": [ + "rotor motion. At each step, the model computes magnetic quantities like the magnetic field distribution inside the machine, mechanical quantities like the rotor position and torque, and circuit quantities like the current in the windings. The air gap flux density data as a function of spatial location is stored at every step, resulting in a two-dimensional table with both spatial and temporal information. Both the Virtual Work and Maxwell Stress Tensor methods were used to compute the torque and observed to provide comparable results. The use of post-processing techniques to improve the accuracy of the results, as in [3], did not produce any significant improvement. The values of the torque and radial force for the no-load and load case obtained from the Maxwell Stress Tensor method (MST), the Virtual Work method, and MST using a local post-processing technique are compared in Table 1.\nForce and torque results appear to have converged for the second order mesh used. The mesh, however, does affect the results obtained as has been noted in [4]. The transient solution method uses discrete time steps solving the field problem at time t, t+At, etc. At every time step, the rotor is rotated and the mesh in the air gap is distorted and the interior angles of the air gap elements are computed. When the distortion becomes too large the air gap is remeshed. This results in time steps with varying degrees of mesh quality. Fictitious torque pulsation have been observed to be caused by the remeshing of the air gap and the slight distortion caused by that remeshing. To avoid this, the size of the airgap element and the time step were chosen so that the rotor turns through an angle equivalent to the distance between air gap nodes. This results in the air gap elements only being renumbered between time-steps. Three layers of elements were used in the air gap to improve the accuracy of the torque computation [ 5 ] .\nA model with a round symmetric stator bore with a periodic mesh pattem as shown in Fig. 2 generates fictitious torque ripples similar to the cogging torque produced by the teeth. To c o n f i i this, the stator was\ncompletely removed and replaced with a Neumann boundary. This results in the flux being normal to the airgap-stator boundary, which is equivalent to having a stator with infinite permeability and, as such, is an approximation to a uniform stator with a high p. Fig. 3 and Fig. 4 compare the torque for the case with the periodic mesh in the stator, and the case with the unmeshed stator, respectively. While there is a distinct torque harmonic due to the mesh pattem, the numbers are observed to be negligible compared to the major torque harmonics.\n209", + "A number of these cases were also studied under full load conditions with the stator winding excited by steady state sinusoidal excitation.\nFour different kinds of unbalances were simulated as follows:\nStatic eccentricity was simulated by offsetting the axis of the stator with respect to the center of the rotation of the rotor, i.e. in Fig. 6, CZ is the axis of rotation of the rotor, and is shifted from C, by a certain percentage of the air gap\u2019 length. The dynamic eccentricity was simulated by offsetting the axis of the rotor with respect to the center of rotation, i.e. in Fig. 6 C1 is the axis of rotation of the rotor. An issue here is the path along which the air gap flux density will be computed. No matter which path is chosen, its distance from either the stator or the rotor will change at every timestep. This, by itself, could introduce some harmonics in the air gap flux density being studied. Since there is no better choice, a fixed path at a constant distance from the stator bore was chosen for the air gap flux density computations. The combined static and dynamic eccentricity was simulated by combining the two afore-mentioned offsets, i.e. in Fig. 6 the axis of rotation of the rotor is neither C1 nor Cz. This results in a synchronized combined eccentricity where the same point on the rotor remains closest to the stator bore. This is the most common type of combined eccentricity in machines caused by a bent or uncentered rotor. Other kinds of combined eccentricities can be modeled using more elaborate schemes. The magnet unbalance was modeled by using a different magnet strength in one of the magnets.\nIn. SIMULATION RESULTS The results from the above cases are compared with those from the base case with a symmetric rotor. The comparisons are based on the torque harmonics and the 2D FFT of the airgap flux density.\nA. No Load Cases\nThe torque harmonics for some of the no-load cases studied are shown in Figs. 7 - 11. The results show that the effect of unbalances on the torque harmonics is very small, for eccentricities up to 30% of the air gap. This can be explained by noting that the effective air gap of the machme includes the pemanent magnet. Since the air gap is very small compared to the magnet length, eccentricities in the actual airgap translate to very small changes in the effective air gap of the motor.\nFig. 12 shows plots of the radial component of the airgap flux density as a function of angular position, at various instants of time for the symmetric no-load case. The six-pole pattern is clearly evident. Superimposed on this are significant spatial harmonics due to slotting and saturation, which can be identified by applying the standard FFT to the waveforms. A spatial harmonic analysis will, however, only present part of the total picture. As seen in the plots, the flux density varies with time as well as with position, and could have time harmonics as well, resulting in a set of sinusoidal traveling waves. The individual traveling wave components can be determined by a 2D-FFT [9].\nTable 2 shows the significant components obtained from a 2D-FFT analysis of the airgap flux density over one electrical cycle, for the symmetric and static eccentricity cases. The symbols used in the table are:\nP = number of pole pairs (3); S = number of stator slots (36); o = synchronous speed (2Q;\n210", + "8 = spatial position in airgap; and, t = time.\nIt is possible to verify the harmonic orders obtained with those predicted by the h4MF-permeance wave method. The first wave in the table is the fundamental wave rotating at synchronous speed, produced by the fundamental MMF acting on the average component of the airgap permeance wave.\nThe next three waves are produced by the interaction of various h4MF components with the fundamental variation in airgap permeance resulting from stator slots.\nE # * - ! c 4. I\n0 6h l 0 B 150 900 250 aor Bp1\nam& wi\nswb w*vi Fig. 12. Spatial distribution of the airgap flux density at a few instants of an electrical cycle.\n21 I" + ] + }, + { + "image_filename": "designv11_65_0001479_iros.2000.893229-Figure5-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001479_iros.2000.893229-Figure5-1.png", + "caption": "Figure 5 shows a sequence VIF -+ VIE -+ VIF with motion direction MD parallel to the left polyhedron face.", + "texts": [], + "surrounding_texts": [ + "ment measured by a wrist-mounted FMS when the grasped DLO changes its contact state with respect to a rigid polyhedral obstacle.\nOur investigation is based on the main assumption that the deformation caused by gravity and contact force is mainly elastic. Therefore, the DLO can be regarded as a bending spring. For the theoretical discussion of force and moment in this section, we assume an ideal forcdmoment sensor, i.e., a sensor with infinite resolution and without internal noise. For now, we neither consider grav-\nstacle not being in contact) and a stable contact state different from N. According to the definitions given above, they are a subset of the initiated transitions. However, it is reasonable to regard them separately because they always form the first step of the assembly process. The possible transitions are:\nAs stated in [7], two types of contact states must be distinguished. Stable contact states, i.e. {N, EIF, EIE, VIF} in Figure 1, are kept up in the presence of distortions and uncertainties. On the contrary, instable contact states, i.e. {EN, VIE, VN}, may change spontaneously. Consequently, we also distinguish two types of state transitions. Initiated transitions are directly caused by a distinct motion of the robot gripper. They start in a stable state and lead to a stable or instable successor. Spontaneous transitions cannot be directly controlled. They start in an instable state and lead to a stable successor. Accordingly, spontaneous transitions always occur as the second transition in a sequence \u201cStable + Instable 4 Stable\u201d.\nBeing F ( t ) and M ( t ) the vectors of force and moment measured by the FMS versus time t , we assume that any state transition is generally accompanied by a characteristic change in the course of these functions.\n2.1 Initial transitions At first, initial transitions shall be regarded which occur between state N (with DLO and ob-\n* Some additional assumptions conceming the contact state transitions which are of minor importance here are given in REMDE et al. [7].\nWe assume that both, force and moment, are zero if the DLO is in state N. When the workpiece touches the obstacle face, it is deformed, causing force F and moment M. Both of them increase as long as the gripper motion is continued, as shown in Figure 2c.3 The same result is obtained for the transitions N + EIF and N + EIE. During the complete motion, F is aligned with the normal of the obstacle face.\nIf any of these transitions is performed in the inverse direction, i.e., a stable contact is loosened, the course of F and M is simply inverse, too.\n2.2 Other initiated transitions Next, we want to consider initiated transitions occurring between stable states different from N.4 The following cases are possible:\nLet us think of a transition EIE -+ EIF. Figure 3 shows the two fundamental cases, distinguished by the motion direction of the gripper, which is MD1 or MD2, respectively.\nFor MD1, the angle a between DLO and face decreases until the transition to EIF occurs for a 4 0. Before the transition, the gripper motion causes sliding of the DLO to the left, resulting in a\nVIF H EIF EIE H EIF\nFor clarity reasons, only the course of F = IF I is shown in the diagrams. It is qualitatively equal to the course of A4 in all cases discussed in the following.\nPlease note that sliding of the DLO on an obstacle edgelface is not regarded here, unless it causes a state transition to another edge/face of the obstacle.\n- 1481 -", + "decreasing curve length of the DLO between gripper and contact point. This causes an increasing flexural rigidity of the DLO and, thus, an increasing contact force. (The flexural rigidity of a (linear) spring fixed at one end (gripper) with distance / t o the point of force application (contact point) is proportional to l / P ) . After the state change, the point of force application moves quickly along the contact line towards the gripper. This results in a steeply increasing contact force. The same effect is observed for a transition VIF -+ \u20acIF.\nIf the motion direction is parallel to the face with which the contact is established (MDz), the situation is different. Both DLO-bending and length / between gripper and contact point are increased while dragging the DLO over the obstacle edge. On the one hand, the increased bending causes an increasing contact force. On the other hand, the increase of 1 results in a decreasing flexural rigidity. Because the second effect overbalances the first one, the contact force decreases. After the transition to state EIF, it stays c o n ~ t a n t . ~\nIn both cases discussed above, the direction of contact force F is perpendicular to the DLO-tangent in the contact point. Thus, F is rotated during the motion and is aligned with the normal of the obstacle face after the transition. Like for the initial transitions (Section 2.1), the course of force and moment is just inverse if any of the transitions is performed in the inverse direction.\n2.3 Spontaneous state transitions The last kind of transitions to be considered are spontaneous transitions starting in an instable state. Because they always occur as second transition in a sequence \u201cStable -+ Instable -+ Stable\u201d, it is reasonable to regard the complete sequence.\nPlease note that independent of the motion direction, the slope difference is always positive. This holds true also if the motion direction is somewhere between MD1 and MD2.\nThere are two fundamental cases: sequences causing a loss of contact (leading to N) and sequences leading to a state different from N. 1) The following sequences belong to the first\ngroup: VIF -+ VIE -+ N EIE -+ E N -+ N\nEIE -+ VIE -+ N EIF -+ E N -+ N\nAs an example, Figure4 shows a sequence VIF -+ VIE -+ N. Because state VIE is a borderline case of state VIF, the transition VIE -+ N means a sudden loss of contact. Therefore, the bending stress of the DLO is suddenly released and the DLO (which may be represented by a spring-mass system) performs a damped oscillation. The frequency of the oscillation depends on the material parameters and length / of the DLO, while the amplitude additionally depends on the force before contact is lost.\n2a) The behavior for the sequences VIF -+ VIE -+ EIE\ncan be traced back to the behavior for initiated transitions (Section 2.2). Let us regard a transition EIE +VIE -+ VIF. This problem is found to be very similar to a transition EIE -+ EIF with motion direction MDz. Depending on the specific gripper trajectory and the geometry of both DLO and obstacle, either of these transitions may occur. Accordingly, the course of force and moment is (qualitatively) equal. The inverse transition, i.e., VIF -+ VIE -+ El\u20ac, corresponds to a transition EIF -+ \u20ac/E. 2b) Finally, we have to regard the following se-\n\u20ac I \u20ac -+ VIE -+ VIF\nWhile the contact force is constant before the change of contact state, the transition obviously causes a tension release for the DLO. Therefore, the contact force is reduced and the qualitative\n- 1482 -", + "course of force and moment is given by Figure k6 F is aligned with the normal of the left face before the transition and is aligned with the normal of the right face after the transition. A similar behavior is obtained for the other transitions of this group.\ngripper)\nIn all of the examples discussed so far, we have assumed a \u201cspecial\u201d motion direction of the gripper (e.g., parallel or perpendicular to an obstacle face) for which the DLO behavior can be derived quite easily. For a more general gripper motion, it must be considered that a motion approaching the obstacle generally increases the DLO bending (causing an increasing contact force), while a withdraw from the obstacle causes a decreasing contact force. Thus, the resulting force and moment can be obtained by superimposing this effect and the behavior discussed above.\n3 Sensor signal processing\n3.1 General considerations After the theoretical discussion of the state transitions in the previous section, it shall be considered how they may be detected by means of a (non-ideal) forcelmoment sensor mounted between robot wrist and gripper. Before discussing the evaluation, some general aspects shall be regarded.\nSo far, we have discussed the forces caused by the interaction between DLO and obstacle. However, in most practical situations, there is a significant additional load due to gravity. As long as we restrict the gripper trajectory to translational motions, gravity only causes a constant offset for both force and moment. Therefore, it does not have to be considered if absolute force and moment thresholds are avoided in the detection algorithm.\nWe generally assume that the direction MD of the (linear) gripper motion is not changed when the state transition occurs.\nIn opposite to the handling of rigid workpieces, especially two additional problems must be regarded: First, the force caused by contact of the DLO with the rigid environment is generally low. This results in the need for a high resolution FMS. However, the resolution is correlated with the measuring range (A high resolution requires a small measuring range and vice versa). Therefore, the measuring range should be as small as possible. Because the force and moment caused by gravity are typically much larger than the contact force, the required measuring range (and thus the obtained resolution) is mainly determined by the gripper mass.\nSecond, all contact forces highly depend on both workpiece and specific situation. Therefore, it is necessary to either have a rather precise a-priory knowledge about both workpiece and task, or to base the transition detection on robust algorithms, evaluating the qualitative course of the force or moment signals rather than their absolute values. In order to obtain easily applicable algorithms, this second approach is preferred.\nFor evaluating the 6D force/moment vector provided by the sensor, there are several possibilities. In the following, we evaluate the components of force or moment vector which reflect the state transition according to Section 2. In many cases, the absolute values of force or moment may also be regarded. The direction of the vectors is not evaluated here because it does not promise to be advantageous and the effort for a vectorial evaluation is rather high.\nAs a fvst step of the signal processing, some low-pass filtering is generally required for noise reduction. In our experiments, we succeed in using two serial moving-average low-pass filters. The first one is a hardware filter which is part of the commercial FMS signal-processing board, the second one is part of the signal-processing software. In the following, we generally refer to sufficiently filtered signal^.^\n3.2 Methods for slope change detection According to the discussion of the single state transitions in Section 2, a contact state transition may either cause a (rather abrupt) slope change or damped oscillations in the force and moment signals. Therefore, two kinds of detection algorithms are required.\nBecause the detection of oscillations of constant frequency is found to be rather simple in both time or frequency domain, this section is focused on the on-line detection of slope changes.\nBeing f(t) the function to be evaluated, t the current instant of time and to the instant of time at which the state transition occurs, we have to decide between the following \u201cno-change hypothesis\u201d Ho and \u201cchange hypothesis\u201d HI.\nThe usage of other (linear) low-pass filters is also possible but is not found to be advantageous in our experiments.\n- 1483-" + ] + }, + { + "image_filename": "designv11_65_0003862_1-4020-3317-6_7-Figure4-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003862_1-4020-3317-6_7-Figure4-1.png", + "caption": "Fig. 4. Corner wrinkles: (a) overall shape; (b) central cross section and definition of half-wavelength.", + "texts": [ + " We will assume that this stress is given by the buckling stress of an infinitely long plate of width \u03bb \u03c3cr = \u2212\u03c02 \u03bb2 Et2 12(1 \u2212 \u03bd2) (6) In the case of fan-shaped wrinkles we will set \u03bb equal to the half-wavelength mid way between the corner and the edge of the fan. To estimate the wrinkle details, we begin by considering a simple analytical expression for the shape of the wrinkled surface. For example, in the case of a symmetrically loaded membrane we assume that at each corner there is a set of uniform, radial wrinkles whose out-of-plane shape can be described in the polar coordinate system of Fig. 4 by w = A sin \u03c0(r \u2212 r1) Rwrin \u2212 r1 sin 2n\u03b8 (7) where A is the wrinkle amplitude, n the total number of wrinkles at the corner \u2014each subtending an angle of \u03c0/2n\u2014 and \u03b8 is an angular coordinate measured from the edge of the membrane. Since the stress in the corner regions is uniaxial there is the possibility of wrinkles forming there. The radial strain is \u03b5r = \u03c3r/E (8) where \u03c3r is given by Eq. (1). The corresponding radial displacement, u(r) (positive outwards), can be obtained from u = \u222b \u03b5rdr + c (9) where the constant of integration c can be obtained by noting that u \u2248 0 at r = R, i" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003651_cira.2003.1222168-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003651_cira.2003.1222168-Figure1-1.png", + "caption": "Figure 1: Dynamic whole arm and body manipulation", + "texts": [ + " The second property enables power assist for helper in the field of social welfare. Lately such systems have been expected by hospital, etc. The studies by Lynch et al. which is referred to as \u201cDynamic Manipulation\u201d [7][8] are also good examples related to whole body manipulation. The purpose of the present investigation is to formulate the basic control algorithm of WAM system under gravity, and in this paper, we deal with WAM system which consists of 2 D.O.F. planar robotic arm with a circular object. As shown in Fig. 1, in the following p a per, we study: 1) how to manipulate the object within the 2 link arm in section 2, 2) how to operate the object according to the environmental constraint in section 3, and 3) bow to we whole body in object manipulation in section 4, respectively. The validity of the control method is investigated by numerical simulations. 2 M o d e l i n g and Control of a Whole Arm Manipulation System In the following sections, we consider a WAM system by 2 D.O.F. planar full-actuated robotic arm as shown in Fig", + " From the simulation results we can see that the control aim is achieved. This manipulation form suggests that by the use of preexist environment effectively, heavy objects are able to be manipulated by small torques. This is very important from the torque-limitation point of view. 4 From Whole Arm to Whole Body Manip ulation It is well known that, when human try to remove some heavy object, he/she actively uses not only his/her hand but the arm and further may be the whole body so as to generate large force as shown in Fig. 1. Such operation is also important to balance the whole body. When the body part do not move, the basic control problem formulation is as same as section 3, however, if we also consider the movement of the body part, for example, to balance the body, then the geometric constraint will be very complex. We are now considering to combine the formulation of sections 2 and 3 to solve this problem by introducing the virtual Link to obtain the Jacobian about the contact between the body and the object. 5 Conclusions In this paper, we have considered basic modeling and control of a WAM system" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000280_ip-gtd:19970911-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000280_ip-gtd:19970911-Figure1-1.png", + "caption": "Fig. 1 1 System setup", + "texts": [ + " The reasons are due to changes of generation patterns in meshed networks and the limitations in constructing transmission lines due to environmental considerations leading to higher stresses in present day power systems. This paper deals with the voltage collapse from a practical point of view. A laboratory power system model is used to investigate how the different system components interact and affect voltage stability. Special attention is given to load composition and to generator current limiters. Voltage collapse usually occurs in a time frame of seconds to tens of minutes, and the model used is particulary suited for these time spans. The power system model (Fig. 1) consists of a power plant feeding different loads over a transmission line configuration which may include one or two transformers with on-load tap changers (OLTCs) [3] . All control systems used are real controllers as applied in power systems. 'This eliminates the modelling of controls which would only be representative of actual control behaviour. The rated voltage for the model is 400V. 0 IEE, 1997 IEE Proceedings online no. 1997091 1 Paper first received 22nd February and in revised form 5th September 1996 The authors are with the Department of Electric Power Engineering, Chalmers University of Technology, S-412 96 Goteborg, Sweden I", + " 2 Static principles, PV-curve and load Voltage stability is often studied by the system PVcurve and load characteristics. It is claimed [6, 11, 12, 141 that the upper side of the PV-curve is the only region where the system can operate and maintain its stability. The critical point, i.e. the point of maximum 258 power transfer, is considered to be the stability limit and the operation region on the lower side of the curve is considered to be unstable. Therefore, the stability limit on the PV-curve for some different loads are investigated. 2. 1 Measurements The line configuration in Fig. 1 is used, and the load is composed of an induction motor and a constant resistance. In all three cases in Fig. 3 the motor load is slowly increased until stalling occurs and the system collapses. Curve (a) shows the case where the load consists of only the induction motor. The motor stalls when it reaches its breakdown speed and the system collapses at the critical point. 0.1 0.2 0.3 0. L active power, p.u. PV curves or dflerent combinations of induction motor load and Fig.3 % Breakdown speed (stalling) constant resistance ", + "7 3 Automatic voltage regulators (AVRs) for large synchronous generators mostly have field current limiters to protect the generator field winding from overheating. The AVRs for the generators in the Swedish nuclear power stations also have armature current limiters thereby avoiding the situation where other protection relays trip the generator for overcurrent. Current limitation mostly implies a decreasing voltage, which generally is devastating for a stressed system. Voltage instability and a collapse may follow. To understand the function of the current limiters, the laboratory setup in Fig. 1 is used. 3.1 Field current limiter (FCL) In Fig. 8, curve 1 shows the result when the load is increased. The field current limit is set to 1 . 0 5 ~ ~ and when it is exceeded (a) there is a 5s delay before the limiter is activated (b). When the FCL becomes active the control changes so that the field current and thereby the excitation voltage is regulated instead of the terminal voltage. This means that the point of voltage control moves behind the machine synchronous reactance Xs, which now becomes a part of the network" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003928_iceee.2006.251859-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003928_iceee.2006.251859-Figure1-1.png", + "caption": "Figure 1. Wind sensor arragement", + "texts": [ + " The direction flow is considered based on the pattern of change of temperature that exhibits the arrangement. This system this developed one specifically for the autonomous navigation of a flying mobile robot, but it can be implemented for any system that needs to know such parameters of the wind in a restricted interval. The sensors arrangement has a forced heat transfer that is due to the fluid flow on a surface [5]. II. SYSTEM DESCRIPTION The wind sensor arrangement consist of LM35 temperature sensors LM35 placed in star form on a disk of 20 cm of diameter like it is shown in the figure 1. This sensors arrangement represents the axes cardinal north, south, east and west. If in a temperature sensor exist a constant temperature and this is exposed to an air flow that causes a variation of their temperature and then exists to a heat transfer that is manifested like a variation gives voltage in the temperature sensors. The system registers these variations of each sensor in a matrix to be analyzed and to be evaluated. III. MEASUREMENT OF THE WIND DIRECTION. A form of the measurement of air flow parameters consists on selecting of the disk the line with a maximum of temperature to evaluate the wind direction" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002866_095440603322769938-Figure2-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002866_095440603322769938-Figure2-1.png", + "caption": "Fig. 2 Normal section of a rack cutter", + "texts": [ + " As suggested by the computer aided manufacturing process, the solid model was employed to generate the cutting path of the gear and pinion geometrical model with ring-involute teeth. The results can be used to manufacture the proposed gear mechanism. F inally, a numerical example is presented in Table 1 to demonstrate the geometric model of a gear with a gear ratio of 3:2. F igure 1 shows the contours of gears with ring-involute teeth. This section discusses the design of a rack cutter with a ring-involute pro le. A normal section of a rack cutter is shown in Fig. 2. The rack cutter consists of two straight lines that form pressure angles fc with respect to the zn axis. To obtain the rack cutter surface, the normal section of the ring-involute rack cutter is attached to the plane yn \u00a1 zn, rotated along the zn axis and translated along the y1 axis with respect to the coordinate system S 1\u2026O1, x 1, y1, z1\u2020, as shown in F ig. 3. Thus, the normal section of the rack cutter surface can be represented by the coordinate system S 1\u2026O1, x 1, y1, z1\u2020 by applying the homogeneous coordinate transformation matrix method", + " Mechanical Engineering Science C02203 # IMechE 2003 at Purdue University on July 22, 2015pic.sagepub.comDownloaded from Thus, regions ab, bc and cd are only considered in the generating process of the gear with ring-involute teeth (convex). Similarly, regions cd , de and ef are only considered in the generating process of the pinion with ring-involute teeth (concave). Based on the description of this section, the equation for the ring-involute rack cutter can be described in terms of the S 1 coordinate system. Region ab as shown in Fig. 2 is used to generate the bottom land of a convex gear with ring-involute teeth. The variable `b is the parameter of the normal section of the rack cutter tooth\u2019s surfaces. In order to generate the complete pro le of the rack cutter tooth\u2019s surfaces, a tooth of the rack cutter will be repeated for cy \u02c6 0, 1, 2, . . . . According to Figs 2 and 3, the equation of region ab, represented in coordinate system S 1, can be written as Rab 1 \u02c6 xab 1 yab 1 zab 1 2 64 3 75 \u02c6 \u2026pm=2 \u00a1 `b\u2020 sin b \u2026pm=2 \u00a1 `b\u2020 cos b \u2021 cypm \u00a1 ac \u2021 r sin fc \u00a1 r 2 64 3 75 0 < `b < w, 0 < b < 2p \u20261\u2020 where w \u02c6 pm=2 \u00a1 bc \u00a1 ac tan fc \u00a1 r cos fc is the distance between each tooth, fc is the pressure angle and ac and bc are the design parameters of the rack cutter", + " The variable `h represents the design parameter of the rack cutter surface. The equation for region cd , represented in the coordinate system S 1, can be written as Rcd 1 \u02c6 x cd 1 ycd 1 zcd 1 2 64 3 75 \u02c6 \u2026bc \u00a1 `h sin fc\u2020 sin b \u2026bc \u00a1 `h sin fc\u2020 cos b \u2021 cypm `h cos fc 2 64 3 75 \u00a1 m1m2 < `h < m2m3 \u20263\u2020 where the parameters m1m2 and m2m3 denote the straight edge of the rack cutter. According to the geometry of the rack cutter, the form parameter can be determined by m1m2 \u02c6 ac=cos fc and m2m3 \u02c6 at=cos fc, where ac and at are design parameters shown in Fig. 2. Region de is the tooth head of the rack cutter tooth\u2019s surfaces and is used to generate the bottom land of the pinion (concave teeth). As shown in F igs 2 and 3, the equation of region de, represented in the coordinate system S 1, can be represented as Rde 1 \u02c6 xde 1 yde 1 zde 1 2 64 3 75 \u02c6 \u2026bc \u00a1 at tan fc \u00a1 r cos fc \u2021 r sin `d\u2020 sin b \u2026bc \u00a1 at tan fc \u00a1 r cos fc \u2021 r sin `d\u2020 cos b \u2021 cypm at \u2021 r sin fc \u00a1 r cos `d 2 64 3 75 0 < `d < p=2 \u00a1 fc, 0 < b < 2p \u20264\u2020 In equations (1) to (4), the position vector of the rack cutter with ring-involute teeth is Ri 1 and the superscript i indicates regions ab, bc, cd and de" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002129_105994901770345187-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002129_105994901770345187-Figure1-1.png", + "caption": "Fig. 1 Schematic of the T joint made from 6061-T6 flange and 5083- H321 web", + "texts": [ + " Contact e-mail: tsrivatsan@uakron.edu. maintaining a constant arc length. Variation in torch-to-work Journal of Materials Engineering and Performance Volume 10(2) April 2001\u2014173 Alloy Si Fe Cu Mn Mg Cr Zn Ti Al 6061 0.4 0.7 0.4 0.15 1.0 0.2 0.25 0.15 Bal 5083 0.4 0.4 0.1 0.4 4.0 0.10 0.25 0.15 Bal piece distance tends to cause significant changes in current with concomitant influence on penetration. The filler material used was wire stock of aluminum alloy 5356. A schematic of the T joint is exemplified in Fig. 1. The fusion zone microstructures were characterized by optical microscopy. Samples for metallography preparation were wet ground on progressively finer grades of silicon carbide impregnated emery paper, using copious amounts of water as the lubricant. Subsequently, the ground samples were mechanically polished using 5 and 1 m alumina-based polishing compound suspended in distilled water. Grain morphology and other intrinsic microstructural features were revealed using Keller\u2019s reagent (a solution mixture of hydrofluoric acid 1 nitric acid 1 hydrochloric acid 1 distilled water) as the etchant", + "[9] Because of the low solubility of the Mg2Si phase in pure aluminum at high magnesium contents, it is often present in the microstructure as the major constituent.[9] The magnesium in Fig. 3 Bright-field optical micrograph showing particle distributionsolution imparts limited solid solution strengthening to the alumiand grain morphology in aluminum alloy 5083-H321num alloy matrix besides facilitating additional strengthening through its influence on work hardening. The conjoint influence of magnesium in solution and the strain hardening arising from 6061-T6 flange and the 5083-H321 web (region 3 in Fig. 1). Both the 5083 web plate and the 6061 flange are in contactcold deformation are responsible for the acceptable strength of this alloy. Besides second-phase particles containing iron, manga- with each other and not in direct contact with the weld bead. The heat of welding was found to have a minor influence onnese and silicon are also present. The presence of manganese results in the precipitation of dispersoids (Al6(MnFe)) during the intrinsic microstructural features. Few of the grains revealed evidence of coarsening coupled with segregation of the coarseingot preheat and high-temperature homogenization treatments", + " This is confirmed by the evidence 3.2 Fusion Zone Microstructures of segregation, presence, and clustering of the second-phase particles at and along the grain boundaries of the 5083-H321-The microstructure of the welded region was examined for joints welded by the semiautomatic technique. Figure 4(a) alloy web plate. Such segregation and concomitant clustering are attributed to be the conjoint influence of heat generatedshows the microstructure at the interface of the 5083-H321 web plate and weld fusion line (region 1 in Fig. 1). A cast structure during welding and its influence on the kinetics of solidification. Furthermore, numerous fine microscopic cracks were observedis readily apparent at the weld bead. Grains of the parent 5083 material, immediately adjacent to the fusion line, show evidence along the grain boundaries of the 6061-T6 flange, both at and adjacent to the interface.of coarsening coupled with segregation of the coarse secondphase particles to the boundaries. The microstructure of the 6061-T6 flange immediately adja- cent to the weld toe (region 5 in Fig. 1) is shown in Fig.The microstructure at the intersection of the 5083-H321 web plate with the 6061-T6 flange (region 2 in Fig. 1) is shown in 5. Numerous fine and microscopic \u201chot-short\u201d cracks were observed. The presence and occurrence of the hot-short cracksFig. 4(b). The weld bead reveals a cast microstructure. The coarse and intermediate-size second-phase particles and other suggests the generation and concomitant influence of heat retention during semiautomatic welding. Furthermore, in the caseconstituents have segregated to the grain boundaries of the parent material, immediately adjacent to the weld fusion line. of this alloy, the thermal excursions in the HAZ will eradicate the prior thermo-mechanical processing (TMP) history, causingThe grains at and adjacent to the fusion line revealed evidence of coarsening", + " Welding Aluminum: Theory and Practice, Technical Advisory Panel on Welding and Joining, The Aluminum Association, Washington, DC, 1992. 3. \u201cWelding of ALCOA Aluminum Alloys,\u201d Internal Technical Report, Aluminum Company of America, Pittsburgh, PA, 1972. 4. R. Martukanitz: ASM Metals Handbook, vol. 6, Welding, Brazing andFig. 5 Bright-field optical micrograph of the fusion zone microstrucSoldering, ASM International, Materials Park, OH, 1993.ture adjacent to the weld toe on the 6061 flange (region 5 in Fig. 1) 5. P. Dickerson: ASM Metals Handbook, vol. 6, Welding, Brazing and Soldering, ASM International, Materials Park, OH, 1993. 6. M.L. Sharp: Behavior and Design of Aluminum Structures, McGrawautomatic technique of GMAW, the following observations can Hill, New York, NY, 1992. be made. 7. E.A. Starke, Jr.: Mater. Sci. Eng., 1977, vol. 29, pp. 99-115. 8. E.A. Starke, Jr.: in Aluminum Alloys: Contemporary Research and Applications, A.K. Vasudevan and R.H. Doherty, eds., Academic Press,\u2022 Grains of the parent material close to the fusion line show New York, NY, 1989, p" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003908_robio.2006.340236-Figure10-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003908_robio.2006.340236-Figure10-1.png", + "caption": "Fig. 10 Experimental system", + "texts": [ + " 8 shows the principle of the variable effective length spring. Effective length of spring H is changed by the length of rigid rod in the blade spring made of Poly(ethylene terephthalate), (PET). B. Propulsion Mechanism in Fluid Fig. 9 shows the fin with variable effective length spring. The fin consists of the aluminum box (length:120mm, variable effective length spring (length L:20mm, height:43mm, thickness:0.3mm). A DC motor, waterproof bearing, slipping screw, resin nut and rigid rod are housed in the aluminium box. Fig. 10 shows a picture of the experimental system. The mechanism is able to move on the linear guide. The pitching H ~~H zx ~ ~Q DQ PET plate Q (a) Apparent stiffiness (b) Apparent stiffness (c) Appairent stiffiiess Maximum (11-0) : Middle (O 7,) ay ay [ l l and the equation of flow takes the form shown below aP whereby q is the coefficient of viscosity, - ax the pressure gradient, h the film thickness and U I and U2 the velocities of the lower and the upper surfaces (Fig. 2). This equation is integrated with respect toy to get the volume flow Q and after performing the necessary substitutions, the expression for the velocity distribution can be written as Power Law Fluid In a power law fluid, the shear stress and the rate of deformation are expressed by the following relationship D ow nl oa de d by [ R M IT U ni ve rs ity ] at 2 2: 02 1 3 A ug us t 2 01 5 Velocity Meaurements in the Crease whereby K is the consistency coefficient and n the flow behavior index, which is quite significant in the determination of the fluid behavior. Its deviation from unity signifies a change in the Newtonian behavior ofthe lubricant. the fluid assuming a pseudoplastic behavior for values of n < 1 and a dilatant behavior for values of n > 1 as shown in Fig. 1. Equation 151 is combined with Eq. 121 to give By integrating Eq. [6] with respect toy. the velocity equation can be expressmi by: where A and B are constants of integration. the values of which could be obtained by using the boundary conditions (Fig. 2): However. with the above boundary conditions. analy~ical expressions for A and B cannot be obtained except for a value of n = I which corresponds to a solution of a Newtonian fluid, hence giving rise to a parabolic velocity distribution equation. APPARATUS The apparatus used is illustrated in Fig. 3. Two transparent cylinders of diameters 80 mm and 150 mm are in contact as shown schematically in Fig. 4 with a grease film napped in between. The outer cylinder is made of quenched glass material. 9 mm thick" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002341_cdc.1990.203964-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002341_cdc.1990.203964-Figure1-1.png", + "caption": "Figure 1: Constraint Surface Description.", + "texts": [ + " We examine these more closely in the section to follow. By duality or a quasi-static analysis using the Principle of virtual work, we determine the relations between the different forces which act on the system. We finally bring this together to derive the dynamic equation for the system in terms of the constraint parameters in the next section. For tasks involving fine motion of a grasped object along a physical surface we require that the contact point on the object does not change. Call this contact point p,, as shown in Figure 1. Following the notation and corresponding constraint surface description of Section two, let us fix the cosrdinate frame C, to the object a t the point p, , aligning its z-axis with the direction of the contact normal, and the x and y axes in the tangential plane. If the object must maintain contact with the surface during motion, the point p , on the object surface must follow the constraint surface. The position and orientation of the frame C, is specified by a vector in X, E R6. We parameterize the constraint surface by U E F, then equation 1 provides the relation between the two variables", + " Thus there is one degree of freedom in orientation and two degrees of freedom in position making a total of three free variables to completely specify the object position/orientation. Thus in this case r = 3. e If the contact point on the object surface is at a point of singularity, such as at the tip of a pencil, then the orientation of the object is free to vary in all directions and the set of parameters U must include the three parameters of orientation of the object. In this case, five variables are required to completely specify the position/orientation of the constraint frame with one variable for force, and the value of r is 5. Figure 1 illustrates the two cases described above. The vector [ h , . . . , SUrl specifies the position-controlled directions and the vector [XI , . . , A 6 4 specifies the force-controlled directions. Note :hat a differentiably smooth contact point has two additional forcecontrolled directions over a point of singularity. Consider now the kinematic constraints on the grasped object. Fix a coordinate frame CO to the object at its center of mass. Let the position of the frame CO relative to the base frame be given by 2, E 923" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003454_1.1757489-Figure4-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003454_1.1757489-Figure4-1.png", + "caption": "Fig. 4 Spacing between the slider and disk surfaces and locations for comparing different forces by two-dimensional and three-dimensional modeling", + "texts": [ + "org/ on 08/30/2017 Term ern negative pressure slider is chosen for this study ~Fig. 1!. The flying height of the slider is 26 nm. A sketch of the slider-disk assembly and coordinates is given in Fig. 2. The pressure profile shown in Fig. 3 is obtained without three-dimensional effects by solving the Reynolds equation for the air-bearing of the slider using the CML code Quick 4. To study particle flow in the air-bearing, we first calculated the spacing function between the slider and the disk. The slider-disk spacing map is shown in Fig. 4, where it can be observed that particles may enter the recessed region of the air-bearing through the leading edge of the slider. Next, the air streamlines at different levels of the slider/disk spacing were calculated using Eqs. ~5!\u2013 ~6! for heights above the disk of 1 percent, 50 percent, and 99 percent of the spacing as shown in Figs. 5~a!\u2013~c!. It is shown that the airflow close to the disk is mainly determined by the disk velocity and slider skew angle, while close to the slider surface the streamlines are mainly determined by the slip boundary conditions of the air", + " The drag force predicted by the three-dimensional model is about 4\u20137 orders of magnitude higher than that predicted by the two-dimensional model due to the transverse air flow in the different regions. The gravity force is 6E-13 micro Newton for a 30 nm particle. Also the lift force acting on the particle in Fig. 8 is between 1E-11 and 1E-9 micro Newton. To further understand the transverse flow Transactions of the ASME s of Use: http://www.asme.org/about-asme/terms-of-use Downloaded F effects, a comparison of drag force and Saffmann lift force at various locations was carried out for the two models. The locations used for this comparison are illustrated in Fig. 4. The results are presented in Table 1. It is shown that the drag force contributes significantly in the transverse flow model ~usually 2 orders higher than in the purely two-dimensional modeling!, with little contribution from the Saffmann lift force. Due to the drag force effect, the transverse flow analysis predicts that the particle\u2019s path is modified at the places of abrupt change in the slider-disk spacing. In this case, the drag force is the main cause of the vertical motion of the particle", + " For the three-dimensional modeling, few particles are observed on the leading pad, and more particles are observed in the transition region between the leading pad and the recess region according to the airflow analysis with transverse effects. Figures 11 and 12 show a comparison of the numerical simulation of a particle contamination profile with experiments for a particular slider design. One can see in both figures that many particles strike the front edge of the rails. These are large particles coming from the entrance. In the simulation results of Fig. 12, 30 nm, and 100 nm particles are used. They have the same density of Table 1 Comparison between different forces by twodimensional and three-dimensional modeling Locations in Fig. 4 Drag Force (1029 Micro Newton! Saffmann Lift Force (1029 Micro Newton! 2-D 3-D 2-D 3-D 1 20.001 20.5 0.003 0.005 2 0.014 4.0 0.014 0.2 3 0.1 16 0.25 0.02 4 20.2 2100 0.1 0.07 5 20.04 20.2 0.014 0.015 rom: http://tribology.asmedigitalcollection.asme.org/ on 08/30/2017 Term OCTOBER 2004, Vol. 126 \u00d5 749 s of Use: http://www.asme.org/about-asme/terms-of-use Downloaded F 4.253103 kg/m3. They are distributed uniformly on the disk surface in the entrance of the air bearing. Mostly the 100 nm particles are deposited in this region" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002711_a:1014049410772-Figure2-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002711_a:1014049410772-Figure2-1.png", + "caption": "Figure 2. Employed autonomous mobile robot, see text for specifications.", + "texts": [ + " Another super positioning method based on connectionist techniques is introduced in Brooks and Viola (1990). Interesting experimental results of an action fusion approach, where collision avoidance is being taken into consideration additionally, might also be found in Egerstedt et al. (1999). Another kind of superposition is demonstrated in Egerstedt et al. (1998) (\u2018virtual vehicle\u2019 control), where a geometrical path generation and following approach is combined with classical track control methods. The physical system employed for all experiments (Fig. 2) has a width of 0.5 m, a length of 0.8 m, a height of 0.6 m and a weight of 60 kg. It is equipped with the following sensor systems and actuators: \u2022 3-axis gyroscope: stability: \u22481\u25e6/s; sampling frequency: 176 Hz. \u2022 3 linear accelerometers: resolution: 5 mG; sampling frequency: 176 Hz. \u2022 2 encoders: resolution: \u224886000 ticks per wheel revolution; sampling frequency: 58 Hz. \u2022 4 wheel drive with differential steering, a maximal linear speed of \u22481.6 m/s and a maximal angular speed of \u2248150\u25e6/s; control frequency: 193 Hz" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003461_1.1809701-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003461_1.1809701-Figure1-1.png", + "caption": "Fig. 1. Stages of the form changing process in the arched strip: (a) setting the initial shape, (b) bending by applying force F in martensite, (c) recovery of the shape upon heating, and (d) recovery of the shape in the presence of gap \u2206.", + "texts": [ + " The aim of this study is to analyze the mechanical stability of an arched strip under constrained conditions [10, 11] in terms of the recently developed theory of bidirectional SME [9, 13] and to find the factors responsible for the click when the temperature is varied. The TiNi strips (the nickel content was 50.5 at. %) of thickness 2h = 0.4\u20130.5 mm, width b = 6\u20138 mm, and length 2l = 19\u201321 mm were mounted in a mandrel with a given radius of curvature R0 and a maximum bending deflection W0 = R0[1 \u2013 ], where L0 = R0sin(l/R0) is half the chord between the ends of the bent strip (Fig. 1a). The fixed strip was annealed at a temperature of 773 K for 30 min in order to give rise to the bidirectional shape memory effect. Further, the strip was rolled to \u03b5\u03a3 \u2248 35% with intermediate annealings and finally annealed at 693 K for 1.5 h. These procedures were aimed at improving the reversibility of the martensitic transformation. As a result, the starting and final temperatures of the direct and reverse martensitic transformations (Ms, Mf and As, Af, respectively) were 1 L0/R0( )2\u2013 004 MAIK \u201cNauka/Interperiodica\u201d 1302 MALYGIN, KHUSAINOV found to be Ms = 307 K, Mf = 279 K, As = 297 K, and Af = 325 K. Next, the arched strip was placed into a deforming machine with hinged grips (Fig. 1b) and bent by applying force F in the direction opposite to the initial bend at 293 K. The amount of the deflection in the opposite direction was equal to W\u03a3/2. In this case, the material of the strip is transferred to the martensitic state. Now, if the strip is heated with force F removed, it recovers the initial state, giving forth a clap (click) (Fig. 1c). The clap is heard when the strip\u2019s ends rest on the stationary hinged grips, and compressive force P appears at the ends upon straightening. This force additionally bends the strip, making its strained state unstable [13]. If one or both ends are free, the transition to the austenitic state and the form changing process occur steadily. Therefore, to stabilize these processes when the strip is in the hinged grips, a sufficiently wide gap \u2206 must be provided in order that one or both ends of the strip are free to move (Fig. 1d). Varying the width of the gap, one can control the constraint. Under the conditions of constrained form changing, the functionality of the arched strip expands. Figure 2 shows the dependence of static reaction force Qr that is exerted by obstacle B (dynamometer) placed at the center of the strip on the total bending deflection W = AB of the central part of the strip during heating. When the form changing process is constrained, the phenomenon of clap is of dynamic character and the strip acquires a kinetic energy in the course of form changing", + " In what follows, we analyze the mechanical behavior of the arched TiNi strip under the conditions of the TECHNICAL PHYSICS Vol. 49 No. 10 2004 constrained SME. The aim is to find the geometry of the strip that favors the effect of click. With allowance for the internal bending moment (which is due to the anisotropic Ti3Ni4 particle distribution over the thickness of the strip) and external bending moment (due to compressive force P at the ends of the constrained strip), the total curvature R\u20131 of the strip (R is the radius of curvature) is given by (1) Here, is the curvature of the strip upon annealing (Fig. 1a); = \u2013(3/4h)|\u03b50| and = \u2013(3/2h) (x, t) are the changes in the curvature due to the elastic, \u03b50 [9], and martensitic, (x, T) [13], strains arising in the strip after the relaxation of elastic bending stresses, respectively; T is the temperature; and x the coordinate measured from the center of the spanning chord (Fig. 1a). The last term on the right of Eq. (1), (x, T) = \u2013M(x, T)/EJ, is the elastic contribution to the curvature of the strip because of the constrained motion of its ends. Here, M(x, T) = W0(x, T)P(T) is the bending moment and W0(x, T) is the bending deflection of the strip in the case of the unconstrained SME. We have (2) In (2) and above, R0(T) and L0(T) are the radius of curvature and half the distance between the strip\u2019s ends when the temperature varies under the conditions of the unconstrained SME, respectively [9, 13]; P(T) = EA\u03b5(T) is the longitudinal compressive force at the strip\u2019s ends (Fig. 1c); E is the modulus of elasticity; A = 2hb is the cross-sectional area of the strip; \u03b5(T) = [l0 \u2013 L0(T)]/l is the longitudinal compressive strain in the strip; 2l0 is the free spacing between the grips, which is related to gap \u2206 (Fig. 1d); and J = b(2h)3/12 is the moment of inertia of the strip\u2019s cross section. The averaged martensitic deformations (T) and (x, T) (under the conditions of the unconstrained and constrained SMEs, respectively) are determined when R 1\u2013 x T,( ) R0 1\u2013 Re 1\u2013 Rm 1\u2013 x T,( ) RP 1\u2013 x T,( ).+ + += R0 1\u2013 Re 1\u2013 Rm 1\u2013 \u03b5\u0303\u0303m \u03b5\u0303\u0303m RP 1\u2013 W0 x T,( ) R0 T( ) 1 x R0 T( ) -------------- 2 \u2013 1 L0 T( ) R0 T( ) -------------- 2 \u2013\u2013 ,= R0 T( ) R0 1 3R0 2h -------- 1 2 -- \u03b50 \u03b5m T( )+ \u2013 1\u2013 ,= L0 T( ) R0 T( ) l R0 T( ) --------------", + " Under the given conditions and parameters of the strip, the maximum sag at its central part is \u00b10.8 mm. The transition of the strip to the 1 \u03c92\u2013 W x T,( ) \u03c9 x T,( ) xd 1 \u03c92 x T,( )\u2013 ---------------------------------. L0 T( )\u2013 x \u222b= R0 R x T,( ) ----------------- 1 3R0 2h -------- 1 2 -- \u03b50 \u03b5\u0303\u0303m x T,( )+ \u2013= + 3 R0 h ----- 2W0 x T,( ) R0 --------------------- \u03b5 T( ) . \u03b5 TECHNICAL PHYSICS Vol. 49 No. 10 2004 fully martensitic state and the change of sign of the sag (point b) by applying force F (Fig. 1b) to the strip at a temperature of 0.9TR (point a) are shown by arrow ab in Fig. 3b. From comparison of the curves in Figs. 3a and 3b, it follows that, in the reduced coordinates, the bending deflection W0(0, T) = W0(T) at the central part of the strip varies with temperature in the same way as the reduced curvature and that these values quantitatively coincide. This coincidence stems from the fact that (i) the deflection of the strip is smaller than its length (\u22480.5W0/l \u2248 0.88) and (ii) the curvature of the strip is also smaller than its length both in the initial (after the annealing) state (l/R0 \u2248 0", + " The free grip spacing is given by the expression 2l0 = 2L0(Ta) + \u2206, where Ta is the temperature at which the R0 1\u2013 R0 1\u2013 W0 T( ) W0 --------------- L0 2 T( ) 2W0R0 T( ) -------------------------\u2248 = R0 T( ) 4R0 -------------- l R0 T( ) -------------- sin 2 l 2R0 ) ---------- sin 2 ----------------------------------- R0 R0 T( ) --------------.\u2248 \u03b5\u0303\u0303m TECHNICAL PHYSICS Vol. 49 No. 10 2004 strip is mounted in the grips and \u2206 is the total gap between the strip\u2019s ends and grips at this temperature (Fig. 1d). For the maximal compressive strain \u03b5(T2) of the strip (see above) and the temperature at which the strip is mounted in the grips, Ta = 1.1TR, gap \u2206 equals 0.2 mm and the compressive (bending) force at the strip\u2019s ends (Fig. 1c) is P = EA\u03b5(T2) \u2248 100N, where E = 100 GPa. When the gap is narrow or is absent at all, the instability of the form changing process builds up, while for wide gaps, it weakens or even completely disappears. The process becomes unstable when the gap is smaller than the critical value, \u2206 < \u2206c, where (11)\u2206c 2L0 T1 2,( ) 2L0 Ta( )\u2013 2 l L0 Ta( )\u2013[ ].\u2248= Figures 6a and 6b demonstrate (according to (2) and (11)) how critical gap \u2206c varies with the initial radius of curvature R0 of the strip in the reduced coordinates \u2206c/2l \u2013 R0/h for l/h = 40" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003633_1.1645294-Figure2-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003633_1.1645294-Figure2-1.png", + "caption": "Fig. 2 Eccentric seal kinematic model and coordinate systems", + "texts": [ + " In most mechanical seals, one sealing face is a thin annular ring and runs against a flat mating face, referred to respectively as the seal ring and the seal seat. The faces are separated by a thin fluid film, and coning and tilt of the faces generate pressures in the film which apply forces and moments to the two seal faces, coupling the dynamic responses of the two rotors. In this work, the seal ring is designated rotor 1, and the seat, rotor 2. They are assumed to be oriented with respect to the system Z axis as shown in Fig. 2. For concentric seals, the forcing functions in the equations of motion result from initial tilts of the rotors, such as occur because of imperfections in the flexible support @13#. These same tilts also lead to forcing functions in the eccentric system, but additional forcing caused by the radial shaft motion must also be included. Thus, the kinematic model of the eccentric seal must be able to describe the relationship between the face tilts and the shaft deflections. Wileman and Green @10# presented the general kinematic model for an eccentric FMRR system. Their definitions for the rotor eccentricities and relative eccentricity are illustrated in Fig. 2. Shaft 1 and shaft 2 have radial deflections from the nominal system centerline, denoted e1* and e2* , respectively. The relative deflection of the two shafts, e*, is obtained as a vector difference ~Fig. 3!. The asterisks indicate these are dimensional values. The normalized eccentricity, e, is defined as e5e*/ro , where ro is the outer radius of the seal ring. Fig. 1 Mechanical face seal with two flexibly mounted rotors \u201eFMRR\u2026 302 \u00d5 Vol. 126, APRIL 2004 rom: http://tribology.asmedigitalcollection" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002720_robot.1992.220099-Figure6-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002720_robot.1992.220099-Figure6-1.png", + "caption": "Figure 6: Probability of collision for second example, front view.", + "texts": [], + "surrounding_texts": [ + "We consider the derivation of the conditional probability of collision, conditioned on the past and present observations. The derivation and calculation is done for increasing degrees of complexity. We investigate the case of errors only on the position component X , of S,, then the more general case of errors in U, (translation errors in position, speed and acceleration), and finally the global case of errors in all components of the kinematic state S, (translation and rotation). This paper present only the simple case of errors in position. Detailed derivations for the other cases are given in [3]. In this case the errors in translation velocity and orientation can be considered negligible compared to the position error on X , . This corresponds to setting the evolution model translation errors V,\" and rotation errors v,\" to (3) In the target frame the vector DL, i.e. the relative position of the robot with respect to the target is given by the change of reference DL = R-a,(Dn - X n ) (4) where R , denotes the rotation matrix of angle a. Dn and a, being fixed, the random vector DL depends only on X , . We may simply write DL = L1(Xn) . Using the previous equation along the target m e tion model we show that the uncertainty at present time on the position of the robot in the target reference and the uncertainty on the motion model evolution are carried to the relative position of the robot at the next time instant, as where: (5) 2442 x, ~ h ( D h ) = R(a,,-an+l)DL+R-an+l ( D n + l - D n - X n - - ) ~53(Vn) = -R-an+I V,X* (7) D i and V: are independent since D i depends only on X, and as a result of the stochastic model embedded in the Kalman filter formulation, X, and V,\" are independent random'vectors. V,\" has a Gaussian d i s tribution, i.e. f\"$(VX) = +l,r:,(vX) with T: given by the probabilistic model. As discussed before in section 2.2, X, is Gaussian. Therefore, D i being an affine function of X, is also Gaussian fD: (D') = N ( E [ D : ] , C f ' ) ( D f ) with mean E[Di ] and variance E!' equal to (9) D' X E n = R-anCn Run where X,,, and E: are output of the Kalman Filter. We now express the probability of collision. We have a collision event for all pairs (DL,DL+l) for which D&+, belongs to the shadow of the target with respect to DL. By (5) it is equivalent to consider all pairs (DL, V:) such that V,\" E D ( D L ) where the domain D(DL) is given by 'D(Di) = { v: : LZ(Di) + L3(V,x) E Sh(DE)} * (10) Finally, using (10) and independence of the vectors, PC( Dn Dn + 1 ) f v2 (VX)dVx dDf (11) V V x eV(Dr) 1 = l v D f D : ( D ' ) (1 Computation is done as follows. The first step consists in the determination of the shadow with respect to a given D i , which can be done beforehand by calculation of the visibility graph of the polygonal target. The above integral is numerically approximated but some of the burden in this computation can be alleviated by direct analytical solution of the innermost integral (basically the integral of a Gaussian over a polygonal region). Alternatively a Monte Carlo method can be used on a SIMD machine to obtain a fast approximation. We give here examples of typical probability of collision profiles for different situations and error values in the case of position error, calculated for a grid of candidate destinations. The shape of the target used is described in Figure 2. The setup is as follows (all vectors are given in the global coordinates system.) The target center of inertia estimated position X,, its covariance E: are assumed to be x, = [ 0\":] ; E; = (12) The target is first assumed immobile: [ ;: ]x, = x, = The current and next target orientation angle: The target position error covariance: 2443 The robot current position is and plans to reach inside the target concavity. For conditions stated above we give the resulting probability of collision associated with a set of candidate robot destinations. This set of candidate destinations is given on a 25 samples grid centered on the position Dn+l = [-0.5,0.0] and of width equal to 1.00. The plot of the probability of collisions is given in Figures 4 a n d 3 . A second example is examined under the same conditions but the target is given now a rotational movement during the time interval of :. The corresponding results are given in Figures 6 and 5. 4 Optimal Destination The type of robotic operation to be accomplished on the target defines a subset of operational s ta t e s AO, comprising the set of states (relative position or kinematic states for instance) for which this operation can be successfully carried out. We measure in a probabilistic sense how close we are to the operational state and incorporate this with the probability of collision in a general cost function. For simplicity, AOcould consist of the positions in the target frame of reference where the robot can.act on the target. The destination Dn+l is operational if DLtl E AO; the conditional probability of being operational is The optimal solution of the problem of trajectory planning defined previously is the destination Dn+l that minimizes a composite cost function as 2444 References Di+1= Argmin(yPC(Dn,Dn+1) + (1 - y)( l - Po(Dn+l))) VDn+1 E Cn+l [l] A.Basu. A framework for motion planning in the presence of moving obstacles. Technical Report CS-TR-2378, Center for Automation Research, University of Maryland, 1989. with 7 E [0,1], where vu 1 (15) The search space Cn+l is reduced to those destinations satisfying the dynamic constraints on the robot. The search for a suboptimal solution (for example a destination meeting an acceptable level of collision probability) might be preferable to balance the cost associated with the computation of the probability of collision by using a scheme that interleaves search and calculation. 5 Conclusion For many robotic applications (robotic satellite maintenance, autonomous vehicle or industrial robot guidance, rotorcraft navigation.. ), the dynamic risk assessment is essential. While the risk is interpreted here as collision risk, we can extend this notion to other hazards that the robotic system might encounter which can be expressed in terms of the relative states of the robot with respect to a target or obstacle. This paper has studied an approach to probabilistic navigation from sensors in dynamic environments considering as optimality criteria the probability of colliding with an obstacle and the probability of accessing an operational state. We describe a computationalframework in which a probability of collision can be effectively derived. This probabilistic description is appropriate for its use with classical decision theoretic tools and can be integrated in a low level trajectory controller that operates with a higher level planner; the high level planner would consider navigational subgoals such as moving toward a position in space, locking in translation or rotation with a target object, going into orbit around a moving object,.. and provide associated levels of acceptable risk. (I4) [2] C.K.Yap. Algorithmic motion planning. In J.T. Schwartz and C.K. Yap, editors, Advances in Robotics vol. 1: Algorithmic and Geometric Aspects. Lawrence Erlbaum, Hillsdale, N.J., 1986. [3] P.Burlina D.Dementhon and L.S.Davis. Navigation with uncertainty: I. reaching a goal in a high risk region. Technical Report 565, Center for Automation Research, University of Maryland, June 1991. [4] G.S.Young and R.Chellappa. 3d motion estimation using a sequence of noisy stereo images: Models, estimation, and uniqueness results. IEEE Trans. PA MI, 12( 8) :735-759, August 1990. [5] T.H.Wu G.S.Young and R.Chellappa. A simple kinematic model based approach for 3d motion and structure estimation. Technical Report 2755, Center for Automation Research, University of Maryland, September 1991. [6] H.P.Moravec. A bayesian method for certainty grids. Technical report, Robotics Institute, Carnegie Mellon University, 1988. [7] J.Borenstein and Y.Koren. Real time obstacle avoidance for fast mobile robots in cluttered environments. In Proc. IEEE Conf. Robotics and Automation, pages 572-77, 1990. [8] J.T.Schwartz and C.K.Yap. Algorithmic and Geometric Aspects of Robotics. Laurence Erlbaum Associates Publishers, 1987. [9] N.C.Griswold and J.Eem. Control for m e bile robots in the presence of moving objects. IEEE Transactions on Robotics and Automation, 6(2):263-68,1990. [lo] 0.Khatib. Real time obstacle avoidance for manipulators and mobile robots. International Journal of Robotics Research, 1(5):90-99, 1986. [ll] T.J.Broida S.Chandrashekhar and R.Chellappa. Recursive 3d motion estimation from a monoeillar camera sequence. IEE Trans. Aero. and Elect. S ~ S . , 26(4):639-655, July 1990. 244.5" + ] + }, + { + "image_filename": "designv11_65_0002529_00037-1-Figure13-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002529_00037-1-Figure13-1.png", + "caption": "Figure 13 Typical flat cruciform biaxial specimen.", + "texts": [ + " A flat plate specimen is usually tested in a cruciform (cross-shaped) configuration. Any combination of tension and compression loading can be applied to the pairs of arms of the cruciform. The test region (gage section) is the central portion at the junction of the four arms. But stress concentrations are induced at these junctions (internal corners). Thus, considerable analytical effort has been expended to define optimum geometries that will minimize these local stress concentrations. Usually, circular cutouts of some shape such as shown in Figure 13 are used. The usual data presentation format in any case is a two-dimensional plot of axial normal (s1) stress vs. transverse normal (s2) stress. If data for all four quadrants of the plot can be generated (i.e., if external pressurization can be attained), the plot is an enclosed area, the perimeter of which at any combination of s1 and s2 defines failure. 5.06.9.2 Triaxial Testing As stated in Section 5.06.9.1, very few experimental results for triaxial loading have been generated to date. As for biaxial loading, either a tube or a cruciform specimen can be used" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000529_41.778253-Figure2-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000529_41.778253-Figure2-1.png", + "caption": "Fig. 2. Phasor representation of stator and rotor fluxes relative to an orthogonal phase reference frame.", + "texts": [ + " As a result of the sinusoidal relationship between torque and load angle, when the motor is operating near the maximum load angle (close to stalling), the slope of the torque/loadangle characteristic approaches zero. This means that a small variation in load torque can cause a large variation in load angle, implying that very little spare torque is available in this region. Consequently, a real motor will need to work with a minimum operating load angle of rather less than /2 to allow it to cope with slight disturbances in load torque without stalling. Fig. 2 shows a phasor diagram for both phases of the motor in steady-state operation, illustrating the relationship between fluxes and rotational EMF relative to a fixed reference frame for phases A and B. As the rotor moves, it induces rotational EMF\u2019s into each of the phase windings as the component of rotor flux in each phase winding changes with time. The rotational EMF in each phase is given by the following applications of Faraday\u2019s law: (7) (8) where rotational EMF induced in phase A (V); rotational EMF induced in phase B (V); number of turns in each phase (same for both phases)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002014_elan.1140010606-Figure7-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002014_elan.1140010606-Figure7-1.png", + "caption": "FIGURE 7. CV curves at the L/L interface with a mixed solvent of o-DCB plus BC in different volume ratios as the Organic phase. W, 0.01 M LiCI; 0, 0.01 M TBATPB. o-DCB : BC: 1, 8 : 2, 2, 6 : 4; 3, 4 : 6; 4, 2 : 8. Scan rate, 40 mV/s; A:Q vs. TBA+ISE.", + "texts": [ + " Figure 6 shows the CV curves of the mixed-solvent systems W/PIT plus BC and W/PIT plus BN (0.01 M LiCl in the aqueous phase and 0.01 M TBAWB in the organic phase) with different volume ratios; these suggest that mixed solvents of PIT with BC or BN in volume ratios of 8 : 2-6 : 4 or 8 : 2-4 : 6 are good for electrochemical measurements at the LIL interface. Increasing BC and BN results in a narrowing of the potential window. Table 4 shows the physical properties of mixed solvents and the values of potential window. WIo-DCB-BC. Figure 7 shows the CV behavior of the WIo-DCB plus BC mixed-solvent system with 0.01 M LiCl in the aqueous phase and 0.01 M TBATPB in the organic phase. Good transfer behavior can be obtained for systems with o-DCB-BC volume ratios of ratios X : 2, 6 : 4 , 4 : 6 , and 2:8. W/CB-BN. The CB plus HN mixture was also found to be a good system for electrochemical measurements at the L/L interface. Figure 8 shows the CV curves of the CB plus BN mixed solvent with different volume ratios; these suggest that reasonable CR : BN volume ratios are 8 : 2-4 : 6" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001951_s0389-4304(01)00102-3-Figure11-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001951_s0389-4304(01)00102-3-Figure11-1.png", + "caption": "Fig. 11. Front view of the linkage layout.", + "texts": [ + " In contrast, because the coupled linkages resulted in nearly uniform movement of the four trunnions, variation in the torque share of the power rollers was reduced. These test results thus confirmed that the use of coupled upper and lower linkages plays an important role in synchronizing the movement of the four power rollers. The dual-cavity half-toroidal CVT was applied to a rear-wheel-drive car fitted with a 3.0-l turbocharged engine. A key objective set for the toroidal CVT layout was to integrate the construction of the upper and lower linkages that support and position the four trunnions. As shown in Fig. 11, the output gear housing and the casing shape were specially designed to create sufficient space for coupling the upper linkages of the fore and aft variators. The large-capacity half-toroidal CVT described here adopts coupled upper and lower linkages for positioning the trunnions that support the power rollers. As a result, thrust forces acting on the power rollers are now cancelled out by the connecting members instead of being accommodated by the linkage supports, thereby improving variator performance" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001124_02783649922066574-Figure7-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001124_02783649922066574-Figure7-1.png", + "caption": "Fig. 7. Manipulator-fixed reference systems.", + "texts": [ + " In general, if the desired three-dimensional terminal location of the target points is defined with respect to a coordinate system fixed on the stationary nonmanipulable body, the intermediate three-dimensional positions of the target points, which constitute the approach trajectory, can be defined by means of a 4 \u00d7 4 transformation matrix Di for the ith camera as follows: t \u2032ixj t \u2032iyj t \u2032izj 1 = Di t ixj t iyj t izj 1 (5) Primed coordinates (t \u2032ixj , t \u2032i yj , t \u2032i zj ) represent the location of the j th target point at an intermediate juncture that is part of the approach trajectory, whereas (t ixj , t i yj , t i zj ) represent the desired location of the same target point as defined with respect to a coordinate system attached to the nonmanipulable body (see Fig. 7). Matrix D has been used in the past (Skaar, Brockman, and Jang 1990) to define the trajectory of the manipulated body toward the workpiece. The present paper introduces a complementary way of defining such an approach at UNIV OF ILLINOIS URBANA on March 10, 2015ijr.sagepub.comDownloaded from trajectory using another 4 \u00d7 4 matrix, called 0, as explained below. Consider now Figure 1, which depicts the general case in which a large object is moved by a robot. As explained above, several cameras are pointed toward different regions of the workspace", + " To obtain the required matrix of moments, the nominal physical location in the previous two expressions can be written in terms of the desired target points (tx i j , ty i j , tz i j ) by using the nominal forwardkinematic model of the manipulator and the known geometry of the grasped object. Without loss of generality, we can consider two independent sets of coordinate systems on the manipulable and the nonmanipulable bodies defined in such a way that they coincide when the task is finished. An example of this is the wheelloading task, which is represented schematically in Figure 7. A procedure for guiding the wheel toward the brake plate (the nonmanipulable body) where it will be assembled involves the definition of a coordinate system XNYNZN on the nonmanipulable body, which will coincide with theXT YT ZT coordinate system on the wheel when the task is complete. An approach trajectory is defined using the XINTYINTZINT reference frame, as depicted in Figure 7. It is possible to obtain a transformation matrix TK(2) (see, for example, the works of Denavit and Hartenberg (1955) and Paul, Shimano, and Mayer (1981)) that relates a coordinate system attached to the robot\u2014stationary with respect to the base, but not necessarily at the base\u2014(X0, Y0, Z0), to the tool coordinate system (XT , YT , ZT ) attached to the manipulated body. This transformation constitutes the forward-kinematic model of the manipulator. Further transformation matrices 0i , defined independently for the ith camera, allow us to define intermediate coordinate systems, XINTYINTZINT, with respect to the tool coordinate system (XT , YT , ZT ), as shown in Figure 7. Then, the following relationship can be obtained, for the j th destination point of the ith camera: at UNIV OF ILLINOIS URBANA on March 10, 2015ijr.sagepub.comDownloaded from ri xj (2) ri yj (2) ri zj (2) 1 = TK(2)0i t ixj t iyj t izj 1 = Ti (2) t ixj t iyj t izj 1 (11) The introduction of this transformation into eqs. (9) and (10) leads to the following expression in matrix form: [ fx(. . . ) fy(. . . ) ] = [ t ixj t iyj t izj 1 0 0 0 0 0 0 0 0 t ixj t iyj t izj 1 ] TiT ... 0 \u00b7 \u00b7 \u00b7 ", + " (31) are the eigenvectors of the 3 \u00d7 3 matrix of central moments Mcm: Mcm = \u2211 (txj \u2212 tx) 2Wj \u2211 (txj \u2212 tx)(tyj \u2212 ty)Wj\u2211 (tyj \u2212 ty) 2Wj Symmetric\u2211 (txj \u2212 tx)(tzj \u2212 tz)Wj\u2211 (tyj \u2212 ty)(tzj \u2212 tz)Wj\u2211 (tzj \u2212 tz) 2Wj . (33) This process is similar to the one used to define the principal moments of inertia and principal axis of the inertia ellipsoid of a rigid body, in the context of rigid-body dynamics, as explained by Meirovitch (1970). Following these ideas, our transformation matrix L defines a transformation between the original coordinate system XNYNZN in Figure 7, in which the set of nt target points (txj , tyj , tzj ) for j = 1, . . . , nt are defined, and a coordinate system whose origin is located at the center of gravity of these target points and is aligned along the principal axis for which the elements off the main diagonal of matrix Mcm in eq. (33) are zero. We refer to this coordinate system as system XpaYpaZpa. Consider the orthographic projection of four points (0, 0, 0), ( \u221a \u03bb1, 0, 0), (0, \u221a \u03bb2, 0), and (0, 0, \u221a \u03bb3), where \u03bbi , i = 1, 2, 3, represents the eigenvalues of matrix Mcm, to each participating camera" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001285_1.2833938-Figure8-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001285_1.2833938-Figure8-1.png", + "caption": "Fig. 8 Experimental apparatus", + "texts": [ + " Figure 7 shows the relationship between time, t, and the shaft displacement caused by the step load, Ae,, comparing the proposed bearing with the conventional bearing with capillary restrictors. Since the proposed bearing has nearly infinite static stiffness, Ae^ is recovered to the initial position shortly after the step load is imposed. But in the conventional bearing, Ae, is not recovered to the initial position because of lower static stiffness. The maximum value of Ae^ caused by the step load of the proposed bearing is about a half of that of the conven tional bearing though the settling time of both bearings shows the almost same value. 4 Comparison With Experimental Results Figure 8 shows the experimental apparatus for obtaining the step response characteristics of the proposed bearing. The shaft is vertically fixed on the base table and the test bearing is pressed into the housing which is supported by an aerostatic thrust bearing. Furthermore, two identical aerostatic thrust bear ings are set on the both sides of the housing in line with the direction of the step load to prevent the housing from rotating. Hence, the housing can freely move only in the direction of the step load" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003065_robot.2002.1014226-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003065_robot.2002.1014226-Figure1-1.png", + "caption": "Figure 1: Camera setting and coordinate frames", + "texts": [ + " The inverse of this matrix is denoted by WIZB. * 2.2 Geometry of the System coordinate frames that the vergence is perfromed - a positive d u e makes the right and left cameras turn towards each other. Additionally, near the two camera mounting points we defined the cameras coordinate frames 1 and r . Finally, the calibration pattern, which defines the world coordinate system W , is placed in front of the stereo head. Our problem formulation starts by creating three additional frames: b, el, and e,., as shown on the right in Fig. 1. These arbitrary frames are necessary to determine the neck-eye calibration independently of the pan, tilt and vergence angles. In this instance, the coordinate frame b is a coordinate frame with the same origin as B. However, while B is fixed, the coordinate frame b rotates according to pan or tilt angles, following the motion of the stereo head. In that way, the HTMs Hb and ErHt, are always constant regardless of the pan and tilt angles and are the first pair of neck-eye matrices that need to be calibrated (Sec", + " 2.3 Calibration Equations Without loss of generality, we will formalize our calibration problem for one single camera. Therefore, we will drop some of the subscripts above and use the symbols E and c for the eye coordinate frame and the camera coordinate frame respectively. In summary, our coordinate frames are now: B, b, E , e, c, and w. Before formulating the problem, it is important to explain the geometry of our stereo head and the choices of all necessary coordinate frames. On the left in Fig. 1 are depicted the camera and stereo head settings. As 2-3-1 Calibration of \u201cHe the reader can observe, the stereo head can move with the neck base, a coordinate frame denoted B is defined. It is around the B coordinate frame that both an and The calibration of the camera with respect to its lowing steps: three degrees Of freedom: P W tilt and vergence. At mounting point, C H e , is obtained by taking the fol- tilt are performed. Next, at the top-left and top-right of B, two additional coordinate frames, denoted E1 and E," + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002006_1.528051-Figure5-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002006_1.528051-Figure5-1.png", + "caption": "FIG. 5. Extension of the chart Yo in the presence of boundary.", + "texts": [ + "4 ) h=' which derives from (2.1 ~ All spin structures on l:g,n have the same S~n(2) bundleF, which is the restriction to l:g,n of the bundle FofRef. 1. In place of4. 12-4. 16 in Ref. 1 we now have bundle morphisms TJlJk: F ..... F defined by TJlJk(e(x\u00bb) = e(x)rijk (x) , (3.5) where r IJk: U ..... SO (2) are such that composed with the loops aA,bA,anddh (regarded as mapsS ' ..... l:g) they have wind ing numbers iA' jA, and kh, respectively. This can be achieved by placing the disksD\" ... ,Dn inZoifg = o and Y, if g> 1 (see Fig. 5, where Y, has been slightly extended) and modifying the functions rlJ on the shaded strips in such a way that crossing the strip between Dh and Dh + \" the function rlJk rotates by 21Tqh with qh = l:~ =, k m \u2022 Every spin structure (ijk) extends uniquely to a 583 J. Math. Phys., Vol. 29, No.3, March 1988 Pin+(2) and a Pin-(2) structure, both of which will be denoted again (ijk). In particular, having fixed the structure group, the pin bundle is the same for all pin structures. The transformation of pin structures under 0 (l:g,n ) can be determined counting the winding numbers of the three terms on the rhs of (4" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002437_6.2001-4427-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002437_6.2001-4427-Figure1-1.png", + "caption": "Fig. 1 Missile coordinate systems for initial turn guidance", + "texts": [ + " During the agile turn phase, the missile is governed by an implicit guidance law. The nominal steering commands are generated from the optimal side jet on-off programs of the open loop time optimal trajectories for a rigid body model in the atmospheric flight. The optimal control problem is defined to transfer the missile attitude from vertical to a given final state in minimum time by using the forward SJC device. The equations of motion for the mathematical model of the missile flight in the vertical plane are as follows (see Fig. 1): u . N . Avv =- \u2014 sm#- \u2014 sm0-\u2014cos# m m m u \u201e N A .v = \u2014 cos#4-\u2014cost ) -\u2014smt t -g m m m a = - (1) where x is the down range of the missile in m, y the altitude in m, vx the missile velocity component along the x coordinate in m/s, v the missile velocity component along the y coordinate in m/s, a the angle of attack of the missile in degree, m the mass of the missile in kg, 6 the attitude of the missile body axis in degree, / the thruster location w.r.t. the center of gravity in m, / the moment of inertia in kgm2 , g the acceleration of gravity in m/s2", + "5 0 , A. ^\u2022^^ 0 0.5 1 1.5 2 2.5 3 3.5 4 Time to go (sec) Fig. 5 Effect of different timing to initiate the terminal SJC for case 1 M iss d ist an ce 0 \u2014 M V 0.5 1 1.5 2 25 3 35 4 Time to go (sec) Fig. 7 Effect of different timing to initiate the terminal SJC for case 3. 10 u g 6 1 4 .2 2 0 ( ^SSN. ^x. \\^ ) 0.5 1 1.5 2 2.5 3 3.5 Time to go (sec) Fig. 9 Effect of different timing to initiate the terminal SJC for case 5. 1.4 ? 12 1 0.8 \u00a3 \u00b06 .a \u00b04 i 02 0 S. Xv \\^ 0 0.5 1 1.5 2 2.5 3 3.5 Time to go (sec) Fig. 1 1 Effect of different timing to initiate the terminal SJC for case 7. 7.5 ? 2 1 '\u2022' Jo , 0 V \u2022 N. ^\\^ 0 1 2 Time to go (sec) 3 Fig. 6 Effect of different timing to initiate the terminal SJC for case 2. 4 I 3 1 1 0 0 \\ v 0.5 1 1.5 2 2.5 3 3.5 Time to go (sec) Fig 8 Effect of different timing to initiate the terminal SJC for case 4 14 \u2022g- 12 ^ 10 i \u2022 4i 6 g 4 S 2 0 \\ L 0 1 2 3 4 5 6 7 Time to go (sec) 8 Fig. 10 Effect of different timing to initiate the terminal SJC for case 6 5 1 4 g 3 .2 0 0 \\ 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000537_s0263574700015411-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000537_s0263574700015411-Figure1-1.png", + "caption": "Fig. 1. W-link serial manipulator and relationship of defined quantities to link k.", + "texts": [ + " In section V, the simulation results developed by the author are presented. The results of this work are summarized and some conclusions are given in the final section. II. NOTATION FOR SPATIAL DYNAMICS In this paper we have assumed that a mechanical system of n links is numbered in an increasing order which goes from the tip of the system toward the base. Joint k in the sequence connects links k and k +1. The fixed base is considered to be link n + 1. At any point at the tip of the manipulator a fictitious joint 0 is attached. The ordering system is shown in Figure 1. The robotic system depicted in Figure 1 can contain only simple revolute and/or prismatic joints. Coordinate systems can be assigned according to a modified Denavit-Hartenberg convention but in general, the equations which will be derived do not depend on the specific coordinate transformation used. The formulation of all equations in this paper is carried out using spatial notation. We have used the notation introduced by Rodriguez14 (Figure 1). The spatial velocity, acceleration and force are all 6 x 1 column vectors. A more detailed definition of these concepts is provided below. Tk and F* are 3 x 1 vectors representing, respectively, the constraint torque and force acting at joint k on link k + 1. The corresponding spatial force is defined to be a 6 x 1 vector and is denoted by xk = [-T* -Fk] T, where T is a transposition operation. Superscript + indicates that the variable is calculated at a point on the link k +1 adjacent to joint k", + " In general, the orientation of a spatial kinematic vector from one coordinate system to an adjacent one may be accomplished by making use of the following spatial orientation matrix (6 x 6) A - where k+ kR is the 3 x 3 rotation transformation between two coordinate systems. Some authors5 combine these two matrices into one spatial transformation matrix. Rodriguez has defined the transition matrix in order to formulate the equations of motion using the discrete linear state space approach. To continue with the symbol definitions presented in Figure 1, A* is a unit vector along joint axis k (not necessarily the zk axis), Ck denotes the mass center of link k, and qk is the joint variable which is positive in the right-hand sense about hk joint axis. Next, Hi is a 6 x 1 vector of the following form: when joint k is rotational and when joint k is translational, 0 denotes a 3 x 1 zero vector. Finally, T* is the actuated torque of joint k. On the basis of the definition of the spatial quantities we now review the inverse and forward dynamics algorithms" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001007_jsvi.1997.0998-Figure11-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001007_jsvi.1997.0998-Figure11-1.png", + "caption": "Figure 11. The directions of rotation of the pulleys: (a) same direction; (b) opposite direction.", + "texts": [ + " Thus, before analyzing the dynamic characteristics of a belt system, it is necessary to determine whether opposing belt movement has occurred. Belt systems are classified into three types, as shown in Figure 10: (1) the series type\u2014the loops are connected in series; (2) the branch type\u2014a series of loops branches at a loop; (3) the circulation type\u2014the loops are connected in a ring. The direction of rotation of a pulley is decided by the direction of velocity of a belt. For the combination of two pulleys and a belt shown in Figure 11(a), the two pulleys rotate in the same direction, but, for the combination shown in Figure 11(b), the two pulleys rotate in opposite directions. The velocity of a belt at a pulley i is vi = riu i . (12) For the combination shown in Figure 11(a), the relationship of the velocity of the belt at pulley a and pulley b is expressed as vb = va . (13) For the combination shown in Figure 11(b), the relationship of the velocity of the belt at pulley a and pulley b is expressed as vb =\u2212va . (14) Therefore, if a belt system consists of n pulleys, the relationship between the velocity of the belt at pulley i and the velocity of the belt at pulley j, when pulley i is located next to pulley j (i, j=1, . . . , n), is expressed by equations (13) and (14). Unless all the equations are satisfied, opposing belt movement will occur. For the belt system in Figure 12, for exmaple, the equations of the velocities of the belts are obtained as v1 = v2, v2 = v7, v7 = v1 for loop L1 v3 = v2 for loop L2 v3 =\u2212v4, v4 = v5, v5 = v6, v6 =\u2212v3 for loop L3 h G G G G J j " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000227_jp970742n-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000227_jp970742n-Figure1-1.png", + "caption": "Figure 1. Potential profile with adsorbed monolayer on electrode.", + "texts": [ + " We would therefore not expect any change in the diffusion coefficient calculated from limiting current data due to the effects that we discuss below. We do note, however, that a lower value of 6.0 \u00d7 10-7 cm2 s-1 has been reported at a 4,4\u2032-bipyridal disulfidemodified rotating gold electrode.9 One aspect that has not been considered is how the potential drop across the modified layer effects the CV response. By virtue of having a modified layer present, the distance of closest approach of cytochrome c to the electrode is altered (Figure 1). This in turn effects the potential difference and hence the potential energy driving the reaction. Since potential difference across the cell will not be the same as the potential difference acting on the molecule at the distance of closest approach, then qualitatively one might expect some distortion of the CV, the amount of which would be dependent on the magnitude of the potential drop across the modified layer. In such a case, this may lead to the resultant increase in peak separation of the CV", + " The potential of the cell is expressed in terms of the initial potential and the sweep rate Arrival at an expression for the cyclic voltammogram requires the solution of the equations for diffusion, a planar surface with the additional boundary conditions given in eqs 7-11. where c*O is the bulk concentration of the electroactive species. and the current is defined as if we assume that reaction 2 is electrochemically reversible so that the concentrations of O and R at x ) d can always be described by (5), then combining (5) and (6) and rearranging yields where \u03c6(d,t) can be derived from Figure 1 Epzc is the potential of zero charge of the electrode, \u0393m is the surface excess of the monolayer with charge zm, f is the fraction of the adsorbed monolayer that is charged, and Cm and Cdif are the capacitances of the monolayer and the diffuse layers, Ox)\u221e z h Ox)d z (1) Ox)d z + ne- h Rx)d z-n (2) Rx)d z-n h Rx)\u221e z-n (3) \u00b5\u00b0Od + RT ln aOd + zF\u03c6d + n\u00b5\u00b0e M - nF\u03c6M ) \u00b5\u00b0Rd + RT ln aRd + (z - n)F\u03c6d (4) E ) E\u00b0\u2032 + \u03c6d + (RT/nF) ln(cOx)d/cRx)d) (5) E(t) ) Ei - \u03bdt (6) \u2202cO(x,t) \u2202t ) DO \u2202 2cO(x,t) \u2202x2 and \u2202cR(x,t) \u2202t ) DR \u2202 2cR(x,t) \u2202x2 (7) cO(x,0) ) c*O exp[-zOF\u03c6(x,0)/RT] and cR(x,0) ) 0 (8) lim xf\u221e cO(x,t) ) c*O and lim xf\u221e cR(x,t) ) 0 (9) i(t) ) nFADO[\u2202cO(x,t)\u2202x ] x)d (10) cO(d,t)/cR(d,t) ) exp[(nF/RT)(Ei - \u03bdt - \u03c6(d,t) - E\u00b0\u2032)] (11) \u03c6(d,t) ) (E - Epzc)Cm + zmF\u0393mf Cm + Cdif (12) respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001696_s0263-8223(01)00168-4-Figure4-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001696_s0263-8223(01)00168-4-Figure4-1.png", + "caption": "Fig. 4. (a) Spring subject to in-plane bending-torsional line load F . (b) Reduced loading system of a semi-spring subject to in-plane bending-torsional line load F .", + "texts": [ + "# : \u00f08\u00de By simplifying the above expression, the strain energy expression becomes U \u00bc F 2R3 L D11D16 D2 66 jDj c2 2 3 sin3 a \u00fe p 2 a \u00fe sin 2a 2 c 2 cos2 a \u00fe sin2 a cos a \u00fe 1 4 p 2 a sin 2a 2 2a cos2 a \u00fe p 4 cos2 a \u00fe 4 sin a cos2 a \u00fe A11A66 A2 16 Aj j F 2 sin a 4 ; \u00f09\u00de where c \u00bc cos a\u00f02 cos a \u00fe sin2 a\u00de 4 3 sin3 a \u00fe p 2a \u00fe sin 2a : Thus, the deflection is dxy \u00bc oU oF \u00bc 2R L D11D66 D2 16 jDj R2 F c2 2 3 sin3 a \u00fe p 2 a \u00fe sin 2a 2 F c 2 cos2 a \u00fe sin2 a cos a \u00fe F 4 p 2 a sin 2a 2 2a cos2 a \u00fe p 4 cos2 a \u00fe 4 sin a cos2 a \u00fe A11A66 A2 16 jAj F sin a 4 : \u00f010\u00de The in-plane bending-shear spring stiffness is Kxy \u00bc F dxy \u00bc L 2R D11D66 D2 16 jDj R2 c2 2 3 sin3 a \u00fe p 2 a \u00fe sin 2a 2 c 2 cos2 a \u00fe sin2 a cos a \u00fe 1 4 p 2 a sin 2a 2 2a cos2 a \u00fe p 4 cos2 a \u00fe 4 sin a cos2 a \u00fe A11A66 A2 16 jAj sin a 4 1 : \u00f011\u00de 2.1.3. Bending-torsional stiffness (Kxz) The composite spring is subject to unidirectional line load \u2018\u2018F \u2019\u2019 in the z-direction perpendicular to the x\u2013y plane at point B as shown in Fig. 4(a). Similarly, a semi-spring model as shown in Fig. 4(b) is considered. The transverse load of F =2 and the indeterminate bending moment MB are all acting at point B. The out-of-plane bending moment M and the twisting moment T at an arbitrary point P of the semi-spring are taken into account in order to formulate the strain energy expression of the complete spring. U \u00bc 1 Do Z n M2 dn \u00fe 1 b Z n T 2 dn BE \u00fe 1 Do Z n M2 dn \u00fe 1 b Z n T 2 dn EF \u00fe 1 Do Z n M2 dn \u00fe 1 b Z n T 2 dn FC ; \u00f012\u00de where Do is the out-of-plane bending stiffness of the spring and b is the torsional stiffness of the spring" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002191_pccon.1997.645598-Figure2-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002191_pccon.1997.645598-Figure2-1.png", + "caption": "Fig. 2. Relation between rotor position and coordinates.", + "texts": [ + " The estimated position angle by the current injection method is not affected directly by which the variation of the motor parameters. In this system, the position angle estimated by the current injection method and its differentiated value in the time domain as the angular velocity are compared with the each estimated values by MRAS. And a corrector revises the estimated values complimentary. This is shown as an estimated values corrector in figure 1. B. Position angle estimation at standstill and lower speed region In this section, a sensorless control method of the salientpole brushless DC motor is presented. Figure 2 shows the nate systems (u-v-w, CY-/? and d-q axes) are defined. I t is difficult to apply the estimating method a t the running condition as to at a standstill condition, because a t the standstill the estimated angular speed is zero and the position angle can not be calculated correctly. The some estimating methods of the position angle a t a standstill are proposed[6][7][9]. On these paper, the estimating methods analytical model of the motor. In this figure, three coordi- of the position angle are based on the winding inductance which changes as a function of the position angle" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001008_0734-743x(92)90486-d-Figure7-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001008_0734-743x(92)90486-d-Figure7-1.png", + "caption": "FIG 7 A gun pendulum", + "texts": [ + ") It was to Robins' great credit that he perceived ball velocities were able to be calculated and measurements of air resistance deduced. It is very laudable that the French, in a number of books authored by eminent 'students' of mechanics, explicitly recognised Robins' new perceptions and achievements and in the last reference are details of the latter's writings. Later stages of development or enlargement of the ballistic pendulum, see Fig. 6, or sometimes in a more useful form as a gun pendulum, are seen in Fig. 7. Since Robins' contribution is unique in that it is the first time in history that a method 302 W. JoHyso~ o E I e o E e'~ ,,.3 r , , a :t ,a , a Benjamin Robins' past and future perceptions 303 of finding the velocity of a musket ball had been achieved, it was then open to perform in the same manner with regard to cannon balls of greater size and weight; it was merely a matter of building heavier pendula and handling larger masses. This was done with slightly more sophisticated apparatus until about 1850 when the latter was superseded by the development of the electro-chronograph typified in Bashworth's work" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002810_1.1711820-Figure6-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002810_1.1711820-Figure6-1.png", + "caption": "Fig. 6 An instance of Sub-type 3", + "texts": [ + "org/about-asme/terms-of-use Downloaded F changeover positions, obtainable from a simple Pythagorean construction employing the center of the circles and the common midpoint of line-segments AB and C8D8, is found to be r822r25ab . In the special symmetrical configuration alluded to above which provides a device for separating the sub-types, and in which the c-links are directed radially, we have r85 1 2 S ab c 1c D , r5 1 2 S ab c 2c D . For the solution to be of Sub-type 1, then, we conclude that r82>r21ab . The next transitional configuration allows passage to Sub-type 3, rendered by Fig. 6. The imposition of symmetry for this case requires that the a-links be radial, and it is recorded in Ref. @7# that the displacement-closure equations will then be ua5ub1uc2p and ~Kca1Kab!~ tb 2tc 21KcaKab!1~KcaKab11 !~Kabtb 21Kcatc 2!50. These equations also apply to Sub-type 4, of which an instance is limned in Fig. 1, when symmetry is demanded. The two cases are differentiated by the a*-loop (AC8DB8), for which we can write 466 \u00d5 Vol. 126, MAY 2004 rom: http://mechanicaldesign.asmedigitalcollection", + "org/about-asme/terms-of-use Downloaded F For the particular doubly symmetrical pose all a-links and c-links are radially disposed, so that r85 a1c 2 , r5 a2c 2 . Consequently, Sub-type 2 is limited by the conditions r21ab>r82>r21ac and Sub-type 3 by r21ac>r82>r21bc . In our representation of the transition from Sub-type 2 to Sub-type 3 ~Figs. 4, 8, 6!, we have introduced an alternative solution in order to accommodate the original diagrams of Bennett and Wunderlich ~here Figs. 1 and 3, respectively!. Figure 6 would have been more clearly visualized as a progression from Fig. 8 if, in the former, we had mounted link A8C on the side of links A8D and A8B opposite from that on which it is shown. Such solution \u2018\u2018branches,\u2019\u2019 and the \u2018\u2018mirror-images\u2019\u2019 referred to earlier, could well merit further investigation in the interest of a comprehensive study of the network\u2019s closures and transitional states. To include all variations here would make the paper too unwieldy, especially since their significance to a network with two degrees of mobility has yet to be fathomed" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003614_02640410410001730232-Figure3-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003614_02640410410001730232-Figure3-1.png", + "caption": "Fig. 3. Diagram of coordinates used to calculate deviation angle.", + "texts": [ + " It was not a requirement for participants to be active golfers. It was an assumption of the study that the average golfer would not possess any difference in the ability to stand perpendicular to a slope than an individual from the general population. The digital images were analysed using the Microsoft Photo Editor software package. The digital camera was levelled with the floor of the laboratory before image capture. This allowed the angles in the digital image to coincide with the actual angles constructed in the laboratory (Fig. 3). The longitudinal axis of the body was determined by constructing a line from a reflective marker that had been fastened between the participant\u2019s eyes, to a point on the platform at the centre of the participant\u2019s stance. The deviation angle of the longitudinal axis relative to a line perpendicular to the slope of the platform was then calculated for each trial (Appendix 2). The measurement error associated with this process was determined using methods suggested by Bland and Altman (1996) (Appendix 2)", + " The scaling factor was applied to predict the length of the second stick. The error for each image was measured as the absolute value of difference between the actual and predicted length of the second stick. The error estimate was determined by calculating the average error for all 16 images. Scaling Error \u00bc \u00bd X16 i\u00bc1 ABS \u00f0Actuali Predictedi\u00de =16 Appendix 2 The deviation angle of a participant was measured as the angular displacement of the longitudinal axis from a line perpendicular to the slope of the platform (Fig. 3). The angle of the participant\u2019s longitudinal axis relative to the vertical was calculated and the deviation angle was determined using this value. If the participant was standing perpendicular to the platform, then his or her angular displacement from the vertical would equal the angular displacement of the platform from the horizontal. Applying the Cartesian coordinates shown in Fig. 3, the following equation was used to determine the participant\u2019s angular displacement from the vertical: Subject angle from vertical \u00bc Tan 1\u00bdX1 X2\u00de=\u00f0Y1 Y2\u00de The deviation angle was determined by calculating the difference between the angle of the platform from the horizontal and the angle of the participant from the vertical: Deviation angle \u00bc platform angle subject angle The error associated with the deviation angle measurement process was assessed by performing repeated measures on each trial. The digital images for all 93 trials were evaluated on three separate occasions to determine the deviation angle of the participant in the image" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002418_880568-Figure21-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002418_880568-Figure21-1.png", + "caption": "Fig. 21 Coulomb friction contribution to crank pin temperature Fig. 23 Improvement of oil route through main journal", + "texts": [ + " This means that sufficiently accurate predictions can be achieved by allowing a margin of \u00b1lO\u00b0C relative to the evaluation baseline value. _ all flow direction ~=I\"~'''''''I 1 Journal - 2 pin lubrication (VB engine etc.) _~~_ =1 8=0.010N' 1 --,A~! G: lubrlcaUng oll flow (kg/min) I I I. '\\r--V N: engine speed x 103 (rpm) 1 Journal - 1 pin lubrication (L6 engine etc,} 10 CONTRIBUTION RATIO OF COULOMB FRICTION TO PIN TEMPERATURE-In this section the experimental equations are employed to verify existence of a Coulombfriciton term. As shown in Fig. 21, the contribution ratio of Coulomb friction to pin temperature increases with increasing engine speed. In the vicinity of 7000 rpm, the contri bution ratio is approximately 70% for all engine types. CONSIDERATION OF BEARING LUBRICATION OIL FLOW-From Eq. (10) it is seen that theoretically there is a theoretical temperature which does not drop regardless of how much the flow is increased. This condition is equal that the crank pin is bathed in oil at that temperature. Then there are only two possible ways to lower the temperature) as can be seen from the equation. One is to make the reciprocal inertial fore term mr smaller. the other is to make the bearing width L larger. The actual pin temperatures measured for each engine are plotted in Fig. 21 in relation to the saturation temperature. These results indi cate that the pin temperature is generally about 10\u00b0c higher than the saturation temperature. Looked at from the opposite perspective, in en gines showing abnormally high temperature measu rements. there is still considerable room for improvement through increasing the lubrication oil flow. TECHNIQUES FOR IMPROVING BEARING RELIABILITY Bearing reliability can be enhanced by mak ing improvements to each of the various factors listed in Fig. 2. This section introduces sever al examples of improvements made using the pin temperature as an index. The effectiveness of these techniques has been verified experimental ly. IMPROVEMENT ACHIEVED WITH LUBRICATING OIL HOLES-In Fig. 21, an extremely high pin tempera- REDUCTION OF OIL TEMPERATURE-Reducing the lubricating oil temperature involves the addition of an oil cooler. As can be seen in Eq. (7). a reduction in the lubricating oil temperature results in a corresponding reduction in the pin temperature. Because of this one-to-one relationship, reducing the oil temperature is a very effective way to lower the pin temperature. In actuality, though, reducing the oil tempera ture also results in reduced flow because of 880568 11 increased viscosity" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002201_iemdc.2001.939386-Figure5-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002201_iemdc.2001.939386-Figure5-1.png", + "caption": "Fig. 5. Experimental Test Section with Details", + "texts": [ + " The radius of the inner end of the gap supports was reduced close to the stator inner diameter. The gap supports were optimized by CFD calculations to provide the most uniform distribution of flow to the two passages windward and leeward of the stator bar while keeping the losses at a minimum. The new configuration is shown in Fig. 4. IV. TEST-RIG DESCRIPTION In order to investigate the effects of the suggested improvements in detail, a rotating test rig was built. The entire facility is shown in Fig. 6, a closeup of the test section is shown in Fig. 5. This test rig consisted of a periodic element of the motor modeled at a scale of 1:l. Two rotor cooling gaps and one stator cooling gap were modeled. The distribution of mass flow in the stator-rotor gap was controlled by two labyrinth seals on the Drive End (DE) and Non-Drive End (NDE) of the stator-rotor gap. The amount of axial mass flow through the rotor was controlled by an adjustable throttle at the DE of the axial rotor bores. By setting this throttle to fully closed the axial mass flow is zero and the flow into the rotor cooling gaps is similar to the flow in the cooling gaps in the middle of the real motor. By opening the throttle, it was possible to adjust the ratio of axial mass flow through the rotor and mass flow through the rotor cooling gaps to a value similar to the one seen by cooling gaps closest to the faces of the real rotor. The test section was instrumented to provide full knowledge of the ventilation characteristics of the rotor, the stator and the entire system. All relevant mass flows, static and total pressures and temperatures were recorded. In Fig. 5 a section of the test apparatus is shown. Fig. 6 shows a picture of the experimental facility. V. TEST-RIG RESULTS A . Ventilation Characteristic Two series of tests were performed. In the first series the ventilation characteristic of rotor, stator and the entire system was determined for high axial mass flows through the rotor and no axial mass flow through the ro- 3 o f 8 tor. Then rotor and stator were modified according t o the suggested improvements described above. The test series was repeated in order to determine the change in the ventilation characteristic" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000533_bf02192245-Figure5-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000533_bf02192245-Figure5-1.png", + "caption": "Fig. 5. Parallel gradient descent versus parallel decision making: (L) parallel gradient descent with r t = 0.5, r2=0.5, ~/= 5; (R) parallel decision making.", + "texts": [ + " Reducing the iteration times may get rid of the oscillations as seen in Fig. 4 (right), but can lead to slow convergence. Finally, let us compare the path of parallel gradient descent with that of parallel decision making. The parallel decision making is expressed by the following iterative algorithm: p~ e argmin F,(p~- 1 , - 1 t- 1 t- 1). , . . . , P i - I , P i , P i + I . . . . . P N Pi In our simulated example, we assume that the sequence generated by this iterative algorithm is as shown by the thick solid path in Fig. 5 (right). For this particular example, it is observed that the evolution of game under an inaccurate search using the parallel gradient descent is very similar to that of parallel decision-making model. Let it be noted that, in this game example, the evolution of the game does not influence where the players end up as a result of having globally strongly convex payoff functions. However, in general, the payoff functions may admit multiple optima and the equilibrium point of the game is strongly determined by the evolution of the game" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000345_1.580546-Figure2-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000345_1.580546-Figure2-1.png", + "caption": "FIG. 2. Transfer movement of wafer using the developed wafer-transfer robot.", + "texts": [ + " When the holder moves toward the central axis, the spring arms change their shape. Two motors, which can independently operate both revolving arms and two pairs of magnet couplings using a permanent magnet that transmits rotation power into the vacuum region, are located outside the vacuum. In the magnet couplings, the magnets are located on both sides of the vacuum partition wall between the atmospheric and the vacuum regions. If the magnet in the atmosphere is revolved by the motor, the magnet in the vacuum region will also revolve. Figure 2 shows the wafer being moved using the wafertransfer robot. The vacuum system composed of three chambers is shown in Fig. 2. It transfers the wafer to chamber B from chamber A. The preparation chamber including the transfer robot is located between chambers A and B. Transfer occurs as follows: ~1! The holder is turned toward chamber B by the revolving motion of the arm @Fig. 2~a!#. ~2! The holder is moved into chamber B by the linear forward movement of the arm @Fig. 2~b!#. ~3! The holder slides under the wafer on the table in chamber B. ~4! The holder with the wafer is moved out of chamber B by the linear retraction of the arm @Fig. 2~c!#. ~5! The holder with the wafer is moved into chamber A by the anticlockwise movement of the arm @Fig. 2~d!#. ~6! The holder with the wafer is moved into chamber A by the linear forward movement of the arm @Fig. 2~e!#. ~7! The wafer on the holder is moved onto the table in chamber A. ~8! The holder is moved out of chamber A by the linear retraction of the arm @Fig. 2~f!#. By the series of repetitive movements described above, a wafer can be moved to any chamber that is attached around the preparation chamber. The spring arm consists of two sheet springs on the upper and lower side, as shown in Fig. 1. It has the following features: ~1! It has strong rigidity in the upper and lower directions. ~2! The operational revolution torque is lower than that of transfer used by the magnet coupling. ~3! The maximum stress generated by the change in shape of the sheet spring is lower than the maximum permissible stress" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002216_icps.2000.854351-Figure3-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002216_icps.2000.854351-Figure3-1.png", + "caption": "Fig. 3. Illustration of a typical CNC machine and controller", + "texts": [ + " CNC MACHINE COMPONENTS AND SUSCEPTIBILITIES A CNC machine is typically composed of a controller with appropriate computer program describing the desired part, servo-amplifiers and positioning motors to control relative movement of the part and shaping tools, and spindle motors that actually work to shape the parts. Positioning of the part or the tooling is typically accomplished by tuming a screw mechanism and moving a nut in one or more axes. Some machines have five axes, 3-dimensions plus a horizontal and a vertical axis of rotation. See Fig. 3. illustration of a typical CNC machine and controller with 2- dimensions and a horizontal axis of rotation[2]. The CNC machine controller contains a computer to provide overall control of the machine. It usually monitors the machined part\u2019s position by feedback from the resolvers or encoders to update the program. Movement is accomplished by positioning motors with velocity-based feedback from tachometers to the computer. The speed of the spindle motor is also variable using an adjustable speed drive with either computer or manual control[2]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003565_p04-043-Figure3-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003565_p04-043-Figure3-1.png", + "caption": "Fig. 3. (a) An unstable system in a smooth surface model; (b) a stable system is possible in a rough-surface model by the different traction P1 and P2 caused by different asperities.", + "texts": [ + " Another insufficiency is that the finite-element method is incapable of dealing with many asperities with an economical computing resource. Komvopoulos [31] studied only three asperities and Akarca [33] studied only one. \u00a9 2004 NRC Canada In the regular finite-element analysis of no-contact problems, quite a lot of attention has been paid to the geometry of the total body and the load distribution, but these aspects have been ignored by most workers on contact problems. The load distribution can cause subtly different results at different length scales. An intuitive understanding may be obtained from Fig. 3. We can see on the left of the diagram that if the model is treated as a smooth-surface contact problem, the system cannot achieve a balance by this kind of loading distribution. However, if the model is treated as a rough-surface contact problem, a stable equilibrium is possible, which results in a disproportionally higher stress on the right-side asperity. The difficulties are illustrated in Fig. 4. To model the contact with a rough surface, the length of the contact side of the element has to be reasonably small compared with the curve length corresponding to an asperity contour" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003261_2001-gt-0255-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003261_2001-gt-0255-Figure1-1.png", + "caption": "Figure 1: 5DOF model co-ordinate systems", + "texts": [ + " The five DOF modeling follows that of de Mul et al [12] in that it is assumed that ball gyroscopic effects considered by Jones [11] may be ignored. All four models consider individual ball rolling elements and use Hertzian contact theory [5] to model the non-linear stiffness of the bearing. Whereas complicated functions are used for stiffness, constant damping is used to represent the effects of lubrication and friction. Cage influences are ignored though the spacing of the rolling elements is assumed constant. Two degrees of freedom model without inertia (2DOF) This was the model investigated previously [10]. Of the five DOF shown in Fig. 1, only the x and y DOF are used. Contact or loss of contact is considered for each rolling element. Ignoring rolling element centrifugal load effects, the inner and outer race contact deformations can be combined so that the overall contact deformation for the j\u2019th rolling element, \u03b4j, is given by: j j j i ex cos ysin c c\u03b4 = \u03c6 + \u03c6 \u2212 \u2212 (2) Copyright \u00a9 2001 by ASME s of Use: http://www.asme.org/about-asme/terms-of-use Down Summing the contact forces for each rolling element in the x and y directions gives: \u2211 = \u03c6 \u03c6 \u03b4\u03b3= bn 1j j jn jjp y x sin cos K f f (3) where j j j 1 for 0 0 for 0 \u03b4 > \u03b3 = \u03b4 \u2264 and ( ) j c 0 b 2 j 1 t n \u03c0 \u2212 \u03c6 = + \u03c9 + \u03c6 (4) Two degrees of freedom model with inertia (2DOF+i) This model can be considered a simplification of the 5DOF model with centrifugal load effects included [12]", + " With known \u03b4i, all parameters, including the inner and outer race contact forces can be obtained and transformed into the global DOF. Note that when inertia is ignored, contact is lost when \u03b4 is negative but with centrifugal load, the inner race can lose contact even though there is still contact with the outer race. The critical value of \u03b4 when contact is lost, \u03b4c, occurs when \u03b4i = 0. Hence contact with the inner race is lost when \u03b4 \u2264 \u03b4c, where: 1 2 n re c p c pe m D 2K \u03c9 \u03b4 = (8) Five degrees of freedom model (5DOF) As shown in Fig. 1, this model [11,12] includes not only the radial displacements of the inner race but also the rotations about the x- and y-axes and the axial displacement. Two coordinate systems are used. The inner race co-ordinate system (x, y, z, \u03b8x, \u03b8y) corresponds to the DOF of the rotor at this point and has its origin at the center of the bearing. The local rolling element co-ordinate system (r, Z, \u0398) defines the position of the inner race center of curvature for each rolling element and has its origin at the nominal position of the inner race center of curvature" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001847_s0141-0229(01)00369-6-Figure10-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001847_s0141-0229(01)00369-6-Figure10-1.png", + "caption": "Fig. 10. Simultaneous hydrolysis of L-PheAMC (0.0021 mM) and D-PheAFC (0.0021 mM) with 100 mg PLE. a: L-PheAMC/D-PheAFC in KPP; b: 1 h after enzyme addition; c: Difference spectrum; d: Course of the reaction (detection of AMC, ex. 345 nm, em. 440 nm).", + "texts": [ + " In the difference spectrum 9e it is clearly visible, that the AFC peak was increased about 50 RFU at lEX 5 365 nm/lEM 5 490 nm and the D-PheAFC was decreased at lEX 5 330 nm/lEM 5 430 nm (former lEX 5 340 nm). Esterase from porcine liver is an unspecific catalyzing enzyme, hence it is necessary to investigate the hydrolysis with equal concentrations of the coumarin substrates. 6 ml of L-PheAMC (0.021 mM) and D-PheAFC (0.021 mM) were placed within the reactor and 100 mg PLE (shaken for 20 h) were added. The maximum at lEX 5 330 nm and lEM 5 390 nm was shifted toward the L-substrate, the high RFI of L-PheAFC totally covers the maximum of D-PheAFC (Fig. 10a). 1 h after the enzyme addition, an increasing peak of AMC at lEX 5 340 nm and lEM 5 440 nm with 33 RFU was detected (Fig. 10b), the increase is visible in Fig. 10c. Even after 18 h, no AFC peak was monitored, D-PheAFC was not hydrolysed. The reaction course in Fig. 10d however shows an influence of the D-coumarin toward the L-substrate. A comparison with Fig. 6c shows a slower increasing AMC peak. Without the D-substrate, the reaction of L-PheAMC with PLE was finished after 25 min, with both coumarins available, the hydrolysis of L-PheAMC was not finished after 35 min of the enzyme addition. In the last experiment, the coumarins were again used in different concentrations. 6 ml of L-PheAMC (0.0021 mM) and D-PheAFC (0.021 mM) were placed into the reactor and the reaction was started with the addition of 100 mg PLE (shaken for 20 h)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000854_s0389-4304(99)00024-7-Figure7-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000854_s0389-4304(99)00024-7-Figure7-1.png", + "caption": "Fig. 7. Test equipment.", + "texts": [ + " The shock absorber's damping force c(x5 1 !x5 2 ), on the other hand, varies nonlinearly in relation to the relative velocity (x5 1 !x5 2 ) of the vehicle body and wheel; thus, the shock absorber's damping force at di!erent piston speeds can be determined by making slight changes in the vibration amplitude and frequency. (See Fig. 4 on the preceding page.) 3.1. Verifying with manual excitation (measuring the damping force) A testing device was built to verify the aforementioned assumptions, As shown in Fig. 7, the tester consists of (1) a wire-type displacement detector, (2) load detector, and (3) computing unit. The load detector was built to independently sense vertical up-and-down motion only (through the oscillation of a suspension link), and was not a!ected by horizontal forces from the area of tire contact. The wire-type displacement detector measured the relative displacement between the vehicle body and wheel, and load sensors detected the tire's vertical load. The output from these sensors was processed arithmetically in the computing unit (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000021_922404-Figure5-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000021_922404-Figure5-1.png", + "caption": "Figure 5. The working points on a Airbus panel", + "texts": [ + " This stage is very important in this method due to the presence of random movements, the trajectories are generally irregular and incoherent; smoothing aims to eliminate all configurations which are not useful to the robot path. Dichotomy techniques or saddle points can be used for smoothing. APPLICATION OF THE METHOD TO RIVETING PATH PROGRAMMING ANALYSIS OF THE PROBLEM - Some Airbus panels are automatically assembled on riveting machines. \"Assembly\" involves riveting structural components (stringers, frames, etc.) to the skin panel (Figure 5). The path planning problem arises when the die must be positioned at the intersection of two structural components. In this case, the machine and die must be moved alternately to generate a path with no collisions. At present, the programmer determines the free path manually. ~ecausethe shape, size and location of the structural elements differ from one panel to another, there is a multitude of escape trajectories. Consequently, a general path planning method, able to find all these trajectories, is required" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000336_analsci.15.537-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000336_analsci.15.537-Figure1-1.png", + "caption": "Fig. 1 Flow diagram. S, sample; L, sampling loop; IC, injector; C, carrier stream; R, reagent (combined) stream; T, transmission line; a, confluence point; B, coiled reactor; W, waste; Window the glass electrochemical cell; 1, 2 and 3, Ni-Cr alloy, platinum auxiliary and Ag/AgCl (3.0 mol l\u22121 KCl) reference electrodes. Black arrows indicate sites where the peristaltic pump is applied. Numbers are dimensions, in mm.", + "texts": [ + "0 mol l\u22121) concentrations and sample-injected volume (10 \u2013 100 ml) on the linear dynamic range, stability and response time of the electrode were evaluated. The proposed system handles about 60 samples per hour and is very stable and suitable for industrial control. Results within 20 and 80%(w/v) glycerol are precise (rsd <2%) and in agreement with the conventional procedures. Keywords Glycerol, nickel-chromium alloy, cyclic voltammetry, flow-injection, amperometry \u2020 To whom correspondence should be addressed. mm i.d. polyethylene tubing. The electrochemical cell is depicted in Fig.1. All solutions were prepared with distilled water and pro analisi substances. Reagent R (Fig. 1) was a 0.5 mol l\u22121 sodium hydroxide plus 2.0 mol l\u22121 sodium chloride stored in an amber bottle. Working standard solutions within 0.10 and 1.00%(w/v) glycerol were freshly prepared by dilution of the concentrated glycerol stock (99.7%(w/w)) with water. Samples (lixivia from Gessy Lever Co, Valinhos SP, Brazil) were collected in 250-ml bottles. The initial dilution was accomplished by accurately weighing 5 g of the solution and mixing with 500 ml of water. Whenever needed, further manual water dilution was performed immediately before sample injection into the flow system", + " It was pretreated according to Stitz and Burchberger:11 the potential was fixed at 3.0 V during 5 min with a 0.1 mol l\u22121 sodium hydroxide solution at 70\u02daC. For the analysis of the nickel-chromium alloy, the wire (about 50 mg) was dissolved in 10 ml of acqua regia and the volume filled to 100 ml with water. For the determination of the minor elements by ICP-AES12, standard solutions (1000 mg l\u22121 in Ni and Cr) were used for corrections. For Ni and Cr determination, the digests underwent further 1:100 dilution in 0.014 mol l\u22121 HNO3. The sample was injected into the flow system (Fig. 1) by means of a 5-cm (ca. 25 ml) sampling loop and pushed by its carrier stream (0.1 mol l\u22121 NaOH at 1.2 ml min\u22121) towards the detector with a fixed potential (0.5 V vs. Ag/AgCl). At the confluent point (a), located 5 cm from the injection port, the sample zone received the reagent (0.5 mol l\u22121 NaOH plus 2.0 mol l\u22121 KCl at 1.2 ml min\u22121) flowing through the 25-cm coiled reactor. Passage of the processed sample through the electrochemical cell produced a transient anodic current proportional to the glycerol content in the sample. After defining the hydrodynamic conditions, the system shown in Fig. 1 was applied to analyses of lixivia collected from different positions of the plant production. Precision was expressed as the relative standard deviation estimated after eleven successive analyses of typical samples (about 40%(w/v) glycerol), and the accuracy was assessed by running samples already analyzed by Gessy Lever Co. The electrochemical behavior of the alloy electrode was similar to that exhibited by the pure nickel electrode. For the interval of +0.3 to +0.7 V vs. Ag/AgCl, voltammograms obtained with a Ni-Cr alloy electrode in the absence of the analyte consisted of a pair of anodic and cathodic waves at +0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003342_robot.2003.1241866-Figure2-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003342_robot.2003.1241866-Figure2-1.png", + "caption": "Fig. 2 The relationship between the base frame and the calibration frame", + "texts": [ + " This is because: 1) the topological architecture of PKM systems varies from one to another; 2) the impossibility to achieve analytically the inverse of a matrix with orders higher than three. However, the Jacobian J appearing in ( 5 ) generally has the form J = A-IB where A is a 6 x 6 matrix. Based upon the theory of linear algebra and coordinate transformation, an alternative approach is presented to justify the rank deficiency of H , . In order not to lose of generality, consider a PKM with 6-DOF mounted on a rigid body to which the base frame 0, - xlylzl is attached, and put the whole device into the calibration frame 0 - xyz in which the machine is calibrated (see Fig. 2). Here, 0 - xyz is placed in a way that its origin 0 is the intersection of the axial axis (the z axis) of the symmetry of the prescribed workspace and the plane ( x - y plane) in which the \u201cflatness\u201d is measured. Note that the small pemrbation in terms of both position and orientation of 0, - xlylzl with respect to the 0 - xyz may occur though they are always expected to be coincident with each other. Thus, the pose error of the end-effector measured in 0 - xyz can he visualized as the superposition of two components: 1) the component due to the geometric parameter errors defined in 0, -xly,zI ; and 2) the component due to the rigid-body motion of 0, -xly,z, relative to 0 - 9 " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001507_iros.1993.583189-Figure5-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001507_iros.1993.583189-Figure5-1.png", + "caption": "Figure 5: first displacements of the grasping point", + "texts": [ + " If we consider the gripper position at time t, the desired motion which is effectively taken into consideration after selection by S is defined along the direction Y of the gripper frame. But the handle has to rotate about its axis. Therefore, the desired position will never be reached. We will show that, because of this, the increment which is used in the control scheme (after S ) is in general different from vAt: consequently, the actual velocity reference value is not v. Let us consider the first displacements of the grasping point R (figure 5). This point is located at the tip of the handle for the sake of clarity in the figure. Also for clarity reasons, the increments are exagerated in figure 5, although this is not realistic. Di is the desired position at time ti (X,(t,)) Ri is the real position at the same time @(ti)) Yi is the Y axis of the gripper frame at the same time Si is the orthogonal projection of Di on Yi-l ti+l = ti + At At the initial time to, Do = R,. The arc represents the trajectory of the grasping point. Let us consider the error vectors E, and <. Generally speaking, the error at time $+1 is: E,(ti+l) = RiDi+l After selection by S, only the component of &,(ti+l) along Y is taken into account: <(ti+l) = RiSi+l Therefore the desired position which effectively acts in the control is Si+l. We always have: DiDi+l= -vAt Yi Si+l is the orthogonal projection of Di+l on Yi. If Ri was the orthogonal projection of Di on the same axis Yi, we would then obtain: = DiDi+l = -vAt Yi But this assumption is wrong, except by chance or for i=l. As a matter of fact, if we suppose that R, is the optimal position, that is to say the position which minimizes the error RiSi, then Ri is the orthogonal projection of Si on Yi (this is the case in figure 5). Since the grasping point trajectory is on the arc, the direction of Y changes as a function of time. In particular, the axes Yi-l et Yi define different directions. Therefore Si, which is the orthogonal projection of Di on Yi-l, is not aligned with Ri and Di. Consequently, R, is not the orthogonal projection of Di on Yi and therlefore: <(ti+,) f -vAt Yi. This is what can be seen in figure 5. This result is of course valid when Ri is not the optimal position. Generally speaking, the value which is applied to the robot is not vAt. 2.4. Correct calculation of b l ~ d ~ Since the mistake comes from the fact that Ri is not the orthogonal projection of Di on Yi, it is sufficient to replace Di by Ri for calculating the new desired po,sition Di+l : Then: E*(t. ) = = -vAt \\Ii. x 1+1 More formally, the relationship RiDi+l = -vAt Yi yields: bXd(t) = bX(t - At) + bLixd( t). The new desired position is calculated froim the last position actually reached, not from the previous desired position" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003565_p04-043-Figure7-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003565_p04-043-Figure7-1.png", + "caption": "Fig. 7. The unit normal vector n of the rigid surface and the normal gap gN(X).", + "texts": [ + " This free-boundary contact adds a nonlinear nature to the contact problem, even for a linear elastic material medium. Rough contact will generate a discontinuous real contact surface filled with a fluid such as air and lubricant in between. Here, the computational domain is defined on the region of the undeformed solid bodies. Thus, the (X1, X2) is defined as the material coordinates of a point X in the computational domain. The rigid surface is modeled as a function \u03c8(X1) . The normal unit vector of a rigid surface is written as N = (N1, N2). This is illustrated in Fig. 7 where N1 = \u2212 \u2202\u03c8 \u2202X1 \u2223\u2223\u2223 X1\u221a 1 + ( \u2212 \u2202\u03c8 \u2202X1 \u2223\u2223\u2223 X1 )2 , N2 = 1\u221a 1 + ( \u2212 \u2202\u03c8 \u2202X1 \u2223\u2223\u2223 X1 )2 (5) The undeformed normal gap between the two contact surfaces at positionX1 is defined following refs. 11 and 12 to be gN(X1) = \u03c6 (X1) \u2212 \u03c8(X1)\u221a 1 + ( \u2212 \u2202\u03c8 \u2202X1 \u2223\u2223\u2223 X1 )2 (6) The classic formulation of the elastic-contact problem in this model is to search for the vector displacement fields ui (X1, X2), defined on satisfying the static equilibrium equation formulated from the \u00a9 2004 NRC Canada vector form of the second Piola\u2013Kichoff tensor \u03c3\u0303ij [37] \u2202(\u03c3\u0303ikF jk) \u2202Xj + fi = 0 (7) whereFij = \u2202xi/\u2202Xj withxi = Xi+ui andfi is the body force density" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002607_978-94-017-0657-5_14-Figure2-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002607_978-94-017-0657-5_14-Figure2-1.png", + "caption": "Figure 2. Kinematic constraint", + "texts": [ + " The relation between the time derivatives of y and s is then given by (3) y = H(y) s, whereby the second matrix equation is the kinematic differential equa tion related to the Bryant angles. With three supporting cables the payload platform can perform sway motions with three degrees of freedom. The sway motions are described, e.g., by the sway coordinates 'TJ = [rx ry 'P3]T. (4) With three longitudinally stiff cables there exist three implicit con straints between the robot coordinates p and platform coordinates y (Fig. 2) i = 1,2,3, or g(p,y) = 0, (5) 128 whereby the cable vectors Ci can be ex pressed by Ci = r + d i - bi , i = 1,2,3 . (6) To derive the feedforward controller model, the equations of motion of CA BLEV are formulated as differential algebraic equations in terms of the co ordinates p and y. They consist of the kinematic differential equation (3), the constraint equation (5), the seven dy namic equations for the assembly of the gantry, the trolleys and the winches Mp(p) P = Kpw + G~(p, y)>., (7) and the six dynamic equations of the platform between p and y Ms(Y) S + ks(y, s) = Q~(y) + G;(p, y)>" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001442_bf01573714-Figure9-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001442_bf01573714-Figure9-1.png", + "caption": "Fig. 9. Mean propulsion force _P~ of a vehicle versus angle y for different values of X (armature non-overlapped winding). [m = 51; gzF = 3.51 m; Z = 0 m]", + "texts": [ + " 8 [1]; in this system the stator three-phase winding is made up of non-overlapped coils and the N b - - T i superconducting winding on the vehicle is used both for the motor excitation and for the vehicle levitation. With reference to Figs. 2b, 3b and 4, the following values are assumed for the sizes and pitches of LSM field and armature coils: 2 = g u A = 4 . 2 m p E = 2 . 1 m P A = 1 .4m L F = l . 3 7 6 m H F = 0 . 4 5 1 m t F = 0 . 0 4 9 m sF = 0.078 m L~ = 1.020 m s~ = 0.080 m Ha ---- 0.620 m tA = 0.080 m The values of the propulsion force given in the following diagrams refer to a single vehicle of tile train, on which two four-pole motors - - one for each side of the guideway -- are arranged. Figure 9 shows the distributions of the mean propulsion force Fp versus the angle y when Z = 0 and four values of the parameter X are considered, while Fig. 10 shows the distributions of/~v versus Z when the same values of X are supposed and when 7 = --0.52 rad. Only in order to give an example of application of the developped method also in case of overlapped winding, we have supposed to replace, in the abovementioned MAGLEV system, the LSM non-overlapped armature with an overlapped winding, for which the following values are assumed with reference to Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002913_1.1479336-Figure6-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002913_1.1479336-Figure6-1.png", + "caption": "Fig. 6 Rotor model and 20 elements in finite element modeling", + "texts": [ + " The accurate results of identification are Table 1 Rotor data Table 2 Eccentricity data OCTOBER 2002, Vol. 124 \u00d5 981 shx?url=/data/journals/jetpez/26816/ on 06/17/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use Downloaded F due to the sufficient measurement information and the noise-free condition. To improve the identification accuracy, the shaft is modeled with 20 shaft elements, which is more accurate in finite element modeling than four elements. The improved finite element model of the rotor is shown in Fig. 6. Again, the middle point, node 11, is measured and the results are compared with those of the previous case. Table 3 shows that the averaged coefficient error reduced to 1.00231024%. This improvement is due to the decrease of the condition number of @Tsys# , where the condition number is reduced from 2.24031011 to 3.89131010. Noise inevitably pollutes the signals in field measurements and enlarges the identification errors. Simulations are conducted with added noise of NSR51% into the signals. 400 sets of signals are averaged" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000725_a:1007941331202-Figure4-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000725_a:1007941331202-Figure4-1.png", + "caption": "Figure 4. Algorithmic singularity (dashed) for 3-link planar manipulator with joint range availability measure and additional formulation singularity (solid) added by compact form of Lagrange multiplier method.", + "texts": [ + " However, drift within the null space can affect the orthogonality of the basis vectors and cannot be fixed with the drift correction right-hand side of (10) [9]. The renew operation, however, eliminates this source of error for the EJM. JINT96B3.tex; 30/04/1997; 15:56; v.7; p.11 One might wonder how the complexity of computation grows with the degree of redundancy. To evaluate this complexity, a planar, n-link manipulator is con- JINT96B3.tex; 30/04/1997; 15:56; v.7; p.12 trolled for position. Hence the degree of redundancy will always be 2 less than n. The configuration of this problem is shown in an insert to Figure 4 and the details of this formulation can be found in [11]. Table III shows the number of floating point computations for the renewed EJM and updated LMM schemes. These data points can be fit well by a cubic which is the expected degree of computional complexity one expects to find from solving linear systems. If redundancy is resolved by minimizing a criterion function and all techniques are bounded by the same algorithmic singularities, then for practical robot control, it is important to know where these boundaries are", + " In [9] it is determined that (9) fails only at kinematic singularities, and that the extended Jacobian formulation reviewed here introduces no additional singularities. However, the most compact version of the Lagrange multiplier formulation, (20), depends on \u2207(\u2207H)\u2212 P not being singular. To examine whether there are actually cases in which (20) fails, the above singularity tracing methods were applied to the planar, three-link manipulator problem. The manipulator was controlled for position only and redundancy was resolved using joint range availability. Figure 4 shows the results of the search for singularities. The solid line shows a locus of points where (19) fails. However, the full size version of the Lagrange multiplier formulation, (18), is not singular here. Basically, this singularity occurs when the matrix used to couple the top and bottom halves of (18) degenerates. Because this apparent singularity only occurs in a particular formulation, it could be said to be a formulation singularity. On this same figure, segments of the algorithmic singularity are shown in dashed lines" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001681_s0022-0728(85)80005-x-Figure10-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001681_s0022-0728(85)80005-x-Figure10-1.png", + "caption": "Fig. 10. Current-potential curves for the reduction of O~ in air-saturated solution (pH = 2.8) at a rotating glassy carbon disk electrode coated with QNH 2 (F i = 3\u00d710 - l \u00b0 mol cm-2) ; v = 25 mV s -1. The disk electrode rotation rate (in rad s -1) is indicated on each curve. ( - - - ) 02 response on a bare electrode. ( ) (a) Q N H 2 / G C electrode in N2-purged solution, (b) Q N H z / G C electrode in air-saturated solution.", + "texts": [ + " electrolyses of dihydrophenazine (PH2) have thus been performed in NA 2+ solution, PH 2 being isolated from the medium by simple filtration. Electrocatalytic reduction of dioxygen at the 2,6-diaminoanthraquinone GC electrode Rotating disk GC electrodes coated with QNH2 catalyze the reduction of 0 2 to H202 at the reduction potential of the anthraquinone ( - 0.4 V at pH 2.8). Even with thin deposits the catalytic efficiency remains unchanged for long periods of time which shows the stability of the electrocatalyst towards H202. Figure 10 summarizes experiments carried out at various angular velocities (~0) in air-saturated buffered solutions on electrodes coated with (sub)monolayers of QNH 2 (F i = 3 \u00d7 10 -l\u00b0 mol cm-2). The dashed line is the response of 02 on a bare gc electrode. The limiting current (iL) being defined as the difference between the currents on a modified electrode at -0.65 V in N 2 purged and air-saturated solutions, the Levich plot (17) of i L vs. \u00a20 -1/2 (Fig. 11) indicates that i L is mass-transport limited up to ~0 = 54 rad 71 s-l" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003375_icmech.2004.1364455-Figure3-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003375_icmech.2004.1364455-Figure3-1.png", + "caption": "Fig. 3. Energy optimal motion with limits on the joint ievoiutiona.", + "texts": [ + " The NLP thus has jrl = 10 additional equality constraints (d'). The detection distances are set to A, = 0. The deuicted maniuulator motion minutes. Higher accuracy yields marginal improvement. Also the necessary run.time increases nonlinearly with the accuracy. Typically with a sliort duration (T = 20) it iz energy optimal to fold the manipulator before turning and afterwards unfolding it. The obstacles mainly interfere within the unfolding motion. Taking into account joint limitations the preceding motions are likely to be impossible. i n figure 3 the revolution of the last four joints is restricted to 450 degree. Thus (e') contains [ / q - q ~ . , J + = 0, a = 2,. . . , 5 . Another scenario is shown in figure 4. The energy optimal control with a safety margin A- = 0 very closely follows the obstacle boundaries. Due to model uncertainties and the finite number of sensors it is nossible that is the result of a numerical solution with a required accuracy of in the objective value. The solution is found within 10 a control scheme using the determined U,, will cause collisions" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002478_ias.1998.732271-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002478_ias.1998.732271-Figure1-1.png", + "caption": "Fig. 1: Schematic of a motor structure of the transverse-laminated type.", + "texts": [ + " This one can be obtained by both the axially-laminated and the transverse-laminated types of rotor construction. The latter type (transverse) is preferable, in practice, because it is more suited to industrial manufacturing. In addition, the rotor can be easily skewed, in this case, thus allowing to get very low values of torque ripple [IS]. The transverse-laminated rotor structure requires, in practice, that the various iron segments in the rotor be connected to each other by thin iron ribs, as schematically shown in Fig. 1. These ribs are magnetically saturated by the stator m.m.f., thus allowing the requested anisotropic behavior. On the other hand, these ribs enhance cross saturation, with reference to the d, q magnetic model of the machine. This in addition to the usual phenomenon, which is common to the other types of machines. As a consequence, an accwte 0-7803-4943-1/98/$10.00 0 1998 IEEE 1 27 modelization of the magnetic behavior is required, for both evaluation of the machine\u2019s performance in terms of torque and definition of the most suited control strategy", + " Let us also point out that the above mentioned torques values have been measured at standstill, by averaging several measurements at different angular positions. The standstill torques have been chosen to neglect iron loss, accordingly with the hypothesis of a conservative system. During the test the motor was supplied by the same digitally controlled CRPWM inverter employed for flux measurement. The supply interval is enough long to get a stable torque measurement, triggered by the inverter\u2019s microcontroller. 70 60 50 40 30 20 10 0 0 10 20 30 40 50 60 Fig. 1 1 : Torque vs current for different 7, checked by stall torque measurements. In Fig. 12 the locus is shown in the (id, iq) plane, f a which the maximum torque is obtained at a given current module. It is commonly referred to as \u201cmax \u201c/A\u201d locus or \u201cmax k7\u201d locus. It is evident, in the figure, the needed increase of d-axis current, for a given torque, to counteract cross-saturation. It can be also observed that the \u201cmax kT\u201d locus (with crosscoupling) is close to a \u201cconstant current-argument\u201d strategy, at least in the considered current range" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000408_ac00066a008-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000408_ac00066a008-Figure1-1.png", + "caption": "Figure 1. Schematic diagram of the thin-layer cell used to evaluate the feasibility of the proposed approach.", + "texts": [ + " The manner in which this system was used in this study is described in the next section. E,, + 20s(III) * E,, + 20s(II) Most prior applications of enzyme-based reactor/sensor systems have involved sufficiently large volumes of analyte solution and sufficiently small reactor surfaces that the time required for all the substrate in the sample solution to react would be very long. Accordingly, it would be very time consuming to make equilibrium-based measurements on such systems. To overcome this problem, we used the reactor/sensor system in a thin-layer flow cell as illustrated in Figure 1. As shown in the bottom part of the figure, the thin-layer cell consisted of two solid blocks (Teflon (top) and stainless steel (12) Mieling, G. E.; Pardue, H. L. Anal. Chem. 1978,50,1333-1337. (13) Harris, R. C.; Hultman, E. Clin. Chem. 1983,29, 2079-2086. (14) Wentzell, P. D.; Crouch, S. R. Anal. Chem. 1986,58,2855-2858. (15) Pardue, H. L. Anal. Chim. Acta 1989,216, 69-107. (16) Mieling, G. E.; Pardue, H. L.; Thompson, J. E.; Smith, R. A. Clin. (17) Hamilton, S. D.; Pardue, H. L. Clin. Chem. 1982,243,2359-2365", + " The reactor/sensor system consisted of the glucose oxidase and electron mediator (Os(II/III)) immobilized on the surface of a glassy-carbon electrode. Solutions of glucose oxidase (2 g/L) and oxidation/reduction polymer (4 g/L) were prepared in HEPES buffer (10 mmol/L, pH 8.2), and asolutionofPEG (2.3 g/L) wasprepared inwater. Thesesolutions were mixed to give a final solution containing 18.3% glucose oxidase, 8.4% PEG, and 73.3% polymer. A 4-pL aliquot of this solution was coated onto the surface of the glassy-carbon electrode and permitted to react as described earlier.1Q~M~22 A schematic diagram of the thin-layer cell is shown in Figure 1. The electrode used in this study was a 3-mm-diameter cylinder of glassy carbon imbedded in a Teflon block which formed one side of a thin-layer flow cell (Bioanalytical Systems, West Lafayette, IN). A silver/silver chloride reference electrode (Bioanalytical Systems) was screwed into a stainless-steel block (Bioanalytical Systems) which formed the other side of the flow cell. The cell thickness was controlled by using Teflon spacers of thicknesses between 14 and 30 pm. These spacers had oval cutout areas of about 1 cm2 surrounding the glassy-carbon electrode (3-mm diameter)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002652_roman.1997.647004-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002652_roman.1997.647004-Figure1-1.png", + "caption": "Figure 1. Image of welfare support systems", + "texts": [ + " The behavior is the forward and the backward displacement of the agent. The emotions of the agent are represented by the status \"like\" or \"dislike, \" depending on the situation when the agent likes or dislikes the person. Finally, the awareness is the behavioral intention of the agent. The agent approaches the person or goes away from him based on the model. Experiments show the efjectiveness of this model. 1. Introduction The concept of systems providing physical assistance to handicapped people is shown in Figure 1. A handicapped person is assisted by mobile robots (intelligent agents) which get information by computer graphics. If the person's instructs an agent by gesticulation, the gesture are transformed to sensor information using a CCD camera, From the sensor information, the agent estimates the person's intention as either \"I go there\" or \"I go away from here\" and the person's emotion as \"happy\" or \"angry.\" The agent determines the behavior of the person based on estimated results. The person will damage the agent by giving wrong instructions or poor directions" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002866_095440603322769938-Figure7-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002866_095440603322769938-Figure7-1.png", + "caption": "Fig. 7 Coordinate systems for a simulation of assembly errors", + "texts": [], + "surrounding_texts": [ + "The in uence of errors is calculated through the tooth contact analysis technique. This technique is employed to evaluate the assembly errors for the gear with ringinvolute teeth. To evaluate the meshing conditions, the coordinate systems S 1\u2026X1, Y 1, Z 1\u2020, S 2\u2026X 2, Y 2, Z 2\u2020, S f \u2026X f , Y f , Z f \u2020 and Sh\u2026X h, Y h, Z h\u2020 have been set up as shown in F ig. 7. The coordinate systems S 1 and S 2 are rigidly connected to the pinion and the gear surfaces respectively. The xed coordinate system S f is rigidly connected to the housing of the gear pair. The auxiliary coordinate system Sh is applied to the evaluation of assembly errors. The parameters of gear surfaces, `h and b, retain the original set to investigate assembly errors. However, the original parameters of the gear surface, `h and b, are replaced by gh and g0 respectively. Coordinate systems S f and Sh coincide with each other if there are no assembly errors. Assembly errors can be evaluated by changing the orientation of the coordinate system Sh with respect to S f . The sensitivity of the centre distance and misalignment of the tilted axis between the ringinvolute pinion and the ring-involute gear can be evaluated by changing the setting and orientation of the coordinate system Sh. For instance, by evaluating the change in centre distance Dd, the origin Oh of the coordinate system Sh may be misplaced by Dd with respect to Of , as shown in F ig. 7. The centre distance between O2 and O1 is de ned as d 0 \u02c6 d \u2021 Dd . In this evaluation, the misalignment angle Dr is decomposed into two components, Dh and Dv, which represent the horizontal and the vertical misalignment angles respectively. To obtain the angles Dh and Dv, the coordinate system Sh may be rotated along the X h and Y h axes at angles Dv and Dh with respect to the xed coordinate system S f . The angle f0 1 is the rotary angle of the pinion. The angle f0 2 is the rotary angle of the gear. By applying coordinate transformations, parametric equations of the C02203 # IMechE 2003 Proc. Instn Mech. Engrs Vol. 217 Part C: J. Mechanical Engineering Science at Purdue University on July 22, 2015pic.sagepub.comDownloaded from ring-involute pinion and the ring-involute gear can be represented by R1 f \u02c6 MfhMh1R1, R2 f \u02c6 Mf 2R2 \u202626\u2020 n1 f \u02c6 LfhLh1n1, n2 f \u02c6 Lf 2n2 \u202627\u2020 Here matrices Mfh and Mh1 represent the coordinate transformations in transition from S 1 via Sh to S f . Matrices Lfh and Lh1 are the 363 submatrices of Mfh and Mh1 respectively. Matrix Mf 2 transforms the coordinate system S 2 into that of the system S f . Similarly, matrix Lf 2 is the 363 submatrix of Mf 2. Position vector R1 is the pinion surface, which is represented in the coordinate system S 1. Position vector R2 is the gear surface that is represented in the coordinate system S 2. The unit normal to the pinion surface is represented by n1. The unit normal to the gear surface is represented by n2. Vector R i f \u2026i \u02c6 1, 2\u2020 represents R i \u2026i \u02c6 1, 2\u2020 in the coordinate system S f . Vector ni f \u2026i \u02c6 1, 2\u2020 represents ni in the coordinate system S f . The basic concept of the technique is that the position vectors and their unit normal vectors should be the same at the point of contact. Therefore, it is necessary to represent the equations of pinion and gear surfaces in a xed coordinate system. When applying the obtained geometrical model and the tooth contact analysis at the point of contact, due to the tangency of the two contact surfaces, the following equations must be observed: R1 f \u2026`h, b, f0 1\u2020 \u02c6 R2 f \u2026gh, g0, f0 2\u2020 \u202628\u2020 n1 f \u02c6 n2 f \u202629\u2020 where vectors R1 f and R2 f indicate the pinion and the gear respectively. Vectors R1 f and R2 f are in the coordinate system S f . Vectors n1 f and n2 f express their normal section in the coordinate system S f . When considered simultaneously, equations (28) and (29) yield a system of ve independent equations, since jn1 f j \u02c6 jn2 f j \u02c6 1. These ve equations relate to six unknowns, f0 1, f0 2, `h, b, gh and g0, and one of these unknowns may be considered as a variable. The results of the TCA yield the relationship between rotary angles f0 2 and f0 1. Since f0 1 is chosen as an input variable, the rotary angle f0 2 is a function of f0 1 and can be written as f0 2\u2026f0 1\u2020. This function is non-linear and its deviation from the linear function f0 1N 1=N 2 de nes the kinematic error of the gear pair, given by Df0 2\u2026f0 1\u2020 \u02c6 f0 2\u2026f0 1\u2020 \u00a1 N 1 N 2 f0 1 \u202630\u2020 where N 1 and N 2 represent the number of teeth of the gear and the pinion respectively. The Df0 2\u2026f0 1\u2020 term represents the kinematic errors in a gear pair induced by the assembly errors. The change of centre distance between the two gear rotating shafts is given by the design value, Dd \u02c6 0:1 mm. It is assumed that the rotating shafts have vertical and horizontal misalignment angles, Dh \u02c6 30 and Dv \u02c6 30 respectively. The solution discussed above is calculated through a general purpose computer program. A function for the solution of the non-linear equations is represented in a computer program. By substituting the values of the assembly errors into equations (28) and (29) and using the computer program, the kinematic errors of the gears with ring-involute teeth can be evaluated and listed as shown in Tables 2 to 5. F rom Table 4, the kinematic error of gear with ring-involute teeth is similar to the kinematic error of spur gears with involute teeth. In other words, the sensitivity of the centre distance errors for the ring-involute form is zero. The change in centre distance between the ring-involute gear and the involute pinion axes does not induce kinematic errors in rotation." + ] + }, + { + "image_filename": "designv11_65_0001863_s0022-0728(84)80344-7-Figure4-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001863_s0022-0728(84)80344-7-Figure4-1.png", + "caption": "Fig. 4. Dependence of the normal ized D N P P current response of the E ~ E m e c h a n i s m on the degree of a t t r ac t ion be tween adsorbed ions of the in termedia te . F r u m k i n ' s isotherm. A~ = A i ( ~ ( t 1 + tp))l/2(FqD1/2A*)-l; t 1 = tp = 50 ms; A E = - - 5 0 mV; a = 0.0 (1), --1.0 (2), --2.0 (3), - 4 . 0 (4) and", + "texts": [ + " The scheme of this excitation signal was given previously [24,40]. The current is measured at the end of the first and second pulses and then the difference between these two measurements is recorded. For adsorption of the intermediate following Langmuir 's isotherm ( A * / F s -- 1.3 \u00d7 103 cm -1, flFs = 1.0 cm), the DNPP current response of the E ~ E mechanism (El \u00b0 = E2 \u00b0 , tpl = tp2 = 50 ms and AE = - 5 0 mV) exhibits three peaks: a small pre-peak, a high main peak and a small post-peak (see curve 1, Fig. 4). The influence of the attraction between the adsorbed ions can be seen in Fig. 4. As in pulse polarography, the separation between these three peaks and the significant increase of both the pre-peak and post-peak can be seen. The main peak also increases, but not as much as the other two. From the formal point of view, we can treat NDPP responses as being similar to the first derivative of the PP response and then the changes of current which can be observed in Fig. 4 can be understood if they are compared with the changes of current from Fig. 1. The pre-peak corresponds to the rising portion of the pre-wave, the main peak corresponds to the main wave, and the post-peak corresponds to the post-wave. The half-widths of the pre-peak and post-peak decrease if parameter \" a \" becomes more negative, which is in agreement with literature data [38b]. The increase of the amplitude of the main peak reflects the increase of the slope of the main wave in PP, which is also in agreement with literature data [11,37,38a]", + " The needle-shaped peaks near the summit of the post-peak, which appear for a < 4, are probably the consequence of the instability of the mathematical model, rather than representative of a certain physical process. Note that in PP (Fig. 1) there is no corresponding sudden jump of the rising portion of the post-wave (curve 7, a = - 4). D N P P is probably more sensitive to instabilities than PP. In Fig. 5 one can see the influence of the repulsion between the ions from the adsorbed layer on the DNPP current response of the E~ E mechanism. All the parameters are equal to those in Fig. 4. Very simple changes can be observed: the pre-peak decreases, the main peak remains virtually unchanged and the post-peak vanishes. The half-widths of the pre-peak and the main peak increases, which is the reflection of the decrease of the slopes of the pre-wave and main wave (see Fig. 2). The disappearance of the post-peak is the result of the fact that the post-peak is 41 always less pronounced because it appears at potentials at which the total current response is negative [24,40]. (1) Finally, the following conclusions can be drawn: At very low ( < 10 3 c m - 1 ) and very high ( > 10 4 c m - 1 ) values of the parameter A * / F s , the influence of adsorption of the intermediate which follows Frumkin 's isotherm is equal to the influence of the adsorption following Langmuir 's isotherm (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002201_iemdc.2001.939386-Figure16-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002201_iemdc.2001.939386-Figure16-1.png", + "caption": "Fig. 16. Details of the Modified Prototype", + "texts": [ + " With the modifications to the stator bar and to the gap supports this reingestion is suppressed. The cooling air is leaving the rotor-stator gap uniformly through both slots 1 1 2 1 P 3 - P 2 ' = - 'P' (W2 - U z 2 ) i- - ' P . C a x 2 - 5 - P - C T ~ ~ (11) 2 2 6 of 8 on the windward and on the leeward side of the stator bars. The flow stays attached to the channel walls (gap supports and wedges) all the way into the plenum. The improved flow is shown in Fig. 15. VI. PROTOTYPE DESCRIPTION Pictures of the modifications applied to the prototype engine are shown in Fig. 16. They are virtually identical to the modifications applied to the experimental test section. A motor of the same type but without any of the modifications was tested as a reference. VII. PROTOTYPE R SULTS Both motors were tested under the same operating conditions. Table I lists the physical parameters recorded under steady-state conditions. The temperature increase of the stator bars with respect to the cooling inlet temperature is reduced from 42.2\"C to 37.9\"C which corresponds to a relative reduction of 10%" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003634_cdc.2004.1430205-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003634_cdc.2004.1430205-Figure1-1.png", + "caption": "Fig. 1. The mathematical model for the hovercraft", + "texts": [ + " In this section we present the model for the system we study, and contrast it with the simpler model initially studied in [1]. We adopt a modification of the approach of the approach of [3] who consider invariant systems on Lie groups. While the simpler model of [1] is an invariant system on a Lie group, the more sophisticated model we present here is not. It is, however, an invariant system on a trivial principal fibre bundle, as considered in [6]. A Hamiltonian setting for this sort of problem is considered in [7]. The system we study is shown in Figure 1. We have a planar rigid body moving in a plane orthogonal to the direction of gravity. Sitting atop the rigid body is a fan which may be rotated via the torque \u03c4 , and which provides a thrust F . The frame {e1, e2} is inertial, and we affix to the centre of mass of the body a frame {f1, f2}, choosing the f1-axis so that along it lies the point of application of the thrust force. The frame {g1, g2} is affixed to the centre of mass of the thrust fan, which we assume to coincide with its point of rotation", + " Therefore, from the preceding two lemmas we know that if we define vector fields Ya = g (F a), a \u2208 {1, 2}, on Q, these vector fields will be G-invariant. We also know that the Levi-Civita connection g \u2207 for g will be G-invariant. The system is then an affine connection control system \u03a3hc = (Q, g \u2207, Y = {Y1, Y2}, U \u2282 R 2), meaning, as we shall see in Section III, that the control equations are g \u2207\u03b3\u2032(t)\u03b3 \u2032(t) = 2\u2211 a=1 ua(t)Ya(\u03b3(t)). Summarising this is the following. Proposition 3: The affine connection control system \u03a3hc = (Q, g \u2207, Y , U) corresponding to the system in Figure 1 is a G-invariant system on the principal fibre bundle \u03c0 : Q = SE(2) \u00d7 SO(2) \u2192 SO(2). This structure leads to an interesting consequence concerning the input vector field Y2. Note that the one-form F 2 annihilates the vertical bundle V Q. Therefore, the vector field Y2 is g-orthogonal to V Q. This, along with Ginvariance of Y2, implies that Y2 is the horizontal lift of a vector field Y\u03032 on SO(2). Indeed, one verifies that Y\u03032 = \u2202 \u2202\u03c6 . This observation turns out to yield the following. Lemma 4: The vector field \u2207Y2Y2 is horizontal" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002689_a:1014805931636-Figure3-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002689_a:1014805931636-Figure3-1.png", + "caption": "Fig. 3", + "texts": [ + "1) are determined in a theorem whereby the solutions of a system of differential equations depend on the parameter \u03b4 continuously. Since the origin of the space of variables \u03d51, \u03d51\u2032, \u03d52, \u03d52\u2032 is a stable singular point for \u03b4 = 0, \u03d51 \u2217 = 0, and \u03d52 \u2217 = 0 [4], all the points of the branch (AC) are also stable. Hence, the points of the branches (CB) and (DA) are unstable and the points of the branches (GB) and (DE) are stable. This dependence of singular points of dynamic system (3.1) on the parameter \u03b4 is confirmed through the direct integration of the system. Figure 3 shows the projections of the phase paths of the dynamic system (3.1) for \u03b4 = 0.1 (Fig. 3a) and \u03b4 = 0.4 (Fig. 3b) onto the plane \u03d51 \u03d51\u2032. They correspond to lines 1 and 2, respectively, in Fig. 2. For small values of | \u03b4 | (more precisely, for | \u03b4 | < \u03b4A) and k = 0.25, the dynamic system (3.1) has three stable focal points F, F1, and F2 and two saddle points S and S1. The curve of stationary states in Fig. 2 shows that transitions from one stable stationary state to another are possible when the parameter \u03b4 changes, i.e., smooth changes in the parameter \u03b4 may lead to spasmodic changes (catastrophes) in the stationary states. The two-dimensional phase portraits constructed for \u03b4 = 0.1 and \u03b4 = 0.4 and shown in Fig. 3 confirm this conclusion. According to Fig. 2, as the parameter \u03b4 further increases, the singular points F1 and S1 merge and disappear when \u03b4 > \u03b4B, with only one stationary state F2 corresponding to line 3 remaining. This conclusion is supported by the two-dimensional phase portrait constructed for the above-mentioned values of the pendulum parameters and \u03b4 = 0.715. The configuration of the curve of stationary states varies with the parameter k. For example, when k = 0.5, the curve in Fig. 2 turns into that for which, if \u03b4 increases, dynamic system (3" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001307_960083-Figure2-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001307_960083-Figure2-1.png", + "caption": "Figure 2: A sectional sketch of the prechamber.", + "texts": [ + " by replacing the fuel injectors with prechambers equipped with ignition plugs (Fig. 1). The catalytic insert, which is a platinum wire (coil l[mm] in diameter and 400[mm] in lenght), was suspended on a special hanger screwed into the prechamber. The prechambers were cooled with water drawn from the cooling circuit of the engine. Table 1 presents engine data for both variations (Diesel and Otto engines). In the Otto version the engine was tested with prechamber (the capacity of which was 24% of the capacity of the main chamber - see Fig. 2). For this configuration of the engine the geometric compression ratio was \u03b5=16. \u2032 In the first stage of the experiments air-fuel mixture was formed in a carburettor. In the second stage of the investigation the engine was fed by an injection system. The fuel was injected into the inlet manifold just before the inlet valves. In the course of the working cycle, the contact between the fresh air-fuel mixture and the catalytic insert is possible only during the compression stroke. During this stroke, fresh mixture begins to enter into the prechamber before the inlet valve is closed, thus entering into contact with the heated catalyst" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001255_20.767379-Figure2-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001255_20.767379-Figure2-1.png", + "caption": "Fig. 2. Croea-section of an electromagnetic valve used in ABS systems", + "texts": [ + " This in turn increases the memory consumption and the CPU time required for the the system of equations. Another problem has when using the BEM-FEM coupling. If the and in consequence their boundary elements of separate subdomains approach each other closely during the motion so called nearly singular integrals occur in the BEM. These integrals have to be evaluated using special techniques 1151, [16]. VI. NUMERICAL EXAMPLE To compare both numerical approaches for the analysis of electromechanical devices an axisymmetric electromagnetic valve used in ABS systems has been investigated, Fig. 1, Fig. 2. The yoke (St4U), the armature and the armature counterpart (both Vacoflux50) consist of different conducting and magnetic nonlinear media (QSUU = 1.818 * 106S/m, qacoflux = 7.693.106S/m), Fig. 3. The stroke of the armature is 0.3 mm. The driving coil of the valve has of 546 turns. The prescribed current in the driving coil is depicted in Fig. 4. The constant hydraulic load acting on the valve needle is 15N. The bias force of the spring is 5.93N. Effects due to friction have not been taken into account" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001847_s0141-0229(01)00369-6-Figure11-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001847_s0141-0229(01)00369-6-Figure11-1.png", + "caption": "Fig. 11. Simultaneous hydrolysis of L-PheAMC (0.0021 mM) and D-PheAFC (0.021 mM) with 100 mg PLE. a: L-PheAMC/D-PheAFC in KPP; b: 1 h after enzyme addition; c: Difference spectrum; d: Course of the reaction (detection of AFC, ex. 365 nm, em. 470 nm).", + "texts": [ + " 6c shows a slower increasing AMC peak. Without the D-substrate, the reaction of L-PheAMC with PLE was finished after 25 min, with both coumarins available, the hydrolysis of L-PheAMC was not finished after 35 min of the enzyme addition. In the last experiment, the coumarins were again used in different concentrations. 6 ml of L-PheAMC (0.0021 mM) and D-PheAFC (0.021 mM) were placed into the reactor and the reaction was started with the addition of 100 mg PLE (shaken for 20 h). Again, the fluorescence maximum in Fig. 11a was composed of both coumarin substrates at lEX 5 330 nm and lEM 5 420 nm with a RFI of 85. After 1 h, only the D-substrate was hydrolysed by the enzyme, visible by the AFC peak in Fig. 11b and in the difference spectrum 11c. This time, even after 25 h, no increase of the AMC product was detected, that means L-PheAMC was not hydrolysed during the experiment. But the hydrolysis of the D-coumarin was again influenced by the L-coumarin. After 30 min, the reaction showed a constant fluorescence intensity of the product peak AFC (Fig. 11d), this constant intensity was reached without the influence of the L-substrate after 15 min (Fig. 7d). The experiments showed, that an on line monitoring of the quasi-enantiomeric hydrolysis with L-phenylalanine-7amido-4-methylcoumarin and D-phenylalanine-7-amido-4trifluoromethylcoumarin was possible. When the coumarins were hydrolysed without the enantiomeric partner, the kinetic parameters could be received via the time scan runs. The results showed, that the L-substrate was hydrolysed 106 times faster than the D-substrate by a-chymotrypsin" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003038_s0094-5765(01)00002-9-Figure6-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003038_s0094-5765(01)00002-9-Figure6-1.png", + "caption": "Fig. 6. Description of the rotations of the body-Axed frame Fi relative to the preceding frame Fi\u22121: (a) rotation about xi along the length of the module; (b) rotation about the axis zi perpendicular to the length of the module.", + "texts": [ + " 5, the vector Rdmi to the mass element dmi on the ith body can be written as Rdmi =Di + Ti(ri + i i); (3) where Di refers to the inertial position of the frame Fi; ri=[xi; yi; zi]T is the position vector to dmi with respect to the frame Fi (in absence of deformation of the body i) and fi(ri)= i i is the Oexible deformation at ri. Here i are the admissible shape functions and i represents generalized coordinates. The matrix Ti in eqn (3) denotes a rotational transformation from the body Axed frame Fi to the inertial frame F0, i.e. F0 =TiFi. It is deAned by the standard 3-2-1 sequence of the Eulerian rotations i1; i2; i3. The rotation i of the frame Fi with respect to the frame Fi\u22121 has three contributions. Fig. 6(a) shows two components of rotation about the xi-axis: rotation of the actuator rotor ( i), which corresponds to the controlled motion of the revolute joint; and elastic deformation of the joint i ( i1). Figure 6(b) presents rotation in the xi; yi-plane: elastic deformation of the (ith \u2212 1) body ( i2); rotation of the actuator rotor ( i), which corresponds to the controlled rotation of the revolute joint; and the elastic deformation of the joint i( i2). Rotations in the zi; xi-plane can be described in the similar fashion. The three revolute joints, in reality, would be physically distinct. In the formulation, they are taken to be coincident at a point. Thus: i1 = i1 + i; (4a) i2 = i2 + i2 + i; (4b) i3 = i3 + i3 + i: (4c) Let i = i1 i2 i3 (5) be the vector of three Eulerian angles indicating the orientation of the rotor of joint i with respect to the inertial frame F0", + " platform and manipulator units) in the system oj vector from frame Fi\u22121 to frame Fi without the eIect of Oexibility O(N ) order N Pc matrix assigning the Lagrange multipliers to the constrained equations q set of generalized coordinates leading to the coupled mass matrix M, eqn (1) q\u0303 set of generalized coordinates leading to the decoupled mass matrix M\u0303 q\u0303i set of generalized coordinates, associated with the ith body, leading to the decoupled mass matrix M\u0303i qs speciAed component of q Q vector containing the external non-conservative generalized forces, eqn (1) ri position vector of the elemental mass dmi with respect to Fi in absence of deformation of the body i Rai inertial position vector of the actuator located at the ith joint Rd Rayleigh dissipation function for the whole system Rdmi inertial position vector to the mass element dmi located on the ith body R;Rn ;Rv transformation matrices relating q\u0307 and \u02d9\u0303q t time T total kinetic energy of the system Ti rotation matrix mapping F0 onto Fi Ti torque provided by the slew-actuator located at the ith joint Ve total strain energy of the system Vg total gravitational potential energy of the system Greek symbols i; i; i rotations of the frame Fi caused by the control action of the actuator located at the ith joint i1; i2; i3 angular motion contributed by joint Oexibility in three rotation directions deAned by Fig. 6 i vector containing time dependent generalized coordinates describing elastic deformation of the ith body /i actuator rotor angle of the revolute joint i with respect to Fi\u22121 i inertial orientation of the actuator rotor on the ith joint true anomaly of the system - vector containing the Lagrange multipliers ! Earth\u2019s gravitational parameter i rotation of Fi caused by the elastic deforma- tion of the (ith \u2212 1) body, Fig. 6 i(xi; li) matrix containing spatially varying shape functions for the ith body platform\u2019s pitch angle i inertial orientation of the frame Fi deAned by the Euler angles i1; i2; i3 angles between Fi and Fi\u22121, Fig. 6 A dot above a character refers to diIerentiation with respect to time. A boldface italic character denotes a vector quantity. A boldface roman character denotes a matrix quantity. Subscripts \u2018p\u2019 and \u2018d\u2019 correspond to the platform and deployable link, respectively. Subscript \u2018s\u2019 refers to the slewing link or a speciAed coordinate." + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002054_ip-b.1989.0009-Figure3-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002054_ip-b.1989.0009-Figure3-1.png", + "caption": "Fig. 3 Commutation capacitor voltage waveforms in the 120\u00b0 CSI and the SCCSI", + "texts": [ + " The SCCSI does, however, have the advantage of lower voltage ratings, which are principally determined by the magnitude of the motor back EMF and capacitor voltage. Note that the magnitude of the motor back EMF between the devices in the SCCSI is due to a single motor phase, rather than two motor phases in series in the star-connected conventional 120\u00b0 CSI drive. The device voltages in the SCCSI are therefore 1/^3 times those of the 120\u00b0 CSI. The capacitor voltages have been analysed in the Appendix for both systems, and the resulting waveforms are shown in Fig. 3a for the 120\u00b0 CSI and in Fig. 3b for the SCCSI. Fig. 4 shows the peak capacitor voltage and total commutation time against the phase-angle , between the phase voltage and current for each system operated at rated flux and 50 Hz stator frequency. The equivalent circuit motor parameters for the 10 kW experimental motor, used to compute all the results of this paper, have been given earlier in Reference 25. A Rated flux operation, zero to rated load a Peak capacitor voltages b Total commutation time Motor back EMF is E sin (cot + ) a) is resonant frequency of commutation cct a Commutation capacitor voltage in 120\u00b0 CSI t, = 1" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002246_iros.2000.893240-Figure6-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002246_iros.2000.893240-Figure6-1.png", + "caption": "Figure 6: Error bounds for coupled system with wn = 8.66. The error bounds in frequency and amplitude are R = {w = [9.6,10.2], D = [0.78,0.9]}, being the estimated values 2ir = 9.9 and D = 0.83. The real values are w = 9.88 and D = 0.85.", + "texts": [ + " It should be stressed here that the graph is build using analytical equations, and not simulation results. This way, stability properties may be inferred with high precision, because the equations may be solved for as many values as the desired. As easily seen in Figure 2, at the intersection with the origin, equation (2) is satisfied, because av/aw > 0, a u i a w > 0, aulae > o and a v / a D < 0. Stability may also be demonstrated by decomposing the system 1 + N(D,jw)G(jw) = 0 + l/G(jw) + N ( D , j w ) = 0. Thus, for 8 = 9.9, U = N, + ( l /G)u, and V = N , + (l/G)v (a plot of N is shown ahead in Figure 6), the same result as before is obtained, i.e., the limit cycle is stable. = -2mw/k = -0.26 < 0, e < 0, e M 0 %/+ = b / k = 0.067 > 0, % M 0, !?& > 0. Stability analysis can also be evaluated using the limit cycle criterion [?I, (the proof can be found in [5] ) : - 1547 - Limit Cycle Criterion: Each intersection point of the curve of the stable linear system G(jw) and the curve - l /N(D) corresponds to a limit cycle. If points near the intersection and along the increasing-D side of the curve -1/N(D) are not encircled b y G(jw), then the limit cycle as stable" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003479_1.1648967-Figure4-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003479_1.1648967-Figure4-1.png", + "caption": "Fig. 4. Curves mz(cy) and = f(\u03b1) for wing 1 with deflected and undeflected flaps (for 1\u20133, see Fig. 2).", + "texts": [], + "surrounding_texts": [ + "It was noted [1\u201314] that the dependences of aerodynamic forces and moments on the angle of attack for rectangular wings with aspect ratios \u03bb = 1 and 5, as well as for aircraft with high-aspect-ratio straight wings, exhibit hysteresis at Re \u2264 4 \u00d7 106. The shape of the hysteresis loop boundaries was shown to depend on the Re [3], profile thickness [4], aspect ratio [5], wing surface roughness [13], and slip angles [5]. Based on visualization data, the flow structures under the test conditions corresponding to the outer and inner boundaries of multiple hysteresis were analyzed [1, 5]. In this work, emphasis is on the effect of the rectangular wing profile curvature on the hysteresis loop boundaries." + ] + }, + { + "image_filename": "designv11_65_0003576_j.triboint.2004.01.004-Figure2-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003576_j.triboint.2004.01.004-Figure2-1.png", + "caption": "Fig. 2. Reflection and transmission of plane waves.", + "texts": [ + " In addition, the Reynolds condition at the film rupture boundary and the JFO condition at the film reformation boundary are adopted to predict the cavitation region of the lubricant in the bearing. A cavitation algorithm [8] is used to implement the above cavitation boundary conditions, and the finite difference method, together with the column method [9], is also used for the analysis. The numbers of grid points in the circumferential and axial directions for half of each fluid film are 151 and 23, respectively. For the sake of simplicity, it is assumed that the acoustic wave for the oil pressure fluctuation is transmitted in the radial direction through the bearing, as shown in Fig. 2, and the acoustic energy loss in the bearing is neglected. In addition, the effect of the cavitation noise generated in the cavitation region was neglected in predicting the bearing noise transmitted at the outer surface of the bearing. Because the acoustic characteristic impedance of the bearing is very much higher than that of the cavity, the noise generated in the cavitation region can be nearly reflected at the inner surface of the bearing. The averaged sound pressure level transmitted to air in the radial direction at the outer surface of the bearing can be defined as [7]: Nb \u00bc 10log 1 A \u00f0 A 100:1N dA \u00f04\u00de where N is the sound pressure level transmitted to air in the radial direction at the outer surface of the bearing, and it can be expressed as: N \u00bc 20log pft pref \u00f05\u00de where pft is the pressure amplitude of the wave transmitted to air at the outer surface of the bearing, and pref is the reference sound pressure, standardized at 20 10 6 N=m2" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002093_0094-114x(90)90118-4-Figure3-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002093_0094-114x(90)90118-4-Figure3-1.png", + "caption": "Fig. 3", + "texts": [ + " There is no moment on spherical paris, therefore, when the first link is an input link, the force on the main pair which is a spherical pair may be solved using the following simpler form: k04 T - I r F34 -= ( [ G \u00a2 , 3 ] ) T . k , ( 3 6 ) where = [[G~,,3]: , [G613];2 [G4,t3]:3], (37) )w T~,, = T, (k) + TI, ~ T2(k) T3(31 . (38) k = l k = 2 After the reaction of the main pair is solved, the forces and moments of other pairs are easy to calculate [9]. M.M.T. 25/2--D 172 H. B, WANG a n d Z, HUANG 4. E X A M P L E The spatial seven-link 7R mechanism with three parallel joint axes $4, Ss, $6 is shown in Fig. 3. Its input and output angular displacements are 0t and 07, respectively. Input angular velocity is ~1 = 2 1/s. The center of the second link has a centralized mass m23 : 1 kg. The masses of other links are not considered. The dimensions of the mechanisms are listed in Table 1. Fixed and local reference systems are illustrated in Fig. 3; the Y-axis of fixed reference is perpendicular to ground. The displacement analysis of the mechanism has been given in Ref. [8]. Figure 4 illustrates a set of displacement curves. Figures 5-7 give the corresponding relative angular velocities, accelerations and the reactions of the seventh kinematic pair. In this paper, the formulas of velocities, accelerations and the reaction acting on a main pair are obtained in terms of kinematic influence coefficients. These formulas with their unified form are suited for all one-DOF single-loop spatial mechanisms with different kinematic pairs and with any inputs\u2022 As all formulas contain only multiplication of six- or three-order matrices, it is very convenient to analyze kinematic and dynamic forces of the spatial mechanisms using this method\u2022 1" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002256_proc-140-437-Figure4-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002256_proc-140-437-Figure4-1.png", + "caption": "Figure 4. Ball on plate geometry. Figure 5. BOP tester (top plate removed).", + "texts": [ + " Between nudges its spiral has a pitch of order 0.0003, and varies between lubricants. No ball retainer or cage is used since balls of slightly different sizes under slightly different loads have equal orbit rates in the ball-between-parallel-planes geometry. Thus test balls touch only the two plates and the lubricant most of the time, giving the simplest possible rolling conditions at the ball-plate contacts. The positioning ball is free on its pin so that mostly rolling contact exists In the ball-ball contact as well (Figure 4). However, during the nudges there is increased sliding at the ball-plate contacts. In these tests each ball experlences its repositioning contacts less than 2% of the total test time. A drill press provides the test stand so that the bottom plate is stationary and the top plate rotates. It is driven by a spindle held in the drill chuck, and loaded by weights on the drill feed. Balls and plates are hardened 440C steel, the same material as the actual bearing. New balls are used in each test. Plates are surface ground and lapped between tests to remove any existing track" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003137_3-540-36268-1_28-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003137_3-540-36268-1_28-Figure1-1.png", + "caption": "Fig. 1. Top: KUKA KR150 industrial robot with spot-welding tool. Bottom: Calibrated test load", + "texts": [ + " This section discusses the implementation and experimental validation of the above mentioned robot load identification procedure on a KUKA KR150 with a calibrated load. Section 4.1 describes the considered test case. Section 4.2 discusses the experiment design and parameter estimation. Section 4.3 summarizes the experimental results. The test case consists of a KUKA KR150 industrial robot equipped with a calibrated test load. This robot, which is used in industry mainly for spot welding applications, has \u00ff rotational joints. Figure 1 shows this robot with a spot-welding tool (left) and the considered calibrated test tool. Table 1, second column, shows the \u00fe \u00ff inertial parameters of the test tool. These parameter values result from the CAD-model of the test load and from weighing the different parts of the load. The unknown parameters for the considered robot load identification are: the \u00fe \u00ff inertial parameters of the load, the Coulomb and viscous friction parameters, and the parameters that model the efficiency of the transmissions" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001810_0094-114x(84)90051-x-Figure12-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001810_0094-114x(84)90051-x-Figure12-1.png", + "caption": "Fig. 12. Forces between the chain and sprocket neglecting centrifugal forces.", + "texts": [ + " Binder [1] presents the geometric progression load distribution (GPLD) analysis to determine the tension in the roller chain links, based on the following assumptions: (1) mechanical clearances are not present, (2) chain pitch is exactly equal to sprocket pitch, (3) friction is neglected, (4) clearances between the pin and bushing are not present, (5) the pressure angle is constant, (6) there is no elastic deformation, (7) the chain and sprocket are considered as rigid bodies, (8) the sprocket has a constant angular velocity, (9) the driving strand between sprockets is horizontal (or vertical) for all positions of the driver sprockets, (10) the line of sprocket tooth-reaction is on the line of the tooth pressure angle, ( l l ) the chain weight is neglected, (12) centrifugal force is neglected, (13) the chain is aligned with the sprocket, (14) the rollers are bedded. Binder defines the average tooth pressure angle qb and the articulation angle a for a new chain by the following relations: 1200 6 = 350 - - - (1) N 360 \u00b0 a = (2) N where N is the number of sprocket teeth that meshes with the chain in one complete revolution of the sprocket. Figure 12 shows the forces between chain rollers and sprocket teeth and defines the tension in the chain links and presents the free body diagrams for two chain rollers. Binder derives the folowing relationships, using the force triangle and the law of sines: t, to to sin + sin(180 \u00b0 - a - +) sin(a + +) where t. is the force in the nth link. Load distribution for steel chains on steel sprockets 455 to sin t l - - sin(ct + ~b) tj sin t2 - - sin(et + ~b) t,_ t sin qb t n - - sin(et + ~) t. = to sin(a + 4~) _ ( s i n " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001491_melcon.1998.692450-Figure2-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001491_melcon.1998.692450-Figure2-1.png", + "caption": "Figure 2: The magnetic flux distribution in the crosssection of the prototype IXPM motor for 8, =20\u00b0.", + "texts": [], + "surrounding_texts": [ + "1. INTRODUCTION\nThe doubly-salient variable-reluctance (DSVR) motor has evolved in various forms over the last decades because of its apparent advantages, such as rugged construction, fault tolerance and mechanical robustness. On the other hand, some of its inherent deficiencies, i.e. excitation penalty, control complexity, noise and vibration, have prompted research into the incorporation of permanent magnets into the basic DSVR motor structure [1,2]. One of the new doubly-salient permanent-magnet (DSPM) motor proposed recently [3,4] has essentially the same construction as a DSVR motor but with highenergy magnets placed in the rotor poles. Electromagnetic torque is thereby developed in\n-410 - 0-7803-3879-0 / 98 / $10.00\nmuch the same way as in PM brushless DC motors, but with an important cogging torque component required for holding the rotor stationary when the power supply is switched off. So, the motor, even loaded, is not in danger of being accidentally moved to a new position. It is the aim of the present paper to study the dynamic performance of the proposed DSPM motor using the nonlinear magnetic circuit and finite element calculations.\n2. DSPM MOTOR DESCRIPTION\nFigure 1 shows the cross section of a newly proposed DSPM motor. Its basic structure is similar to that of a three-phase doubly-salient 6/4-pole DSVR motor, except the four pieces 1 of PM, buried inside the rotor poles 2, and thus introduced into the main flux path of the stator windings 3 . High-energy bounded-NdFeB magnets with linear demagnetizing characteristic and 0.6 T remanence are used. Both stator and rotor are built fiom laminations. The rotor pole, unlike that of DSVR motor, serves essentially as PM flux guide. Considerable magnet flux concentration is achieved under the overlapped pole pairs where electromagnetic torque is produced. Each rotor pole cutting accomodates a PM and enables two saturated isthmuses 5 preventing the PM magnetic short-circuit by the rotor pole. The isthmuses are framed with epoxy-resin fillings 6, required for PM fixation and rotor pole strengthening.", + "The windings on the stator are of particular simple form, i.e. two diametrically opposite stator poles 4, having the same magnetic polarity, carry coils connected in series to constitute each phase. Both one-phaseon and two-phase-on feeding schemes are adopted. A significant cogging torque component is also developed, since the airgap reluctance seen by the PM excitation is intentionally configured to vary with rotor position. It is obvious that both hybrid and cogging torques are nonuniform. Nonlinear magnetic circuit analysis and finite element calculations are used in dynamic performance prediction.\n3. DYNAMIC MODELING OF THE DSPM MOTOR\nThe circuit equation which describes the DSPM motor dynamics is :\nwhere ws represents the stator flux linkage. which can be considered as the sum of the flux linkage from the stator winding and the mutual flux PM-stator winding; it is dependent on current and rotor position, i.e.\ndtp Sy/(is,&) d i s Sv(is,t&) d& Clt sic. dt -= - Sa dt (2) -- + Nonlinear magnetic circuit analysis is employed for predicting dynamic performances of the novel DSlPM motor. Salient pole geometry, magnetic saturation, fr;nging and PM demagnetization eflects are fbllly taken into account. The airgap permeance consists of two parts, with the bulk corresponding to the overlapped pole area and the rest being the fringing around the pole. Calculation of the h g e part at various rotor positions is based on the circular arc-straight line method. Its accuracy is improved by introducing a correction term through finite element calculations. In its turn, the permeance of the iron is by nature nlonlinear and needs to be determined iteratively. One-phase-on and twophase-on excitation schemes, at the rated current having a density of 10 \"m', are both considered for the prototype motor. Figure 9, presents the magnetic flux distribution in the cross-section of the prototype DSPM motor for e, =20\". The first term of the s u m in Eq. (2) can be written as\n-411-", + "+(is,&) dis dis - = L(is, a) - ,\n6is dt dt (3)\nwhere LS(is,Br) is the dynamic (incremental) inductance of the stator winding at certain rotor position. Figure 3 shows the flux linkage for the three phases of the DSPM motor. It was calculated by nonlinear magnetic circuit analysis based on the equivalent magnetic circuit of the motor for 90 rotor positions and 13 current values. The dynamic inductances were calculated for the same number of rotor positions as\nLs(is,&> = ly.(is, k + 1, a) - rys(is, k, a>\n- (4) i s , k + 1 - is, k Figure 4 shows the winding inductances\nfor one-phase-on excitation of the three phases of the motor. The values of the phase winding inductance and flux linkage for one-phase-on ( Figure 5 ) and two-phase-on ( Figure 6 ) excitations were obtained by interpolation. In the case of two-phase-on stator excitation, the DSPM motor dynamics is described by two voltage equations: usl = RA, + -, dwi dt\n( 5 )\nThe flux linkage of the phase i is dependent on stator currents, is, and isj, and rotor position, so that it can be written : dpi G p i ( i s l , i s j , 6%) disi --- - dt S i s l dt 6~l(isl,isl,&) dis, 8pl(isl,isj,&) d& (6)\nThe partial derivative of the flux sa dt\n+ -+ &SJ dt\n6 ~ l ( i s l , is,, 6%) Si, linkage, represents the mutual\ninductance between the two phases, i and j , d p j dt Lg(isi, isj, Or). The expression for - is similar to Eq. (6) . The dynamic model of the DSPM motor also involves the expression of the electromagnetic torque . It is derived as follows based on a per phase physical model of the DSPM motor, which neglects secondary effects such as saturation, demagnetization and iron losses. The terminal voltage equation for an active stator phase winding ( for one-phase-on excitation ) is\nus = Ris i- e x e = d W, 7. (7)\nHence, dp 6y(is,a) dis GV(is,&) d& dt 6iP dt 68 dt\ndt 664 dt\n- -+ e = - - -\ndis &(is,&) d 8 * (8) = L,-+ -\n- 412-" + ] + }, + { + "image_filename": "designv11_65_0001548_bf00938605-Figure5-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001548_bf00938605-Figure5-1.png", + "caption": "Fig. 5. Stochastic pursuit-evasion game in the plane.", + "texts": [ + " Thus, the equations of motion of E and P are given by dXE=DECOSOEdt+o'EdW1E, t>O, (99a) dyE=VEsinOEdt+o'EdW2E, t>0 , (99b) and dxp:VpCOSOpdt+o'pdWlp , t>O (100a) dyp= Vpsin 0p dt+o'p dw2p , t>O, (100b) where (XE, YE) and (xp, yp) denote the coordinates of players E and P, respectively; tr E and fro are given positive numbers; and Wp={Wp(t)=(W,p(t), W2p(t)), t->O} are RZ-valued standard Wiener processes. It is assumed that WE and Wp are mutually independent. JOTA: VOL. 51, NO. 1, OCTOBER 1986 155 by P and by considering the relative motion of players P and E in polar coordinates, where r is the range from P to E and/3 the bearing of E from P (see Fig. 5), we obtain the following equations for r and/3: dr = [ VE COS( 0 E --/3 ) -- /)p COS( 0p --/3 )] dt + cr cos/3 dBl + o\" sin/3 dB2, (101) r d/3 = [ VE sin(0E --/3) -- Vp sin(0p - /3 ) ] dt - o\" sin fl dBl + cr cos/3 dB2, (102) where B = {B(t) = (Bl(t) , B2(t)), t > 0} is an R2-valued standard Wiener process satisfying crB,(t)=trEW~E(t)-trpW~p(t), t>-O, i= 1,2. (103) Equations (101)-(102) can be written as (by using the same reasons as given in Remark 4.1 in Ref. 3) dr = [ VE COS(0E --/3) -- Vp COS(0p --/3)] dt+ or d W1, (104) d/3 = [(rE~r) sin(0E--/3)--(Vp/r) sin(0p--/3)] d t+( t r / r ) dW2, (105) where is an RE-valued standard Wiener process" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000425_physrevlett.70.1445-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000425_physrevlett.70.1445-Figure1-1.png", + "caption": "FIG. 1. Geometry of an arbitrary confocal texture with vir tual focal surfaces (drawn as cusps outside the cell). The whole texture can be entirely deduced from one of its lamellae.", + "texts": [ + "\" In practice, it is reasonable to accept as confocal any texture with both a negligible dila tion, i.e., e<^ 1, and a ''passive'' dilation energy, i.e., Let us now consider a smectic-^ limited by boundaries that the lamellae can freely intersect without feeling any positional constraint. In the absence of any imposed lamellae dilation, it is natural to look for an elasticity in volving only curvature energy in the continuum of confo cal textures. To simplify, we consider a smectic limited to a planar cell \\\\{x,y) of thickness d (Fig. 1). The exter nal forces are an external field Ellz coupling to the smec tic dielectric anisotropy As, and two plate ' 'anchorings\" YxiO) and yiiO) assumed to favor only preferred orienta tions but to give no positional anchoring [17]. We look for an equilibrium confocal texture assumed invariant along y, with focal surfaces all virtual and well outside the cell. Integrating Eq. (2) along a generator of length L gives e\u2014laiX/a)^, a being the curvature variations length scale. Then condition (3) reads l^<^{a^/X)'^. Since / < a ', passive dilation requires only (4) Under this condition, the dilation energy can be neglected {j Be^<^ J Ko^) and the free energy per unit length reduces to [18] \"^ ^ [ 2 2 4;r J i = \\ , 2 ^ is) Our confocal texture with virtual focal surfaces is entirely determined by the shape of one of its lamellae; it can be equivalently parametrized by the angle Six) at which the lamellae intersect the x axis (Fig. 1). From confocality, the lamella curvature radius <7~' varies linearly along the generators. Calling C, the coordinate along the genera tors, we have ( j~ ' (x,C) = c r ~ K x , 0 ) \u2014 i\u0302 , CIT=CIX(\\ \u2014 ^a)^dcosO, dSi=dx{\\ \u2014 Qo) (/ = 1,2), and the boun dary curvatures cj/ (/ = 1,2) verify ai \u2014 cy\\=^la\\a2 and ar ' -}-cj2~'=2a'~*|^=o. Without any assumption on the lamellae shape, the free energy (5) can be integrated along the generators: (3) Fleix)]- {dxco^e\\\\Kcj\\n\u2014^ Z / = 1,2 CT/ C O S ^ where Yi{e)-yi{e) + \\^E^d^\\n^e", + " This value is (laM^^'^--\\0 times lower than the Helfrich-Hurault [1,19] so-called \"ghost' ' transition. In practice, due to permeation [1], the above discussed transitions should be slow (unless the smectic order is melted near the boundaries [20]). The equilibrium confocal textures may sometimes at tract the focal surfaces and yield the nucleation of focal conies. First, we expect the most general confocal texture to have virtual surfaces degenerating into real lines as they penetrate the smectic bulk (Fig. 2). This raises some interesting problems. Consider in Fig. 1 the pene tration of a focal-surface cusp. Along the cusp generator, since da/ds =0, we have d^a/ds^= -cj'^d^{cj~^)/dO^. From confocality, d^(cj~ ' ) /9^^ = A is a (macroscopic) generator's constant; then d^a/ds^ = \u2014 Aa^. Integrating Eq. (2) up to a distance r ( == cr ~ ' ) to the cusp, e diverges as \u2014A>\\.VA*^, and the stability condition (3) requires r^ rc'^(AX)^^'^. Closer than r^ which is a mesoscopic distance, the texture is no more confocal and dilation will repel the singularity (unless it is topologically required)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000765_s0378-4754(97)00149-3-Figure5-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000765_s0378-4754(97)00149-3-Figure5-1.png", + "caption": "Fig. 5. The SRM.", + "texts": [], + "surrounding_texts": [ + "As we have seen (Fig. 3), the initial instant of the perturbed trajectory must coincide with the corresponding instant of the steady-state. By omitting details [6], we can obtain from (8) and (13): Xn 1 A1 Xn B1 Un (22)" + ] + }, + { + "image_filename": "designv11_65_0003998_1.1839923-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003998_1.1839923-Figure1-1.png", + "caption": "Fig. 1 Schematic view of test three lobe hybrid gas bearing", + "texts": [ + " 18 for a discussion of the test rotor and gas bearings, operating conditions, and major experimental findings. In addition, the reader may find it useful to consult Ref. 19 for detailed coverage of the research program on gas bearings, including test results for other bearing configurations. The test gas bearings are modeled using a FE computational program for solving the ideal gas thin film Reynolds equation 17 . A rotordynamics software suite 20 provides the stability and response predictions for the studied rotor-bearing system. Gas Bearing Model. Figure 1 shows the test hybrid hydrostatic/hydrodynamic gas bearing consisting of three bearing lobes, each with a pair of feed orifices at the lobe apex. The gas film thickness h is a function of the nominal clearance C , pad preload (rp), pad offset angle ( p), and journal center displacements (eX ,eY). The Reynolds equation defines the pressure field P in an ideal gas film with constant viscosity 17 , i.e. \u2022 h3P 12 \u2022 P \u2022R 2 \u2022 x Ph t Ph RgT \u2022m\u0307OR A , (1) where (m\u0307OR), a mass flow source due to local hydrostatic feed, is a function of the pressure ratio P\u0304 P/Ps and the orifice geometrical configuration see Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003674_j.mechmachtheory.2004.05.002-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003674_j.mechmachtheory.2004.05.002-Figure1-1.png", + "caption": "Fig. 1. The reference and current configurations of a rigid body.", + "texts": [ + " Such a combination of translation and rotation is called a screw displacement [22]. Moreover, a screw displacement is understood here as if the corresponding body under motion were fastened to the bolt of a screw whose axis is the screw axis. In what follows, we will apply the concept of screw displacement in order to specify the particular displacements associated to three types of kinematic pairs, namely, helical, prismatic and revolute. To this end, we first present a general derivation of a finite screw displacement, which is built on [21,23,24]. Shown in Fig. 1 is a graphical representation of the general displacement of a rigid body between its reference B0 and its current configuration B. Moreover, we assume that the screw axis of such a displacement is represented by line LQ. It should be noted that, for a general displacement of a rigid body, the screw axis does not necessarily pass through the origin O of the fixed frame, i.e., in general, krQk 6\u00bc 0. From the geometry shown in Fig. 1, we can formulate the following equations: e \u00bc Qe0; r \u00bc rQ \u00feQ\u00f0r0 rQ\u00de \u00fe dk \u00f01\u00de where Q is a proper orthogonal matrix [25] which represents a rotation about axis eQ. Moreover, it should be noted that the translation vector d of an arbitrarily selected point, fixed on the body, can be conveniently decomposed [24] into vectors parallel dk and perpendicular d? to the rotation axis eQ, as shown in Figs. 1 and 2. In fact, Fig. 2 may be considered as a particular representation of the displacement corresponding to a helical joint because of it was assumed a screw displacement, and, for that reason, vector dk depends on the rotation angle /" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001505_icec.1998.699751-Figure5-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001505_icec.1998.699751-Figure5-1.png", + "caption": "Figure 5: Robot Puma 560", + "texts": [ + " Then the fitness of every individual is calculated using Equation (8). Now an abort cirterion must be examined. In this paper this is the maximum number for the generations gmaz. If this number is reached, the GA stops and the individual with the best fitness is the solution of the inverse kinematic transformation for the target TCP. If this number is not yet reached, the genetic operators reproduction, crossover and mutation are used to generate another population. 6. Tests For the tests of the proposed method the six-degreeof-freedom robot Puma 560 is used. Figure 5 shows a three-dimensional illustration of the robot Puma 560 and its six rotational axes. One target TCP is used for the examples. Table 1 shows the data of the target TCP. In all tests the following values are used: 6.9 7.0 -77.3 -98.2 -103.3 -186.0 -184.5 -185.9 5.9 -70.2 6.3 100.4 176.2 18.3 -4.1 -16.9 1.0 -27.6 167.6 47.4 172.9 -29.0 176.0 22.7 The last column in Table 3 shows the error function values of every configuration. These values proof that the multi population GA computes every configuration with a sufficient numerical precision" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001383_978-3-642-60832-2_17-Figure9-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001383_978-3-642-60832-2_17-Figure9-1.png", + "caption": "Figure 9: Mechanical field weakening scheme for DSPM machine magnets in such a machine can also", + "texts": [ + " \u2022 (1) c Q) 2 \"0 ~ Q) 3: 0 1 a.. 0 0 2000 4000 6000 Rated Mechanical Speed (rpm) Comparisons between 4-pole induction machines and 8-pole axial flux toroidal PM machines, Figure 5, is presented in Figure 13. These curves indicates the following ratios for design of axial flux PM machines to squirrel cage induction machines [16] at ns =1500 rpm ~AFl'PM / ~IM = 2.64 (17) at ns =6000 rpm (18) Comparisons obtained between 4-pole induction machines and a 4-rotor-pole/6stator-pole clouble salient PM machine of the type shown in Figure 9 were carried out in Ref. 15. The results are shown in Figure 14. The results for designs at the two sample speed points are as follows at ns =1500 rpm (19) at ns =6000 rpm (20) 219 The comparisons between the 4-pole induction machines and the 4-rotor-pole / 6-stator-pole double salient PM machine with field weakening capability, typified by Fig 10, can be computed in the same manner. At the two sample speed points, [15] at ns =1500 rpm ~DSPMFW / ~IM = 1.36 (21) at ns =6000 rpm (22) It is clear from the results presented in the previous section that this new family of converter fed machines can be a serious competitor with squirrel cage induction machine used for drive applications with a predicted increase in power density ranging from 140 to 327% when compared with typical induction machines" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000007_piae_proc_1922_017_075_02-Figure17-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000007_piae_proc_1922_017_075_02-Figure17-1.png", + "caption": "FIG. 17.-Front hub for 25/50 car. Alternative arrangements.", + "texts": [], + "surrounding_texts": [ + "374 THE INSTITUTION OF AUTOMOBILE ENGINEERS.\nThe point arises here as to how the hand of the gear should be selected for the bevel or crown-wheel so as to impose the heavy thrust on one or the other. The conclusion just reached regarding the double-purpose unit at the crown-wheel is on the assumption that the light thrust load oome's in this position, and this is in accordance with general practice. In order to make full use of the possibilities of the single-row bearing, however, the hand of the gear might be reversed, and certainly there will, in many cases, be more opportunity for obtaining a large capacity in the crownwheel position than behind the bevel-pinion\nPilot Bemirig. Although the pilot bearing, Fig. 20, does not appear on any of\nthe graphs, it will be remarked that it is in many casea an& example of a bearing applied to relieve a heavily loaded one without kaking into account the possibility of its own failure, which frequently occurs.\nClutch. The bearing selection for the clutch is not so ainenable to the same treatment, but it will usually be found that a light type single-row bearing w i t b u t fillibg-slot has su5cient capacity .to deal with a load which is only in occasiond operation.\nFigs. 17 to 20 show the aotual arrangements of the front-hub, rear-hub. gear-box and rear-axle bearings of the car referred to as No. 12.\nSUMMARY. The difficulties in the way of fixing definite factors of safety which will meet all cases are fully appreciated, but it is hoped that the preceding recommendatipns will form a useful basis for investigating existing designs. For convenienoe, the factors have been collected in Table I., the letters in Fig. 15 qorres o d ' with those used thmughout the drawings of the Sizaire-brm? chassis.\n2016 at Kungl Tekniska Hogskolan / Royal Institute of Technology on June 5,pau.sagepub.comDownloaded from", + "THE ENDURANCE ,OF BALL HEARINGS. 375\n2016 at Kungl Tekniska Hogskolan / Royal Institute of Technology on June 5,pau.sagepub.comDownloaded from", + "376 THE INSTITUTION OF A UTOMORILE ENGINEERS.\nFIG. lg.-Gear-bux for 25/50 h.p. car.\n2016 at Kungl Tekniska Hogskolan / Royal Institute of Technology on June 5,pau.sagepub.comDownloaded from" + ] + }, + { + "image_filename": "designv11_65_0003839_3-540-26415-9_101-Figure3-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003839_3-540-26415-9_101-Figure3-1.png", + "caption": "Fig. 3. Graphical 3D viewer", + "texts": [ + " The two links between the three modules are actuated with two pneumatic pistons [6]. 2 Alicia 3 Simulator \u2013 global description The Alicia3 simulator developed in Simulink\u00ae allows testing locomotion strategies, but can be also used to tune parameters for locomotion control for the entire robot, as well as for the base system module Alicia II (Fig. 2.). A graphical 3D viewer, developed in Delphi \u2122 7 using the GL-Scene free package and a Mex-function to integrate it in Simulink \u00ae , visualize Alicia 3 sequence of operations (Fig. 3.). Basic blocks that represent the behaviour of the main system components have been implemented. Connecting together all these blocks, a realistic model of the Alicia 3 robot behaviour has been obtained. The implementation of these basic blocks will be described afterwards. For each block, a set of parameters has been de- fined and can be adjusted according to system components. In the Simulink \u00ae diagram there is the possibility to modify all these parameters. The control algorithms implemented in the simulator are the same of that of the real robot" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002081_1521-4109(200203)14:6<405::aid-elan405>3.0.co;2-c-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002081_1521-4109(200203)14:6<405::aid-elan405>3.0.co;2-c-Figure1-1.png", + "caption": "Fig. 1. Dimensionless square-wave voltammograms of the reversible redox reaction complicated by the reactant adsorption on the spherical electrode surface. A) z 100 and B) z 1000; 0.1 (1), 1 (2), 3 (3), 5 (4), 10 (5), 30 (6) and 500 (7); Eacc E 0.2 V, tacc 80 /f, b 0.01, a 0, n 1, Esw 30 mV, E 5 mV, Est Eacc (the scan runs from right to left).", + "texts": [ + " In square-wave voltammetry (SWV) the dimensionless response of fast and reversible redox Reaction 1 depends on five dimensionless parameters: the adsorption constant of linear isotherm z maxf1/2D 1/2, the relative surface adsorption capacity maxf1/2D 1/2(c ox 1, the electrode sphericity b D1/2r0 1f 1/2, Frumkin coefficient a and the relative duration of accumulation period tacc N/f, whereN is a certain number. Note that the formal stripping scan rate is v f E and the scan duration is (Eacc Efin)/f E. The parameter is inversely proportional to the bulk concentration of the reactant. It determines the relationship between the diffusion controlled part of the response, or the main peak, and the response of the surface redox reaction, or the post-peak. This is shown in Figure 1, for Langmuir isotherm (a 0). In case of very short accumulation (tacc 80/f), well developed and clearly separated main peaks appear if 10 and z 1000. These values correspond to c ox 10 4 mol/L and max 0.3 cm, if max 5 10 10mol/cm2, D 10 5 cm2/s and f 100 Hz. For many surface active ions and organic substances the value of the product max is between 10 2 and 1 cm [5]. The forward (reductive, red) and backward (oxidative, ox) components of the responses consisting of the post-peak and the main peak are shown in Figure 2", + "Within these limits of , thedimensionless post-peak is diminishing as is decreasing and the real net peak current of the surface reaction is not a linear function of the bulk concentration of the reactant. In the narrow range 5 3 the ratio ( P,net/ )post-peak is constant, meaning that the real peak current of the post-peak is independent of the reactant concentration. This is a consequence of the surface saturation. The separation between the main peak and the postpeak depends linearly on the logarithm of the parameter z, with the slope EP/ log z 2.3 RT/nF, if z 100.However, for rather weak adsorption (z 100) these two peaks are unresolved. The responses shown in Figure 1A possess a single maximum, but the main wave, or the post-wave may appear as a shoulder (see also Fig. 2A). Curves 2 \u00b1 4 in this figure show that the variation of value causes the shift of the potential of the maximum response for 80 mV. This is also shown by curve 1 in Figure 3. The potential of the main peak is 0 V vs. E . A negative logarithm of corresponds to the logarithm of the reactant concentration. Curve 2 in Figure 3 shows the potentials of the post-peak and the main peak for strong adsorption (z 1000, Fig. 1B). With the development of the main peak (5 1) the post-peak is shifted negatively for 40 mV. Under the influence of lateral attraction in the adsorbed layer (a 0), the separation between the diffusion and adsorption peaks increases, as can be seen in Figure 4. The difference between peak potentials is linearly proportional to Frumkin coefficient: EP/ a 15mV. Figure 5 shows that the influence of attraction is not simply additive to the influence of the adsorption constant. At very low surface coverage the attraction is negligible (compare curve 7 in Fig. 5 with curve 7 in Fig. 1A), but as the relative bulk Electroanalysis 2002, 14, No. 6 concentration of the reactant increases the post-peak is shifted negatively for 80 mV. This is shown as curve 3 in Figure 3. Under the conditions of full coverage two small post-peaks may appear (see curve 2 in Fig. 5). Figure 6 shows that this phenomenon is caused by the separation of small and narrow reductive component from bigger and wider oxidative component of the response of surface reaction. Hence, the continuous shift of the post-peak in negative direction, together with the splitting of the postpeak, are the indications of lateral attractions in the adsorbed layer", + " The difference Epzc can also be either positive, or negative because the reactant can be either a cation or an anion. There aremany possible combinations of these variables. To obtain a qualitative information about the influence of capacitive current on the total response, that is valid within an order of magnitude, a standard set of values is chosen: C 0 30 F/cm2, C 1 C 0 20 F/cm2, Epzc, 0 E 0 V, Epzc 0 V, n 1, max 5 10 10mol/cm2, z 1000, 3,b 0.01, a 0 andEsw 30 mV.The faradaic response of the corresponding redox reaction is shown as the curve 3 in Figure 1 B. The dependence of the adsorption induced component of the dimensionless capacitive current on the change of the potential of zero charge is shown in Figure 11. The capacitive current is significant in the potential range of the post peak, inwhich the adsorbed reactant is reduced, but in the range of the main peak it is negligible, because of d / dt 0. Comparing to themaximumof the faradaic response, the maximum capacitive current makes only about 4% of the total response. If Epzc, 0 E 0.5 V and the other parameters are as in Figure 11, the maximum capacitive current increases to 17% of the maximum faradaic current", + " The influence of various C* on the capacitive current is shown in Figure 12. If the change of electrode charge Electroanalysis 2002, 14, No. 6 capacity is positive, the current is negative and its minimum is about 14% of the maximum faradaic response. Figure 13 shows the influence of the variation of the bulk concentration of the reactant (or the relative surface adsorption capacity ) on the dimensionless capacitive current c and dimensionless total response F c. The parameters of redox reaction are as in the Figure 1B. The biggest capacitive current appears when the surface coverage is low (the curve 6 in Figure 13A), but its minimum makes only 2% of the maximum of the total response. Figure 13B is almost identical to Figure 1B. This shows that the adsorption induced capacitive current does not distort the square-wave voltammograms significantly. The distortion may appear only if the manolayer of the adsorbed reactant is condensed [37, 46]. It is assumed that a metal ion M2 forms four labile complexes with the ligand X and that only a neutral complex MX2 is adsorbed to the mercury electrode surface: 3 MX2 4 ne , max, a (A1) M (Hg) (MX2)ads In a great excess of the ligand, the concentrations of free metal ions M2 and the labile complex MX2 are: cM cM tot 1 B (A2) cMX2 cM tot W2 (A3) B 4 j 1 Kj [X ]j (A4) W2 1 B K2 X 2 (A5) cM,tot cM cMX cMX2 cMX3 cMX4 (A6) Kj is a stability constant of the complex MXj and [X ] is a bulk concentration of the ligand X " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003324_icsmc.2003.1244217-Figure7-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003324_icsmc.2003.1244217-Figure7-1.png", + "caption": "Figure 7: Two bicycle states evaluated as different", + "texts": [ + " 5 is used by the planning algorithm to verify if the current state has already been explored (it is in closed-states) or is already in the list of states pending exploration (it is in open-states). If the state has not been explored but it is part of the list of states pending exploration, then the algorithm checks whether its cost needs to be updated based on the current path. If the current state is part of neither the list of closed-states nor the list of open-slates, then the state is added to the list of open-states and set for pending exploration. In Fig. 6 and Fig. 7 we show a graphical representation of how the function in(state, list-of-states) evaluates whether a state is present in a list of states or not. In Fig. 6 we see two states considered as equal. The consideration is based on the fact that the Euclidian distance between the two center of gravities is less than a given position threshold and the difference of the two heading angles is less than a given angular threshold (heading threshold). Fig. 7 instead shows the situation when two states are not considered equal because either one (or both) of the two differences is greater then the corresponding threshold. In the next section, we proceed by showing some simulation results for the planning algorithm. 5 Simulation Results Motion path results for two different situations are shown on Fig. 8 and Fig 10. Fig. 9 and Fig. 11 show the feedfonvard control values found for both situations. In Fig. 8, the system moves from the intial point (0, 0) to the goal point (1, 6) avoiding an area of slippery surface (dark rectangle)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002201_iemdc.2001.939386-Figure10-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002201_iemdc.2001.939386-Figure10-1.png", + "caption": "Fig. 10. Control Volume for the Rotor Cooling Gaps", + "texts": [ + " These losses are artificially introduced into the ventilation characteristics by many authors (e.g. [2], [l], [4]). The shock loss coefficient is usually used to fit the theoretical ventilation 5of8 characteristic to the experimental data. That is why the coefficient assumes values in the wide range from 0.5 to 12. For the case of axial bores in the rotor there is no need to use this approach. Simply applying the conservation For the pressure increase from Position 1 in Fig. 9 (Rotor Inlet) to Position 3 in Fig. 10 (Rotor Exit) 6 and 11 yield: 1 2 equations leads to a relatively accurate determination of the losses associated with the transition of the cooling air p 3 - p1 = - . p . ( U 3 2 - U 2 2 ) + c . m2 (12) into the rotor. The second Part of the Pressure change in the rotor takes place in the rotor cooling gaps. The Control Volume is shown in Fig. 10. In 7 the conservation of rotational moAll pressure drops that depend on the mass flow have been sw\",rized in the last term of 12. For the pressure rise without any m a s flow follows: mentum in the axial direction without frictional effects is stated for the rotating frame of reference: Applying this to the control volume shown in Fig. 10 leads to an equation for the moment acting on the fluid: . . 1 2 ( P 3 - Pl)p ,o = 5 . p . ( U 3 - U Z 2 ) Or expressed in dimensionless parameters: + ( c p = O ) = (1 - (2)') (14) The ratio of mean inlet to outlet radius of the test rotor was 0.61. For $I (cp = 0) follows the value of 0.632. Fig. 11 shows the comparison of the rotor characteristics with and without the modifications for the case without axial mass flow. The improvement is significant for high mass flow rates. For low mass flow rates (cp -+ 0) both curves converge to the limiting value of $I = 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001826_s0021-8928(01)00025-9-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001826_s0021-8928(01)00025-9-Figure1-1.png", + "caption": "Fig. 1", + "texts": [ + " After this the problem of synthesizing the optimal control both of the swinging and damping of the double pendulum is solved without using the Pontryagin maximum principle. The problem investigated here is of interest both from the point of view of theoretical mechanics and also for modelling the motions of a gymnast on a horizontal bar. Such motions have been investigated in a number of publications (see, for example, [6] and the bibliography given there). Consider a plane double physical pendulum, suspended at the point O by an ideal (frictionless) cylindrical hinge (Fig. 1). The sections, which are absolutely rigid bodies, are also connected to one another by means of a cylindrical hinge at the point D. The axes of the hinges are horizontal and parallel to one another. The centre of mass of the first section will be assumed to be situated on the section 0i) . We will denote by m~, Ii, rl, I~, the mass of the first (upper) section, its length OD, the distance of the point O to its centre of mass and its moment of inertia about the point O, respectively. Suppose m~, l~, r~, I~ are the similar quantities for the second section" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002913_1.1479336-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002913_1.1479336-Figure1-1.png", + "caption": "Fig. 1 Coordinate systems of the rotor-bearing system", + "texts": [ + " 124, OCTOBER 2002 Copyright \u00a9 rom: http://gasturbinespower.asmedigitalcollection.asme.org/pdfaccess.a dynamic analyses ~@13#!. This paper used polynomial curves for eccentricity distribution with finite element modeling in the eccentricity identification derivation of shafts. The merits of our approach are the derivation is straightforward and easy to incorporate into the existing finite element rotordynamic programs. Finite Element Modeling The coordinate systems used in this paper are depicted in Fig. 1, where XYZ is the fixed frame, and xyz the rotating frame. The rotor rotates about the z-axis with an angular velocity V counterclockwise. u, v , w are the displacements of the rotor in the X, Y, Z directions, ux and uy the rotational angles in X and Y-axes. Figure 2 shows a cross section of the shaft with mass center M c , geometric center Gc , eccentricity e, and phase angle of the unbalance fu . Assuming the eccentricity curves are finite, piecewise continuous, and of m-degree polynomials, the local eccentricity distribution for each shaft element f (s) can be expressed as f ~s " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003221_50011-7-Figure11.2-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003221_50011-7-Figure11.2-1.png", + "caption": "Figure 11.2. Stanford robot", + "texts": [ + "- The concatenation of the equations of several planes in a unique system of equations increases the number of identifiable parameters [Zhuang 99]. The use of three planes gives the same identifiable parameters as in the position measurement method if the robot has at least one prismatic joint, while four planes are needed if the robot has only revolute joints [Besnard 00b]. * Example 11.1. Determination of the identifiable parameters with the previous methods for the StMubli RX-90 robot (Figure 3.3b) and the Stanford robot (Figure 11.2). The geometric parameters of these robots are given in Tables 11.1 and 11.2 respectively. For the plane constraint methods, we assume that the plane coefficients are known. Tables 11.3 and 11.4 present the identifiable geometric parameters of the two robots as provided by the software package GECARO \"GEometric CAlibration of Robots\" [Khalil 99a], [Khalil 00b]. The parameters indicated by \"0\" arc not identifiable, because they have no effect on the identification model, while the parameters indicated by \"n\" are not identified because they have been grouped with some other parameters" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000076_0045-7825(93)90131-g-Figure12-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000076_0045-7825(93)90131-g-Figure12-1.png", + "caption": "Fig, 12, Equivalent stress assessment (t = 300 s),", + "texts": [ + " Example of simulation of a torsion-tension test We consider the same geometry as in Fig. 4. The material obeys the rheoiogical law of (49). The rotational velocity is low (ltr/min). A tension velocity is prescribed at the other extremity of the sample (1 mm/min). The thermal computation shows that deformation is nearly isothermal. Time steps are constant and equal to 5 s. As can be found in the analytical solution of the tension test, the dependency of the longitudinal velocity on z is linear (Fig. 11). The distribution of equivalent stress (Fig. 12) 432 A. Moal et al., A simulation of' the torsion and torsion-tension tests A. Moal et al., A simulation of the torsion and torsion-tension tests 433 shows that the deformation remains homogeneous in the longitudinal direction of the worked part of the sample. The reduction of section is approximately constant along the specimen. Comparisons with experimental results (evolution of torque and force with the deformation) could be done in the future in order to evaluate the validity of the numerical computation for the simulation of torsion-tension tests" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000230_0020-7403(94)90070-1-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000230_0020-7403(94)90070-1-Figure1-1.png", + "caption": "FIG. 1. T w o - l o b e bearing.", + "texts": [ + " in each lobe of the bearing and an initial temperature distribution are assumed so that the oil viscosity at the nodal points may be calculated. Using the assumed attitude angle and temperature distribution, the solution scheme initially determines the nodal pressures (from the solution of Eqn (1)] for the flow field. Using iterative procedures the solution scheme establishes: (a) the boundaries of a positive pressure zone satisfying the condition (OP/~O) = 0 at the trailing edge of the positive pressure zone in each lobe (Fig. 1); and (b) the equilibrium locus (i.e. attitude angle) of the journal centre for a vertical load support. Nodal pressures obtained after the establishment of the film extent and the attitude angle are used to establish the temperature field by iterating the energy and the heat conduction equations. The iteration between the energy and heat conduction equations is terminated when the boundary conditions at the bush and oil interface is satisfied within the specified tolerance. The temperature then obtained is used to modify the viscosity" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003221_50011-7-Figure11.1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003221_50011-7-Figure11.1-1.png", + "caption": "Figure 11.1. Description of frame RQ relative to frame R.j", + "texts": [ + "3] The nominal value of Pj is zero. If Zj.i and Zj are not parallel, pj is not identifiable. We note that when Zj.j and z, are parallel, we can identify either rj.] or rj (\u00a7 11.4.2), thus the number of identifiable parameters for each frame is at most four. 11.22. Parameters of ihe base frame Since the reference frame can be chosen arbitrarily by the user, six parameters are needed to locate the robot base relative to the world frame. As developed in \u00a7 7.2, these parameters can be taken as (YZ, bz, ttz, dz, Oz, t^ (Figure 11.1): Z = -iJo = Rot(z, Yz) Trans(z, bz) Rot(x, Oz) Trans(x, dz) Rot(z,e2)Trans(z,r2) [11.4] The transformation matrix \"^Ti is given by: -iTj = -^To \u0302 1 = Rot(z, Yz) Trans(z, bz) Rot(x, a^ Trans(x, dz) Rot(z, Gz) Tran$(z, t^ Rot(x, aO Trans(x, di) Rot(z, GO Trans(z, ri) [11.5] Since ai = 0 and di = 0, we can write that: -^To ^ 1 = Rot(x, Oo) Trans(x, do) Rot(z, Go) Trans(z, ro) Rot(x, a*i) Trans(x, d\\) Rot(z, G'l) Trans(z, r'O [11.6] with 00 = 0, do = 0, Go = Yzj 0\u0302 = ^z,a'l -Oz, d\\ = dz, &x = Gi + Gz, r*i = ri + TZ Equation [11" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003626_1350650041323368-Figure5-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003626_1350650041323368-Figure5-1.png", + "caption": "Fig. 5 Test specimen configuration and plan view of arrangement of contacts", + "texts": [ + " Two piston treatments, tinplating and anodizing, were also compared for their resistance to microwelding. A Plint TE77 high-frequency reciprocating tribometer was adapted to simulate the percussive impacts between the components, both with and without sliding. It was desirable to use samples that had the actual metallurgy and topography encountered in service, so test specimens were sectioned from production pistons and rings. A mechanism was designed to load and slide a sample of piston groove against piston ring sections located in a stationary oil bath, as shown in Fig. 5. To recreate suitable contact pressures with the Plint TE77, the groove width was narrowed to 1.0mm and the piston ring was replaced by three 10mm long sections held in a lower piston ring section holder, essentially an annular groove machined in a baseplate. Three clamps each hold a section of piston ring 10mm long at 1208 intervals around the groove (Fig. 5). The groove is 2mm deep, allowing a small volume of oil to cover the ring sections. The piston sample holder has a spherical ball-and-socket joint, applying the same load to each contact. The fulcrum lies in the plane of the sliding contacts to eliminate any turning moment that could distribute the load unevenly. The standard constant vertical loading arrangement of the TE77 tribometer was replaced with an eccentric Table 1 Common conditions of combined sliding/percussive tests Test frequency (Hz) 25 Test duration (h) Until observation of material transfer Period of cycle (s) 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000533_bf02192245-Figure3-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000533_bf02192245-Figure3-1.png", + "caption": "Fig. 3.", + "texts": [ + " P u )[:2, the parallel gradient algorithm for this case is specified as follows: ~+1 = A~'(p')(p~), ~+1 = A~'(p')(p~). For a particular set of r~, rz, {7'}~=I, the game evolves as shown by the dots in Fig. 2 (right). Each dot with its associated stage t indicates the status of the game for that stage. It is observed that the evolution of the game corresponds to a very fast decaying oscillation around the Nash equilibrium point (p*, p*). Proposition 2.1 states that there is upper limit on the stepsizes of the players when they are resorting to parallel gradient descent. Figure 3 (left) JOTA: VOL. 90, NO. 1, JULY 1996 55 (L) Nash equilibrium; (R) parallel gradient descent with r~ = 0.5, r2 = 0.~, 7' = 5. Parallel gradient descent behavior with respect to stepsize: (L) parallel gradient descent with rj=0.8, ~z=0.8, ~,'=5; (R) parallel gradient descent with r~=0.1, r2=0.1, 7 '=5. shows the evolution of the game when the players use larger stepsizes. As expected, there are more oscillations and the amplitudes of the initial oscillations are larger. Note that, in this case, the position at time = 1 has fallen outside the range shown in the figure. Making the stepsizes smaller can get rid of the oscillations as seen in Fig. 3 (right). Proposition 2.2 states that there is a lower bound on the iteration time so that a given error level can be achieved. Increasing the iteration times Parallel gradient descent behavior with respect to iterations at each stage: (L) parallel gradient descent with r~ =0.1, Vz=0.1, ?,t= 15; (R) parallel gradient descent with r~=0.5, r2=0.5, 7'=1. can also compensate for the small stepsize as seen in Fig. 4 (left). However, there is a practical limit to how large the iteration times can be chosen as it will increase the magnitude of the oscillations" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001018_0954411981534619-Figure2-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001018_0954411981534619-Figure2-1.png", + "caption": "Fig. 2 Joint coordinate system for seven degrees-of-freedom model of the human upper limb. Successive rotations between coordinate systems are described by Denavit\u2013Hartenberg parameters as outlined in Table 1", + "texts": [ + " reference (5)] is all but removed. The standard Denavit\u2013Hartenberg notation (6) was used to define the location and orientation of joint axesThe MS was received on 16 May 1996 and was accepted for publication on 27 May 1997. (Tables 1 and 2, Fig. 1). The joint coordinate system for H03796 \u00a9 IMechE 1997 Proc Instn Mech Engrs Vol 211 Part H at Kungl Tekniska Hogskolan / Royal Institute of Technology on July 5, 2015pih.sagepub.comDownloaded from defined as being along the Z i\u22121 axis (see Table 2 and Fig. 2). This procedure is continued until all remainingFig. 1 Definition of Denavit\u2013Hartenberg parameters joints and links have been defined.for definition of linked segment coordinate Body segment parameters for human biomechanicalsystems modelling can be determined by a variety of techniques, including direct measurement and regression equations (7). A software program (8) which calculates body seg-a three link, seven degrees-of-freedom, three-dimenment parameters based upon the results of Zatsiorskysional model is shown in Fig. 2. The technique used does and Seluyanov (9) was used to calculate parameters fornot impose any restrictions on the complexity of models individual subjects. Body segment lengths, masses andwhich can be specified. Coincident degrees of freedom locations of the centre of mass for the upper arm, fore-(such as those found in the human shoulder) are repli- cated by using a link length of zero. arm and hand were used to define the linkage geometry. at Kungl Tekniska Hogskolan / Royal Institute of Technology on July 5, 2015pih" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000785_s0094-114x(98)00074-3-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000785_s0094-114x(98)00074-3-Figure1-1.png", + "caption": "Fig. 1. Edge of regression of a spur involute surface.", + "texts": [ + " The envelope Sr to the family of tool surfaces is represented in Sr as [4, 6] r u; y;f ; f u; y;f 0: 1 The working part of the tool surface Sr is a regular one, and ru ry$0. Important contributions to the study of parametric envelopes are represented in di erential geometry [8] and in theory of gearing [3\u00b16]. Our investigation is directed at the discovery of singular points on Sr that may form: (1) an edge of regression; (2) an envelope Er to contact lines (characteristics) on Sr that is simultaneously the edge of regression. Fig. 1 shows an edge of regression in the case of spur involute gears. Fig. 2 shows that, in the case of an involute helical gear, the edge of regression, the helix on the base cylinder, is simultaneously the envelope Er to contact lines on the helical tooth surface. Contact lines on generating surface Sr may also have an envelope designated as Er. The main di erence between the envelope Er from Er is that Er is formed by regular points of the generating surface, while envelope Er is formed by singular points of generated surface Sr" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001138_1.2833218-Figure7-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001138_1.2833218-Figure7-1.png", + "caption": "Fig. 7 Case 3, radial ioad acting on bearing due to component rotation", + "texts": [ + " But the ratio of the net displacement to rolling element radius has changed little in the two cases considered here. We have included the effects of inertias of different bearing components and found that equations of motion have nonlinear and time-varying coefficients. If the application involves very high speed then the inertia of the rotor is also to be considered. 2.3 Case 3. Bearing as an Interface for a Rotating Com ponent. This case simulates the most common application of a bearing. A radial force is allowed to act on the bearing due an external rotating mass Mf, attached to the outer ring (see Fig. 7) . Hence the line of action of force is rotating with the outer ring. This results in continuous change in the position of the load zone around the circumference and the number of rolling elements carrying the load. The inner ring is fixed and the outer ring is moving. The bearing of Case 1 is used here. The expression for the radial force is given as F = MfRf^l cos (^J + MfRf^l sin (/>,,/ (17) The position vector of Mf from the global origin O is rf = (Rf cos 4)1, + Xi,)T+ {Rf&ia cf),, + y^)! (18) The virtual displacement is given by 6rf = {Sxh \u2014 Rf sin (l)i,6(l)i,)i + (Syh + Rf cos cj)i,64>i,)j (19) 326 / Vol" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000166_9.256404-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000166_9.256404-Figure1-1.png", + "caption": "Fig. 1. Multiple robot arms handling a common object.", + "texts": [ + " Under some reasonable assumption, it is shown that the number of saturated actuators on any finite time interval along the optimal trajectory is 1 + 3 ( 0 - 1Xrn - 1) in which D is the number of dimensions (two or three) and rn is the number of robots in the system. Furthermore, some remarks on computational algorithms for the numerical solution and a numerical example cited from [16] are given in Section V to verify the theoretical result developed in this work. Finally, conclusions and suggestions for further research are presented. 11. REDUCED-ORDER DYNAMIC MODEL: A CLOSED CHAIN FORMULATION In this work, we consider multiple robot arms holding a common object as a single mechanical system which consists of kinematic closed chains, as shown in Fig. 1. Suppose that rn robots are used to handle the same object in order to lift it, turn it, or transfer it, etc. It is apparent that the dynamic behavior of the robots is dependent upon each other because of the kinematic and dynamic coupling among them. Our main objective here is to derive a dynamic model of the whole system for control purposes. For convenience, we first make the following assumptions regarding the dynamics of the multiple robot arms which will be used throughout the paper. Assumption 2" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003375_icmech.2004.1364455-Figure5-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003375_icmech.2004.1364455-Figure5-1.png", + "caption": "Fig. 5. Energy optimal motion with increazed safety margin.", + "texts": [], + "surrounding_texts": [ + "I. INTRODIJCT~ON Optimal control of robotic manipulators will be an issue in fut,ure rnaniynlation systems, even though it is hard to be found in today's work processes. A manipulator traject,ory may he optimal w.r.t. several criterions. Such criterions are the overall energy consumption, the path lcngt,h, the manipulator stiffness or dexterity. Commonly energy optimal and shortcst path cont,rol are of major interest. The main reason why optimal control is virtually insignificant for practical applications is the complexit,y of the accompanying mathematical optimal control problem (OCP). There are basically t,wo approaches to solve the OCP, direct and indirect methods. Indirect, methods attempt to solve the corresponding Euler-Lagrange equations, a two point boundary value problem. The numerical solution of thc later is a very ambitious t,a& Boundary value solvers essentially pursue shooting methods. For nonliiicar control problems, as in case of robot manipulators, single shooting methods are likely to fail. Advanced niultiplc shouting algorithms [3];[12] may be able to solve the boundary value problem provided that the nuiriber of multiple shooting nodes is sufficiently high. This. however, tends to increase the complexity of the underlying optimization problem dramatically, even for simple examples. In this paper the direct method [1],[2],(19],[22] to solve the original OCP is applied. Instead of solving the corresponding Enler-Lagrange equations the OCP is taken as it is and approximations of the generalized coordinates of the nranipnlator are determined that minimize the respective objective. Boundary constraints are incorporated as eqrialit,y constraints. Thus the OCP is restated as algebraic nonlinear optimization problem subject to equality constraints. n'ith a finite approximation generally a near optimal solutiou C a l i be obtained. Its quality depends on the approximation order. Howevcr: since for practical applications an almost optimal is ;ts good as ttrc theoretically optimal cont,rol t,he direct approach is a computational efficient. strategy for optimal manipulator cont,n,l. Yet there is no established methodology fur collision avoiding vptirrial control of manipulat,ors. Crucial is an adequate math- 'The autlior thanks Prof. P. hlaisser for his manifold suppoit 0-7803-8599-3/04/$20.00 02 04 IEEE 299 ematical description uf spatial objects that can be easily incorporated in the OCP formulation. More precisely a continuous mcasure for the distance of spatial objects is necessary. Artificial potential functions were used for obstacle modelling in path planing schemes [S] ~ [ 101 [l l] ~ [la] ~ 1251, [26]. .4 repelling potential is assigned to each olatac!e and an attractive potential to thc goal positiori. The combination of both yields an overall potential field in which the manipulator moves. The manipulator is 'controllcd' by potential forces originating from the attraction and the repelling potentials respectively. The time integrating of the motioii equations yields the driving forces that steer the manipulator from its initial to the goal configuration. However, with t,hc addition of multiple obstacle potentials there may bc local miniina causing the manipulator to stall before reaching the terminal configuration. Also the particular trajectory depends on the given duration of niot,ion and on the magnitude of the potential forces, that have to chosen hy chance. Thereupon opt,imal control is not achievable by this approach. But nevertheless artificial pot,entials facilitate the definition of an adequate distance measure between objects. This paper is organized as follows. Section 2 reviews the kincmatics and dynamics of robot manipulators. Particular attention is paid in section 3 to the description of spatial objects and distarice measures with artificial potent,ials. This distance measlire is applied in section 4 to collision detection. The actual energy OCP and its accompanying nonliuear programing problem (NLP) is derived irr'scct,ion 5. Section 6 out,liiies the numerical solution via t.he direct approach. The paper concludes with two examples in section 7. 11. ROBOT KINEMATICS :ANT) DYNAMICS A serial manipulator (SM) is a holonomic multibody system (MBS) consisting of a single kinematic chain with 71 bodies connected by n joints. In the following only rigid bodies and ideal joints are considered. All n rigid bodies B,: a = 1 ; . . . ~ 7~ of the SM are niimhered in increasing order starting with the ground Bo. Usually the terminal body B, caries the nianipulator's end-effector (EE). Rigid body configurations are described by homogenous matrices. They coristitute the Lie group S E (3) generated by its Lie algebra se (3 ) [16],[23]. The configuration of B, is expressed by the niatrix The rotation matrix R E SO (3) describes t,he relative rotation of a B,-fixed refererice frame (RFR) w.r.t. to an inertial frame (IFR): or world frame. The position vector of the RFR origin expressed in thc IFR is Pt R3. With the assumption that all t,echnical joints can be assembled with one-DOF joints the SR4 contains nn revolute and n p prisrnatic/screw joints. Their joint variables qa; a = 1: . . . , U are generalized coordinates on the configuration space V\" = I\"\"\" x E?\"'. On B,-I is also dcfined a where J is the inertia matrix of B, w.r.t. to its RFR. The Lagrangim motion equations (LME) of a force cont,rollcd Sh4 are where Fah, := (&gac + a,y,b - a,,,,) are thc Cliristoffel symbols of first kind, Q. the generalized forces and uu the gcneralked controls. Due to (5) the LME can be given derivative free in closed form [14],[17]. It sliorild be not,iced that the LME (7) or their generalization, the Bolzmann-Hamel cquations, of an electromcchanical system with elastic bodics are consistently of the form (7) [15],[13]. 111. OBSTACLE MOUELLINC Denote with p = (p1>pz>fJ) t R3 the position vector of a point in E3 cxpressed i n the IFR. Boundaries of three dimensional solids are implicitly described as v-ellipsoids in E', determined hy thc shape function g a d + r ,4bqc = Q. + uo: (7) joint frame for joiiit U. Expressed in this B,_,-fixed joint franre is the screw vector X E W6 of joint a. The matrix X E ae(3) is giveii in terms of the screw vector X via' X = X j Ej, where E l ; . . . ~ Es is t,he basis for se ( 3 ) dednced from the joint frame. The transformations between 6-vectors and .se (3)-mat,rices are denoted with X = ; and X = X . V = (U? U ) is the twist vector. Denote with A1 the coiist,ant transformation from the RFR to tlic JFR on B,. The product of exponentials (POE) mapping C: V\" + SE(3) [17]:[16],[20] D e - a n ( L a XVU C ( q ) = M eTq' . . M e* , y E SE (3) ?E se ( 3 ) , (2 ) yields thc configurat,ion of B,. The end-effector (EE) attached to B, is represented by an EE frame. Let M be the transformation from the RFR on terminal body B,, to its EE frame. Their the EE configuration is E = C M . Throughout the paper A.f = I is assumed for simplicity. The body velocity V = C-'C of B, is [I71 n e e a \" V = K u j b , with Ka =Ad,-,.(X);b 5 a (3) 0 0 a b b Therein Ad : SE (3) x se (3) - se (3) is the .4djoint operator and Ka denotes thc kinematic basic fnnction (KBF), also termed the geometric Jacobim. V t se(3) contains the skew synimetric angular velocity tensor 3 t so ( 3 ) and the linear velocity vector The EE velocity is Adn,-I(V). The left invariant metric on SE ( 3 ) : i.e. a metric for which ( X , X ) = X T G X , 2 E se ( 3 ) , is independent from the particular IFR, is r n This metric, howevcr, depends on a scding fiLct,or u / p that scales rotation vs. translations. In vector form of V the KBF of D, is the matrix K = E W6.\". The end-cffector DOF of the SM is E := maxyeVN rank(K(q)). The partial derivative of the KBF is given algebraically [I71 Kb 1 8, K b = [Ka, KC]: b < c 5 a.: ( 5 ) D m a where the Liebracket on se (3 ) is [XI Y ] = XY - YX, or in vector form [X,Y]\" = [ux x W Y , W X x WY + wx x WY]. The mctric G gives rise to 2r left invariant distance measure d (.4,B) = l/(logR,'RB)\"ll + P [IPA ~ Pa/ ) > (6) A? B t SE (3). Here log R = & (R - RT) is the log fnnction on SO ( 3 ) ) wit,h the angle 6 of the rotation determined by R. The kinetic energy T ( q , 4) = fg.a@\"db of the MBS is given in terms of the generalized ~iiass mat,rix where fi? i = 1 , 2 , 3 is a function for the length of the i th serniprincipal curis. Such function werc called super-quadratic [ll] If f; '= c, are constant and rri = 1 then the set of points for which S = 1 yields an ellipsoid. For m going to infinity this becomes a box with respective side lengths Zc,. The factor .$ stretches the m-ellipsoid so that it touches the corner of the box. Thus S = 1 yields the surrounding not the inscribing ellipsoid of the box. With appropriately chosen fi the shape function Scan describe almost any mirror symmetric shape. For large m the contour S = 1 very closely follows the varying semi-principal axis length fi ( p ) . Values S > 1 correspond to points outside the object while inside S < 1 and S = 0 is the object center. With the pseudo-distance 6 ( p ) := S ( P ) k - 1 (9) the object bomidary is the set of points for which 6 = 0. The pseudo-distance is only a proper distance nieamre outside the object since it takes negative values inside. Figure 1 shows the contour 6 = 0 and two contours with 6 > 0 surrounding an object with order m = 3. Having found a suitable dist,ance meanre a potential field can he straightforwardly assigned to the rcspective object. This potential is assumed to attain a certaiii finite value A which is used for collision detection. The potential has range [O. A]. The paranleter D > 0 determines trow fast the potential decays with increasing pseudo-distance. The psendodistance between a point p and body B, is where the shape function of the body B, is This is simply the shape function S transformed in the moving RFR of Ba using the rotation matrix R (y) and the position vector p(q). I.e. the shape fnnction and thus the pseudo- distance of a point arid B, depend on the configmation q: exccpt for the ground Bo. To body B, is %signed the potential Pa ( q ; p ) := A,eCD\" (13) Several such potentials with corresponding shape functions are used to describe complicated geometries. On t,he gronnd in particular several objects constitute the environment. Iv. POINT WISE COLLISION DETECTION Each body carries a potcntial ficld. In ordcr to dctcct contact of inutually colliding solids, say B, and Bb, it would consequent,ly be necessary to considcr t,he corrcsponding potentials Pa and PL in every point in their proximity. A collision would occur if in a point bot,h potentials simultaneously attain a certain value depending on t,he parameters A and D. Clearly a point wise evaluation is in practice impossible. Sensor points ,axe defined on each body. If B, carries 1, such points thc sensor points are denoted with p,, , a,, l l . . . , l a . The potcntial fields arc only evaluated at these points. Generally it is not necessary that a sensor measures every potential: cspecially not that of the body to which it is fixed. A measure graph r dctcrmincs which potentids a sensor nicasurcs. If a sensor ub on Ba is measnring the potcntial Pc of B, then edge (a:b) t r. It is natural that the potential of two bodies are mutually measurcd. So r is nondirected and each edge of r corresponds to a pair of possibly colliding bodies. denotes thc number of edges of r. The order rn determines the gap between the object's actual surface and t,he houndary 6, = 0. The gap only vanishes for m i CO: but the boundary is always surro~inding the object. The parameter D , dctcrmincs tlic 'visibility' of B, from far distance. It is not iinportaiit, for the condrol strategy introduced here. A, can be det,ermiried such that at a certain pseudodistance A,, from the boundary 6, = 0 the potent,ial Pa attains a certain value that indicates contact with B,. This value is arbitrary and set t,o 1, so that (14) A ._,DaA= Thus P,, (y ,pyl) 2 1, (a, b) t r if the pseudo-distance of B, and sensor point a b on Bb is less or equal to A,. This provides a safety levcl for the collision detection. a .- V. THE COri'mOi. PROHI.EL~ A broad class of optimal control problems for miilt,ibody (and general mcchatronic) systems is described as optimal control problem of a system of differential cquations subject to equality and inequality constraints. Such probleins have the general forni .I (U) i inin ( a ) 4% (4, i, 4: t ) = U,, i = l _ . . . ; n ( b ) (OCP) f j ( q : < j ) = 0 , j = 1 . . . . , T (C) Pa((rl:pob)- < 1, v t E [ O , T l , ( a , b ) E r , (4 lqal 5 4&: a = 1 ( e ) T n J ( U ) := [ wou:dt, (15) I where the objective fuIict,ion is with positive weighting coefficients wa. The differential eqnations (a) describe how thc system dynamics is controlled by generalized driving forces us. These equations arise from the LME of a general mechatronic system. A solution q ( t ) of (OCP) satisfies the system of differential equations (b ) while minimizing the objective (a). This objective is usually! but, not neccssarily, a quadratic form in U,. Resolving (7) for the driving forces yields (16) P a := gab4 ..L + rczbc\u00b6bi' - Qa. Therewith a solution of (OCP) is a motion with minimal ovcrall driving forces. The r equality constraints (c) include initial and terminal donstraints on t,he states and/or constraints on the EE-configuration but also interniediate constraints on the states and/or EEconfiguration at 0 < t < T . It is customary to require the manipulator to attain a desired initial and terminal state, usually it should he at rest. This gives 411 boundary constraints f ' , i = 1,. . . ,4n in terms of the generalized state variablcs (f') := ( Y (0 ) - yo; q (TI - YT; 4 (0 ) - h, 4 (T) - 4 ~ ) Often it, is more convenient to provide t,tre terminal E E configuration. Especially in case of redundant manipulators the configuration is not unique nor is it clear which configuration from t,he self motion rnanifold 151 is best in terms of the objective J . Alternatively the 371. + 1 boundary coiistraint,s require the terminal EEconfiguration E (T) hut not the terminal configuration q (T) of thc SM. Here d ( E ( T ) , E T ) is the distance nieasure (6). In principle the initial configuration inay also be replaced by t,he EEconfiguration but this is usually not nicaningfnl since the Shl should start from a well defined posture. If the SM is redundant the (OCP) solubion is such that the overall energy consumption is minimized. Collision avoidance, according to the sensors at the manipulator and environment, is ensured by the inequality constraints ( d ) . Throughout tlic paper thc t.erniinal time T is considered to be fixed. I t can also he a free parameter. VI. NUBlERIC.41. SOLUTION Either direct or indirect methods can be applied to solve OCP. Indirect methods solve the corresponding Euler-Lagrange equations, a two point boundary value problem [4]. Their solution raises considerable difficulties. Even with advanced multiple shooting algorithms [3];[12] this rcmains a challenging task. Here thc direct solution of the OCP is pnrsned. Ansatz fnnctions are defined far the generalized coordinates and the OCP is reformulated as an NLP. Obvionsly the cxact soliition q ( t ) of the OCP may only be found if it lies in the function space spanned by the ansata. In general only an approxiination of the exact solution is achievable. A . NLP refomidation of the optimal contml problem The objective .J ( U ) is approximated by (NLP) 'p, (r;?, i'. t ) = U 0 > a = 1 ; . . . , n, (b') f j ( T : i ) = 0, j = l , . . . , r: (c') T 1 s ([pa (rlP=,) - 11,) = 0, ( % b ) Ob 0 S ([I.\" - q ~ . l l + ) = 0: U = I , . . . , n , r, (d ' ) (e') T associated NLP is much easier than solving the associatcd EulerLagrange equations of thc OCP. even though the large number of parameters. However, the price to he paid is that only an approximatc solution can be fonnd since the solution space is determined by t,he ansatz function basis. This is not necessarily a disadvantage since a good approxiination of the optimal trajectory is snfficient from a practical perspect,ive. The quality depends on the used ansat,z functions. Connuon choices for smooth energy hoiinded q are orthogonal function bases such as Fuuricr-. Chebychev- or wavelet bases. The approximation r converges to the solntion q if it belongs to the space spanned by the ansat,z fnnct,ions. Splines will not have this property. Since the generalized driving Forces U , are dcdnced from an a g proxiinatc solntion r via thc differential eqnations the smootliness of uu depends on that of r and pa. The LME (7) for an nnconstraincd (either holonoinic or nonholonomic) multihody system are hounded w.r.t. to q, 4, q and t t,he snioothness of ua depends on smoothness of T . A crncial part is the placement of a snfficient number of sensor points on the bodies. For simple geometries like boxes this is easy while for more complex objects an automatic placement is preferable, similarly to FEM mesh generation. Advantageously a rcfinenient of the covering with sensors does not increase the NLP complexity since the number of constraints (d') depends on the pairs of ohjects considered for collisiox B. Analytical ezpressions for the gradients For the numerical solution of t,he NLP it is advantageous to have analytic derivativcs of J ' , pa> Pa and f j w.r.t . the parameter. The partial derivatives of re and f' w.r.t. the paranleter A;> 0;; C;: D; are readily found. The derivatives of pa merely correspond t,o the linearized motion equations 1141. Like the LME they can he given algebraically due to ( 5 ) . In order to find derivatives of Pa the KBF is split as With a,aRT= - RT6b and a q b P = Uo [17] it follows that l i a ( I n . . splinc will be nscd. For a cubic splinc ansatz s; ( t ) = Ai + BE (t - t i ;) + C; ( t - tk)' + 0; (t - ti;)3 is a cuhic polynomial defined on the interval ti. < t < t b c l . From the sDline conDenote with T~~ the position vector ofthe sensor point p,, w.r.t. the Bh-fixed RFR. so that vn. = p + R rnr. Then &=v,,. = . . ~ \" b 6 ~\" . .\" - .~ ditions Sk ( t k + l . ) = S k i 1 ( t k t l ) , i k ( t i + l ) = Sk+l ( t k + l ) Si; ( t k + l ) = s k t l (&+I) follows that ol1ly M + 3 coefficients of TO ( t ) are independent. The sensitivity of ra w.r.t. parameter changes sumests to choose D r 3 . . . , DE.,, A;, B:, Ct as indepeii- p + R n = r ~ , b b for b > 0 and a,=p,, = 0 if b = 0 (i.e. p,, is on the ground). The gradient of S ( p ) is easily found from (8). This rives - ._ .. I dent parameters of the spline ansatz T O . The NLP thus condains ( M + 3) . n variables. Cnhic splines yield continuous piecewise linear accelerations F i t ) . To achieve continuous ierks requires a apPa ( q > p m b ) = - D 2 a (q ,Po, )$=Sa (q,Pq). (20) _ I fourth order spline ansatz. It is cnstomary to choose the same numbcr A,f of segments for all P. Only with some a-priory knowledge about the character of the optimal trajectory a particular number may he chosen accordingly. To sum up the OCP is reformulated as NLP, which is an algebraic nonlinear mjnimiiation problem in ( A t + 3 ) . n varishles subject to r + /r( eqnality constraints, where Irl is the nnmher of pairs of possibly colliding bodies. Additional n constraints appear if joint limits (e') are inclnded. The NLP problem is continuous. It should be emphasized that the solution of the 6. Ve@ication of optirnality A NLP solution T of is generally an approximate solution of the OCP. Further, a numerical solution of the NLP may yield a 1- cal minimum and thus only bc a snbopt,imal solution. If q is a solution of the OCP it satisfies the Eoler-Lagrange cquations of the OCP. It can thus he checked wliet,lrer an ohtained solution T is near optimal hy evaluating the Euler-Lagrange eqnations. The deviation from being satisfied constitntes a measure of o p timality. VII. EXAMPLES The approach has been applied to a variety of nonredundant and redundant planar and spatial SMs. For the sake of clarity a planar manipulator kconsidered here. Results for cubic splines are reported in order to elucidate the effect of non-smooth jerks. The NLP is solvcd with a standard package for constrained o p timization [S],[Zl]. No effort was yet made on advanced and tailorcd optimization methods, that shall be expected to reduce the necessary computation time for solving the NLP. Figure 2 depicts the top view of a redundant planar 5R manipulator. The shown object linings are the isosurfaces for 6, = 0 corresponding t,o shape finictions. The five links are considered as aluminium rods l m in length with constant rectangular cross-sections. The task in figure 2 is to cont,rol the EE on BE from the shown initial to the goal configuration in the presence of two squared obstacles. The duration is T = 20. A spline ansata with M = 10 is used. Prescribed is the initial configuration q (0) and the terminal EEconfiguration E (T). So (c') in (NLP) contains the five initial comt.raints qn (0) = 4: and the terminal constraint d ( E ( T ) ,E T ) = 0. The metric scaling a l p = 1 is set in (4). All controls have equal weights, U,, = 1. Considered are collisions of manipulator a i d obstacles according to the measure graph r. Black dots on the bodies indicate sensor point,s. The NLP thus has jrl = 10 additional equality constraints (d'). The detection distances are set to A, = 0. The deuicted maniuulator motion minutes. Higher accuracy yields marginal improvement. Also the necessary run.time increases nonlinearly with the accuracy. Typically with a sliort duration (T = 20) it iz energy optimal to fold the manipulator before turning and afterwards unfolding it. The obstacles mainly interfere within the unfolding motion. Taking into account joint limitations the preceding motions are likely to be impossible. i n figure 3 the revolution of the last four joints is restricted to 450 degree. Thus (e') contains [ / q - q ~ . , J + = 0, a = 2,. . . , 5 . Another scenario is shown in figure 4. The energy optimal control with a safety margin A- = 0 very closely follows the obstacle boundaries. Due to model uncertainties and the finite number of sensors it is nossible that is the result of a numerical solution with a required accuracy of in the objective value. The solution is found within 10 a control scheme using the determined U,, will cause collisions. For safety the collision detection should bc more conservative as shown in fignre 5 . Therein the same task is accomplished with A, = 0.1, so that the manipulator always remains at this distance from the obstacles. The effect of using cnbic splines is rcflected i n the driving forces (figures 6). If non-smooth controls are not desirahle fourth order splines shall be applicd. Then the number of variables in the NLP only increrzcs by n if t,lie conditions S I (ti+i)=Si;+l (ti;+I) are incorporated in the spline defintion. Gcncrally spline functions do not constitute a basis for energy hoiinded functions, like Fourier polynomials do; and so will not always yield the hest approxiniation of the exact solu... t,ion. VIII. SUMMARY A direct approach to collision avoiding energy optimal control of serial manipulator is proposed. The original optimal control problem is restated as constrained algebraic nonlinear o p tiniization problem. The Free variables of t,his prohlem are the paranleter of ansatz functions for the generalized coordinates of the manipulator. Contours of spat,ial objects are described by generalized ellipsoids. Thc presented approach to optiinal control is not restricted to serial manipulator nor is it restricted to rigid body mnltihody systems. The Lagrangian motion eqnations of a general niecliatronic system are consistent,ly expressed hy Lagrangian niotion equation? that govern the dynamics. It is straight forward to extend the approach to manipulators with kiireniatic loops, like parallel manipulators, where the necessarg closure constraints must, be incorporat,ed ils fiirtlier equality constraints. REFERENCES [ I ] V.Y. Berbyuk. h1.V. Demidyuk: Parametric optimization in problems of dynamics and conrrol of an clastic inaiiipulators -ith dialiihuted ~ararnetun. 1. Mechanics of Solids. Vol. 21, 1988. pp. 78~86 J.T. Belts: A direct approach to solving uptinial ~ o n t i o i pioblenm. IEEE H G. Bock: A mii l i iple shooting algorithm for direct ndution of optiinel conlrol probioms. Int. Federation A i i f o m a t i ~ Control, 9th World Congress. Budapest. 1984 141 A.E. B r y ~ u n . Y.C. Ho: Applied Optimal Control. Taylor and Francis, 1975 151 J: W. Burdick. On the inverse i ineinalics of redundant manipulators: Charrctorirstion of the self-motion manifold. Proc. I E E E Int. Conf. 0 1 1 Robotics and Automation, hlay 15-19, Scottrdale. AZ, 1989. pp. 284-270 Y.C. Chen: Solving robot trajectory planning I)roblemr with uniform cubic 8-splines, Optimal Control Applicaticins & Methods. Vol. 12, 1YB1, pp. 247-262 171 S.I. Chni. B.K. Kim: Obstacle avoidance control for redundant manipuiafora using collidability measures. Rohoticn, Vol. 18, ne. 2. 2000, 1 4 3 ~ 1 5 1 181 P.E. Gill. W. Murray. h1.H. Wright: Practical O~t,imisntion, Acadcmir 12) 131 J. Computing in scienca k ErlgineerinG, Vol.,, no.3, ,999, pp. 73-75 161 ~~ Press. Lurrdon. 1981 IS] N. Hogan: Impedance control: An approach to manipulation. J . Oyn. Sys- tems, hlaaisurernents and Control, 107. 1985. PP. 1-21 1101 0. Khatib: Real-Tinie Obstacle Avoidance for hlanloulators axid hlobiie . . R.,botS, i n t . J. 01 nobotics ~ ~ ~ ~ ~ ~ ~ h , ~ ( 1 ) . I B R ~ [ill P. Khasls. R . \\'dpo: Superquadratic Artificial Potentials for Obstacle Avoidance and Approach. Proc. I E E E Int. Cunf. 011 Robotics i n d Automation. Philadelphia, 1988 ple shooting baaed S Q P ~tartegy for large dynamic process optimization. Cornpiiter k Chemical Enginemring, Vol 27. 2003 113) J.E. Slarsden. T.S. Ratiu: Introduction t o mechanics and s).lnnletry, 1141 P. h4uisser: A differential geometric approach io the multibody system dynamics. 2. ang. Math. hlech. (ZAMY). 1991, pp. 118-1111 1151 P. hlaisrer. 0. Enge. H. Freudcobeig, G. Kirlsu: Electrvmechanical Infera ~ t i o n ~ in Multibody Systems Containing Electromechanical Drives. Journal hlultibody System Dynamics. Vol. 1. no. 3. 1997. pp. 281-302 I161 R.hl. hlurrsy, 2 . Li. S.S. Sartry: A msthematical Introduction to robotic bliaipalation, CRC Press. 1993 [I?] A. Muller, P. hlairser: Lie group formuistion of kinematics and dynamics of constrained hlBS and ita application te analytical mechanics. Multibody System Dynamics. Kluwer. Vol. 9. 2003, pp. 311.552 1181 A. hluiler: .< dexterity maximizing ~unf inusf ion nietliod fur the inverse kinematic8 of redundant manipulator, 11th I E E E hlediterranean Coiif. an jizi n . 8 . ~ e i ~ . ~ ~ b o r . I. B ~ ~ ~ ~ . H.G. s a c k , J . P sciliedFr: A,, efficierlt l l l U i t i ~ Springer, 19Y1 Control snd Auromation, Rhodes, 2003 1191 k4.L. Nagurlu. Xr. Yam: Fourier-Based Optimal Conlroi of Nonlinear Dy- namic Svsterns. AShlE Journal of Dvnamical Systems. &lessuremant. and . , Control. Vol. 112, 1990, pp. 17-26 1201 F.C. Park J.E. Bobrow, S.R. Ploen: A Lie Group Forinulrtion of Robot Dynamics. In,. J. Robotics Remarch, Vol. 14. No. 6, 19!15, pp.: 609-618 Calcuiations. Nnmerical Analysis, in G.A. \\\\'atson (ed.)- Lecflire Notre in blatbemrrics. \\'\"I. 630, Springer Veriag. 197H 1221 D. Schmitt:. A.H. Sun;, V. Siiniunran. 6. Naganathim: Oiitirnal hlution Programmrng of Robot Blanipulators, ASME Journai of 8lechanisms. Trn~smirrionr. and Automation in Derign, Vol. 107. 1985, pp. 239-244 [23] J.hl. Selig: Geometrical foundations of robotics. World Scientific. 2900 1241 A Visioli: Trajectory planning of robot manipui i l f~rr by using algebraic and trigonometric spiner, Robofica, Vol. 18. 2000, pp. 611-631 [ZS] R Volpe, P. Khosia: A s t r i l t r g ~ for Obrtrcle Avoidance a n d Approach using Superquadratic Potential Functions. Proc. 5th Int. Symposium of Robotics Research, 'Tokyo. 1989. pp. 93-190 (261 C. Waren: Global path planing using srtificiai potential fields. Proc. IEEE C o d Robotics end Automation, 1989. pp. 316-321 1211 h1.J.n. ~ ~ ~ ~ 1 ; : A ~~~t A I ~ ~ ~ ~ ~ I ~ ~ ~ for N ~ I ~ ~ ~ ~ ~ ~ constrained optimizat ion" + ] + }, + { + "image_filename": "designv11_65_0003302_ip-a-1.1981.0012-Figure14-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003302_ip-a-1.1981.0012-Figure14-1.png", + "caption": "Fig. 14 Crossarm potential detector", + "texts": [ + " Ambiguity therefore arises when an indicator, which is intended to be operated by a local distribution conductor, in fact is operated remotely by a transmission conductor. In other circumstances, a voltage indicator may show that a metal crossarm is live (Fig. 13); the potential may, however, be harmless, for example as the result of unbalanced tiny leakage or capacitive currents, or lethal if it arises from a punctured insulator. A solution here is a limiting resistor used to indicate whether or not the source impedance is dangerously low (Fig. 14). Apparatus intended for checking the insulation of such things as hot sticks both should be able to indicate relatively small changes in large numbers \u2014 a seriously defective 400 kV component may still have a resistance of 10M\u00a32 against a capacitive reactance only ten times larger \u2014 and be able to recognise concealed hazards like internal tracks or metallic inclusions. Even more difficult is to detect a long air-filled void, for example within a foam-filled tube, which is not a hazard until it is either wetted, carbonised or ionised" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001215_pime_proc_1996_210_489_02-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001215_pime_proc_1996_210_489_02-Figure1-1.png", + "caption": "Fig. 1 The TU-3 valve train system. Exhaust components viewed from the front of the engine", + "texts": [ + " The positions of minimum contact velocity also frequently coincide with the initiation of high-rate wear processes such as scoring of the surfaces. This is well illustrated in the Peugeot TU-3 fired engine test (8), the principal European test for valve train wear in the current specifications for gasoline engine oils of the Association des Constructeurs European d\u2019Automobiles Part J: Journal of Engineering Tribology (ACEA). The TU-3 engine has a rocker follower OHC (overhead camshaft) valve train system, which is shown diagrammatically in Fig. 1. It is most informative to consider the occurrence of surface distress on the follower surfaces, as before (3), because the positions of the critical values of the main parameters that are likely to affect wear are better spatially separated on the follower than on the cam surfaces. The appearance of a follower wear track in the very early stages of failure is shown in Fig. 2. This is an electron micrograph of an area approximately 6 mm wide at the extremity of the contact path closest to the axis of rotation of the follower [identified as the E, position in previous analyses (3)]", + "______________ Maximum value in T U - 3 3 - 2 - 0.8 1.4 c I VE I I -0.2 I I 0 100 200 300 400 Crank angle deg Contact velocities in the reciprocating Amsler for a stroke of 5 mm at 400 r/min Fig. 10 Part J: Journal of Engineering Tribology retical oil-film thickness varies over quite a narrow range. The variation of contact duration for one half-cycle of the reciprocating Amsler block, calculated in the same way as for the TU-3 in Section 2, is shown in Fig. l la . The second half-cycle is the mirror image of Fig. 1 la. Values range from 1.5 to 8.5 ms for each half-cycle, with a maximum continuous contact of 17 ms near the reversal point, at x = b. The corresponding sliding distances are shown in Fig. l lb . Values vary between 0.7 and 4.7 mm, with a maximum continuous contact of 9.4 mm. Figure l l a and b is not quite symmetrical about the centre of the stroke because of the effect of the crank system. Values at the outer end of the block (x = 0) are slightly greater than at the inner end, although in this short-stroke condition the differences are barely perceptible", + " However, if their variation over the full stroke of the Amsler is examined, it is found that contact dura- tions and sliding distances fall within the range of the more severe parts of the valve train contact cycle. The maximum values of these parameters calculated for the TU-3 and VE systems are indicated in Fig. l l a and b. In the TU-3 test, scoring consistently appears first in the area within 1 mm of El on the follower (Fig. 2). The range of sliding distances occurring in this area is covered within 0.7 mm of the extremities of stroke in the reciprocating Amsler, as indicated in Fig. 1 lb. One potentially important difference between the conditions in the reciprocating Amsler and in the valve train systems should be noted. In the valve train systems, the oil films are thick enough to separate the surfaces throughout most of one sweep of the cam over the follower surface (for example when the contact is moving from left to right in Fig. 3c), but are very thin during the other sweep. Thus, each point on the follower surface makes intimate contact with only one point on the cam surface, except at the El, El, R, and R, positions, during one revolution of the cam" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003437_0954406042690498-Figure3-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003437_0954406042690498-Figure3-1.png", + "caption": "Fig. 3 Structure of the blade-disc", + "texts": [ + " 2b represents a pinion with concave\u2013concave tooth traces and generated by two blade-discs of which the strips formed by cutting blades are crossed. The reason for designing the pinion as C04304 # IMechE 2004 Proc. Instn Mech. Engrs Vol. 218 Part C: J. Mechanical Engineering Science possessing concave\u2013concave tooth traces is to be able to consider reducing the bending stress. To avoid collision of the cutting blades on one blade-disc with those on the other blade-disc, two blade-discs have to rotate synchronously. The sketch map of the structure of bladedisc is shown in Fig. 3, where (1) represents the blade supporter disc, (2) the cutting blade, (3) the wedge block, (4) the fastening bolt, (5) the adjusting bolt, (6) the sleeve and (7) the fastening bolt. Several rectangular straight slots are radially distributed over the end face of the blade supporter disc. Each slot holds one blade, which is fixed by wedge blocks and bolts. The radial extending length of the cutting blade is controlled by an adjusting bolt, which is installed in one sleeve of a ring. The ring with radially distributed sleeves is fixed on the blade supporter disc by several fastening bolts" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002093_0094-114x(90)90118-4-Figure2-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002093_0094-114x(90)90118-4-Figure2-1.png", + "caption": "Fig. 2", + "texts": [ + " (27) The relation between equivalent force or moment and local equivalent force or moment can be written as 1 6 = (28) This shows that equivalent force or moment is different from local equivalent force or moment. This is the reason why we present the concept of the local equivalent force or moment. At the same time, equation (28) gives another method for calculating the equivalent force or moment. In order to calculate the reactions of all kinematic pairs, the seventh basic pair is chosen as the main pair of the force analysis, as shown in Fig. 2. The reaction F67 = (T6Tv F~7) T of the main pair are decomposed along the fixed reference O-XYZ. If the reaction of the main pair can be solved, the reactions of other paris will be easily obtained. According to the above-mentioned method, the local equivalent force or moment with respect to the generalized coordinate ~. due to the reaction F67 acting on the main pair can be deduced as the following T. (07) = - ([6 ~6];, )T F67. (29) From the principle of virtual work, the work done by the local equivalent force or moment and input force or torque in infinitesimal virtual displacements &b, (n = 1, 2 " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003676_apec.2004.1295970-Figure2-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003676_apec.2004.1295970-Figure2-1.png", + "caption": "Fig. 2 Added linear generator arrangement plan", + "texts": [ + " The harmonic injection traveling-wave filed cuts the linear generator winding to generate the energy that the train needed. An extra linear generator winding is introduced to improve the efficiency of the linear generator. Harmonic injection method can solve the problem of Maglev power supply when Maglev is running in low speed. It needn\u2019t any change to the ground connection and low cost. At first, this paper introduces the harmonic injection method. Then output voltage model of linear generator is given. Finally, experiment results are given. II. PRINCIPLE OF THE HARMONIC INJECTION METHOD Fig. 1 and Fig. 2 show the principle of the harmonic injection method. Three-phase current in the long stator windings are 1( ) 2 sin( )ai t I t\u03c9 \u03c9= , 1 2( ) 2 sin( ) 3bi t I t \u03c0\u03c9 \u03c9= \u2212 and 1 2( ) 2 sin( ) 3ci t I t \u03c0\u03c9 \u03c9= + , where 1I is stator current, \u03c9 is stator current angular frequency, which drives linear motor to move maglev train. Three-phase high frequency alternative currents are added to original stator alternative currents, which are ( ) 2 sin( )ah h h hi t I t\u03c9 \u03c9= , 2( ) 2 sin( ) 3bh h h hi t I t \u03c0\u03c9 \u03c9= \u2212 and 2( ) 2 sin( ) 3ch h h hi t I t \u03c0\u03c9 \u03c9= + ", + " The injection current generates the harmonic traveling-wave field, whose velocity is 2h hv f \u03c4= \u22c5 , where hf is injection current frequency. Cutting speed between traveling-wave field and harmonic traveling-wave field is 2 ( )h hv v f f\u03c4\u2212 = \u2212 . The voltage induced in the linear generator (on-board power collecting windings) varies in amplitude and in frequency in proportion to hf f\u2212 . So the voltage induced in the linear generator can be controlled by the injection current amplitude and the injection frequency. Fig.2 shows the added linear generator on the magnetic pole and the pole span of the added linear generator is equal to the stator polar distance \u03c4 in order to improve the electric power generation efficiency of the linear generator. III. OUTPUT VOLTAGE MODEL OF LINEAR GENERATOR There are two part of the linear generator output voltage. One part voltage 1genU is generated by the stator spline harmonic magnetic field cutting the linear generator winding. The other part voltage 2genU is generated by injection harmonic traveling wave magnetic field cutting the original linear generator winding and added linear generator winding" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002855_s0261-3069(01)00038-3-Figure4-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002855_s0261-3069(01)00038-3-Figure4-1.png", + "caption": "Fig. 4. GS44 Si N FDC bar sample, (a) Schematic of missing roads and (b) X-ray radiographs before WIP\u2019ing.3 4", + "texts": [ + " Two different specimen geometries were used to evaluate the effect of WIP\u2019ing on the elimination of FDC build related defects: (1) rectangular bars; and (2) square plates. The rectangular bars were used to assess the ability of WIP\u2019ing to remove FDC inter-road defects, such as missing roads or incompletely bonded layers. To ensure a sufficient population of defects in the FDC bars, missing road defects were intentionally introduced in the center section of the bars in a staggered position at every other layer. (note: roads are running at \"458 to the specimen edges). Fig. 4a shows a schematic of the defects. One of the typical defects identified in the FDC process is the sub-perimeter void that was frequently encountered early on in the development of the process. These voids, as shown in Fig. 3b, develop in the area between the perimeter (specimen contour) and the vec- tor paths due to an inadequate impingement of the vector paths on the perimeter. Square plates (25.4=25.4=6.35 mm) with a centrally located hole (12.7 mm in diameter) and sub-perimeter voids surrounding the center hole were used to test whether WIP\u2019ing could effectively remove these defects", + " In order to quantify the extent of healing due to WIP\u2019ing, four point bend strengths were measured on some of the sintered WIP\u2019ed rectangular bars. Control specimens without defects were built and WIP\u2019ed under the same conditions. Standard 4=3=50-mm samples were cut and diamond ground from sintered bars. For all of the parts, the binder burnout was performed using methods developed earlier w2x and sintering was performed at Honeywell. Figs. 4\u20136 show the results of X-ray radiography tests conducted on the bend bars and square plate FDC samples, before and after WIP\u2019ing at 708C and 35 MPa. Fig. 4b shows the built-in missing road in both views (at\"458 angles in the top view) along with a schematic showing the missing roads (Fig. 4a). Fig. 5 shows subperimeter and inter-road void defects at the junction between the contours and vectors. From the X-ray radiography results, it was observed that WIP\u2019ing at temperatures of 30, 40, 50 and 608C had no effect on removing the existing build defects. The samples\u2019 shapes and dimensions also remained unchanged after WIP\u2019ing at Ts608C. On raising the WIP\u2019ing temperature to both 70 and 908C, the build defects in the FDC parts could no longer be detected in the X-ray radiographs. As can be seen from Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002008_12.474436-Figure6-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002008_12.474436-Figure6-1.png", + "caption": "Figure 6. More HUR-Badger Platform Functions. (Note: Model does not show tracks.)", + "texts": [], + "surrounding_texts": [ + "Once inside the building Badger robots could clear obstacles, rubble, open doors, clear booby traps and make other preparations for manned entry. In standard con gurations the Badger is capable of pushing, digging and carrying moderate loads. They can be coordinated to cooperate to move even larger amounts. Badgers could place explosive charges to blow obstacles too large to be moved normally, blow holes in walls or oors, open barricaded doors etc. Badgers could also deploy unattended sensors, chemical and biological weapons detectors, smoke and tear gas. While these robots may not yet engage the enemy, they could disrupt the enemy's operations and compromise their hidden locations." + ] + }, + { + "image_filename": "designv11_65_0001776_0094-114x(89)90016-5-Figure3-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001776_0094-114x(89)90016-5-Figure3-1.png", + "caption": "Fig. 3. The seven-bar linkage.", + "texts": [ + "max~)l \"= ~ /C --lrlmin (24) G a r e min (~re, min ~ ] m a \u00d7 ' = ff)l - - ~[ ]r ( 2 5 ) Then C - Imm Ir G2 (26) re, max C --/max L - ~ . (27) a re, mill Substituting (26) into (27), yields 1 C - 2 : (G;,. . . . . l ' , . .x- : ' G . . . . . . - - G . . . . . in ' G . . . . in /rain)\" (28) Constant C can be calculated by eqn. (28) after selecting appropriate G . . . . . . and Gr~, m~o- This is easier. 5. A N U M E R I C A L E X A M P L E To take a seven-bar linkage for a numerical example. The linkage is shown in Fig. 3 and its dimensions and parameters are listed in Table 1. The linkage added a dyad is shown in Fig. 4, The added dyad group is HIJ and the roller is IJ. The curves of the input torques of the linkage, running at 80 r.p.m., are shown in Fig. 5. Curve 1 and curve 2 are the input torque before and after balancing respectively. The parameters of the dyad and the values of the input torques are listed in Table 2. The fluctuation value of the input torque is reduced by more than 80% and its maximum value is decreased by more than 54% after balancing" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003434_0020-7403(77)90048-0-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003434_0020-7403(77)90048-0-Figure1-1.png", + "caption": "FIG. 1. Design of a dumb-bell shaped projectile.", + "texts": [ + " Retardations during plough and rotations at exit were measured for each projectile and curves of plough development with time and angular velocity, with intersphere distances were obtained. Crater profiles and volumes were also measured for d.s. projectiles and are compared with those of two spherical projectiles. 555 556 G . H . DANESHI and W. JOHNSON Spherical and dumb-bell shaped projectiles (d.s.)t were fired into a fine grain sand (150-300t~ or 50-100 Tyler mesh) at 5 \u00b0, 10 \u00b0 and 15 \u00b0 angles of entry and caused to ricochet. The dumb-bells consisted of two spherical balls 1 in. in diameter connected by a light weight straight rod (see Fig. 1). The projectile velocity vector and the direction of the connecting rod were initially the same. The specification of each projectile and the speed range in which it was tested is given in Table 1. A specially adapted cartridge-gun was used for firing the projectiles. A spirit-level angle selector and a pair of braking wires, 5 cm apart, activating an electronic clock, were used to measure the impact angle and entry speed. Tests were carried out in a darkened room. Six spark flash condenser units were discharged at pre-determined intervals to illuminate the target area" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003567_bf00012615-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003567_bf00012615-Figure1-1.png", + "caption": "Figure 1. Cracked linear viscoelastic plate.", + "texts": [ + " To overcome this difficulty in elastic problems, the singular element in which the displacement function is taken from the analytical crack-tip expansions has been employed in the finite element computation for a number of problems involving transient stresses of a crack [1-4]. In the present paper, the above method is extended to the dynamic problem of linear viscoelastic solids and the time variation of the stress intensity factor is calculated for cracked plates. 2. Crack-tip stress and displacement fields Consider a plate of homogeneous, isotropic and linear viscoelastic material. The plate contains a stationary traction-free crack along the xl-axis as shown in Fig. 1. In order to obtain the near crack-tip field let us examine the nature of the viscoelastic boundary value problems. Applying the Laplace transform in time fo ~ f ( s ) = f ( t ) e -st d t (1) to the equation of motion, to the strain-displacement relations and to the constitutive relations leads to the following statements in t ransformed space. 0376-9429/801020097-13500.2010 Int. Journ. o[ Fracture, 16 (1980) 97-109 98 S. A o k i et al. ff~ij,i = s2pui (2a) gij = x,(ai,j + ~j,i) (2b) g~i = 2st2eo (2c) 6\"kk = 3sK~-kk (2d) where indices i and j indicate the Cartesian components of the stress o-, strain e and displacement u, and p is the mass density", + " The set of transformed equations (1)-(4) has a form identical with that of linear elasticity if s ~ ( s ) and s K ( s ) are associated with the elastic moduli/~ and K. It follows that the Laplace transform of the viscoelastic solution is obtained directly from the transform of elastic solution by replacing p~ and K with s ~ ( s ) and s/((s) , respectively. Thus we can obtain the stress and displacement fields near the crack-tip in viscoelastic solids as will follow. For the transient elastic problems as well as static ones, components of stress and displacement in the neighborhood of a stationary crack are given as, in a polar coordinate system as shown in Fig. 1, crij(t) = K l(t) f~j(O) + K n ( t ) f]](O) + Kin ( t ) eni ta ~ (5a) V2~rr ~ ~/27r-----~'it ,~, Int. Journ. o f Frac ture , 16 (1980) 97-109 Dynamic analysis of cracked linear viscoelastic solids 99 Kt(t) ./ Kiii(t) r I K I I ( t ) ~/~_~g[l(o ' is) ~/~g]li(0)r (5b) v2-~ gi(O, v) + + ui(t) = ~ ~ 2tz where Kt(t), Kii(t) and Kin(t) are the dynamic stress intensity factors for mode I, mode II and mode III, respectively, and v is the Poisson's ratio. Performing the Laplace transform we obtain KI(S) , /~n(S) " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000015_joaiee.1922.6594381-Figure2-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000015_joaiee.1922.6594381-Figure2-1.png", + "caption": "Fig. 2 indicates diagrammatically a two-pole rotor tapped at A B C for three-phase slip rings; the line 0 F represents the stationary polar center line. Assume the rotor to be driven in a counter-clockwise direction", + "texts": [], + "surrounding_texts": [ + "52 F I E L D : P O L Y P H A S E C O M M U T A T O R M A C H I N E S Journal A . I. E . E .\nconductors, the field strength, and the relative speed of field and conductors, or t o / i / p .\n7. The generated voltage of the second frequency, measured between consecutive brushes on the commu tator, corresponds similarly to the number of conduc tors, the field strength and the same velocity.\n8. Thus if we have the armature tapped for six phase slip-rings, and brushes on the commutator for six phases also, the generated voltage of frequency fl9 between slip-rings connected to consecutive taps, will be the same as that of frequency f2 between con secutive brushes; the currents also will be equal, barring magnetizing current and a small component which provides part of the power lost in the conversion.\n9. The commutator frequency f2 may be either higher or lower than the slip-ring frequency / i . But if one of the two frequencies is to be a very low one, it is advisable to adopt the lower for the commutator, and this becomes essential if the lower frequency is to be subject to variation passing through zero, that is, changing its direction of phase rotation. In deciding which frequency to assign to the commutator, it will be noted that:\na. The mechanical speed is the same in either case. b. The core loss is not greatly affected by the choice, since the field rotates at a speed corresponding to the one frequency, relatively to the stator, and the other frequency relatively to the rotor.\nc. The relative speed of rotor and field should be high, for minimum size of machine; i. e., the slip-ring frequency should be high.\nd. With reference to commutation, the transformer action in the coils connected to the segments under the brush will be the smaller the lower is the frequency on the commutator side.\ne. The frequency can be brought down to zero at the commutator, but not at the slip-rings.\nOn the whole it is usual to assign the higher fre quency to the slip-rings, but no assumption on this point will be made, and the results will apply equally either way. As regards the two possible rotor speeds, we shall find that by adopting a suitable convention with regard to the sine of , the angle of displacement of current on the commutator side, the results obtained apply equally for either speed, i. e., for f2 positive or negative.\nNow the currents in the rotor will result from the combination of the slip-ring currents at frequency flf and the commutator currents at frequency / 2 . Generally, in the case of superposed currents of two frequencies, the determination of the mean PR loss averaged over a few cycles, is simple, as each set of currents involves its own loss, and the net resultant loss is merely the sum of the two separate ones, irre spective of their relative magnitude or phase. In the present case, these simple relations do not hold, for although in any one conductor we have merely current of the first frequency introduced via the slip-\nrings, and for brief periods, we similarly have the actual current of the second frequency superposed in the same conductor, yet before one cycle is complete, the rela tive movement of commutator and brush has trans ferred this conductor into another phase group of the / 2 system.\nIt becomes necessary to view the cycle of events for atypical conductor, and then by a system of averages\u2014 or simple integrations\u2014 to arrive at the resultant loss.\nWe may start from the basis that should power be mechanically transmitted to or from the rotor, enabling the one system to operate alone, then the distribution among the arms of the delta-connected rotor of the Y-currents introduced at the slip-ring taps, or at the equivalent taps corresponding to the instantaneous positions of the brushes, will necessarily follow the ordinary Y\u2014A courses; and by the symmetry of the system, the voltage upon the idle taps or brushes, as the case may be, will form a symmetrical three-phase or six-phase system. As this applies to the case of operation with power being supplied either via the commutator or via the slip-rings, the result of super posing the two will leave balanced conditions and we may therefore consider that each set of currents upon entering the rotor divides up among the delta arms in the orthodox way, so that at any instant the actual current in any conductor is the sum of the two corresponding instantaneous currents. The same rea soning may, of course, be applied to other numbers of phases than three or six on either or both sides.\nWe now require a simple way of viewing the life history of a typical conductor, determining for any one instantaneous value of the fx current the range of values of the / 2 current. We can most readily do this by reference to the only feature which is common to the two frequencies, viz.: the armature reaction. To render this process more clear, we shall on paper bring the air-gap field to rest in space by rotating the brushes on the commutator, and by readjusting the rotor speed suitably.\nAssume the field to be rotating clockwise, viewed from a given side of the machine. Superpose upon the whole machine a counter clockwise rotation of f2/p revolutions per second, thus bringing the field to rest in space while the brushes rotate counter clockwise at this speed. Note that / 2 may be negative.\nWe can now readily picture the current distribution in the rotor, since the armature reaction of the / x currents, and again that of the f2 currents, will be represented each by a diagram fixed in space (although slightly pulsating or varying). Although produced by currents of different frequencies, these two diagrams will have the same number of poles, and on the average will cancel one another, the one being, at every instant, nearly superposable upon the reverse of the other\u2014 if for the moment we ignore the magnetizing current.\nWe have two methods open to us of investigating the I2 R loss; we may do so (1) by considering the loss", + "Jan. 1922 F I E L D : P O L Y P H A S E C O M M U T A T O R M A C H I N E S 53\noccurring, from moment to moment, in the succession of conductors which occupy a point fixed in space, upon the rotor periphery; then averaging this up with respect to time, and again with respect to angular position.\nOr (2) we may consider a marked conductor upon the rotor; average the loss occurring in it; and again average this up for all positions on the rotor. This latter method enables us to discriminate between conductors near to, and far from the slip-ring taps, and thus has an advantage. We shall adopt this procedure, and we shall first omit the magnetizing current, in order to make the procedure more clear.\nWith this omission, we have the same power factor on the two sides of the machine. That this must be so, can be seen by considerations along the following lines: The same air-gap field is responsible for causing the generated voltage of each frequency, and it is stationary in space. The armature reaction diagram of the currents of frequency fi must be generally similar to that of the currents of frequency / 2 , except for sign since the sum of the two produces the working field and we are neglecting the magnetizing current fo^ the present. But the armature reaction of the watt component of current of either frequency has its polar center lines at right angles to those of the air gap field. Therefore, these two component armature reactions, for the two frequencies, have their polar center lines coin cident; the watt component of current of frequency / i must be equal to the watt component of current of frequency / 2 , in order to effect balance of mechanical torque; and therefore the armature reactions of these two watt components balance one another; hence the\nor lagging, is the same; while for negative values of f2 the power factor is the same on the two sides, but a lagging current on the slip ring side is associated with a leading one on the commutator side, and vice versa.\nRotor Velocity Air-gap field is stationary\nF I G . 2\narmature reactions of the two remaining components must balance one another, i. e., the two wattless cur rent components must be equal and therefore the power factor must be the same on the two sides.\nWe shall now show, still neglecting the magnetizing currents, that for positive values of / 2 not only is the power factor the same on the two sides, but the sine of 0, which \u0302 determine whether the current is leading\nmechanically and to be acting as an a-c. generator, then the current in every conductor of the A B phase band will reach its maximum when the center line 0 L of the phase band has some definite angular position, say in advance of 0 F, where \\p may be positive or negative. If the load be noninductive \\p will be zero and the current will be a maximum in all conductors of the phase band at the moment when 0 L coincides with 0 F. If the load be inductive and the current be lag ging by an angle a, then when the current reaches its maximum in the conductors of the AB phase band, the center line 0 L will have advanced beyond OF by the angle a, and in the well recognized manner, the con ductors of the phase band will occupy such a position that the reaction of the currents in them has a com ponent along 0 F opposing the main field.\nIf now the points ABC, instead of being fixed tap points rotating with the rotor, should correspond to the momentary positions of brushes on the commutator which are rotating in a direction the reverse of that of the rotor, then the movement of the center line 0 L by the angle a will be in a clockwise direction in the figure. But as the voltage induced in the conductor is due to the motion of the conductor, and not that of 0 L, it will remain in the same direction as before, and hence also the current in the conductor if the machine is still acting as a generator via ABC. In this case, therefore, a lagging current will produce a reaction having a component along OF, tending to increase the field strength although the machine is acting as a generator arid not a motor.\nHence we, have the condition already outlined for", + "54 F I E L D : P O L Y P H A S E C O M M U T A T O R M A C H I N E S Journal A . I. E. E .\nthe relative phase displacement on the two sides, viz., that when f2 is positive i \u2014 2, but when f2 is negative 0i = \u2014 fa. We have seen that the power factor is the same on each side of the machine; it follows that the amperes per slip ring equal the amperes per brush stud, assum ing the same number of phases each side, for the volt age between rings equals that between brush studs.\nFig. 3 shows a two pole rotor with chorded windings; the right hand diagram shows six phase taps brought out, and the left hand three phase taps. The region occupied by conductors of one phase, which we may call a phase band, is shown by heavy lines, dotted for the inner layer conductors and full for the outer. The tap is shown as though located opposite the slot con taining the outer layer conductor to which it is con nected, but the inner layer conductor to which it is also connected is indicated by a similar letter with a circle around it; this is displaced from the outer layer conductor by one coil span. 0 L' represents the center line of the outer layer phase band, and 0 L\" the center line of the inner. The effective center line for deter mining the phase relations is OL which is displaced from OL' by half the angle of chording. It will be understood that the tap points shown may be either actual fixed taps connected to slip rings, or may repre sent the momentary position of brushes on the com mutator; in either case the voltage generated in the phase band is the same, being produced by the move ment of the conductors in the field and not by the movement of the center line of the phase band. If we are dealing with a commutator phase band the center line 0 L may be rotating oppositely to the rotation of the rotor itself; viz., when f2 has a negative value.\nIn the figure, 0 F represents the polar center line of the main field and ^ the angle, measured counter clockwise, of 0 L from 0 F. We take the component of current in the phase band which has frequency / , to have a root-mean-square value of unity and we define its phase angle as follows:\nIf / is positive, + c/> represents the angle by which the current leads with respect to conditions for unity power factor.\nIf / is negative, + a> represents the angle by which the current lags with respect to conditions for unity power factor.\nIn either case the power factor is cos and a mag netizing armature reaction is associated with generator conditions when fa is positive.\nIt will be found that with this definition the current component of frequency / in every conductor of the phase band A B is represented by\n\u00b1 V 2 cos + ) (1) for either direction of rotation of 0 L. Note that \\p is always to be measured in a counter-clockwise direction.\nWhen we are dealing with a specific conductor P in say the outer layer, it becomes convenient to specify\nthis by its co-ordinate measured from the center line of its phase band in the layer in question, or O L ' for the outer layer, instead of from 0 L, and consequently it becomes convenient to deal with the angular co ordinate of OL', rather than 0L, with respect to the fixed field axes. It will be noted that we may still use expression (1), and define \\p therein as the angle measured counter clockwise from some suitable zero line 0 F' fixed in space, to the phase band center line OL'; further the datum line OF' will be the same whether 0 L' is a center line of a slip ring phase band and therefore is fixed in the rotor, or whether it repre sents the momentary position of the center line of a commutator phase band and is rotating in the same direction as the rotor or the reverse. In other words, the chording angle does not enter into consideration in dealing with the difference of phase of the currents of the two frequencies in a conductor, either for positive or negative values of f2.\nB 2\n- =^-Hotor Velocity =3^- Brushes, assuming f2 Positive\nF I G . 4\nTHREE-PHASE CASE\nConsider now the case of three-phase slip ring taps and commutator brushes 120 deg. apart, Fig. 4, and omit magnetizing current for the present. P in the figure represents a definite conductor of the rotor; it is shown as occurring in phase band AxBi of the slip ring side, and in A2B2 on the commutator side. It remains permanently in the AiBx phase band, but passes successively through the various phase bands of the second frequency (commutator).\nCurrents then arise in the conductor P having values which have already been discussed and represented by (1). They have respectively the values: V 2 cos ( ^ i + ) of frequency / i and V 2 cos (ir + fa + fa of frequency f2 The difference (\\p2 \u2014 may clearly have any value between 7r/3 + 6 and - 7r/3 + 6. Further, every value within this range is an equally probable one and will occur for the same number of seconds per hour of operation, on the average. We therefore require the average value of 2 [cos yp - cos (^ + 7) ]2 (2) for all values of y within the range (6 d= 7r/3), giving equal weight to each value. We shall next require" + ] + }, + { + "image_filename": "designv11_65_0001136_s0020-7403(98)00106-4-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001136_s0020-7403(98)00106-4-Figure1-1.png", + "caption": "Fig. 1. Typical semi-trailer with lashings.", + "texts": [ + "15 m) ZS(i) de#ections of the spine relative to the deck ZS;S unstrained height of the suspension g gravitational constant a amplitude of the roll angle of the ship (303) b amplitude of the pitch angle of the ship (83) h(i) roll angle of the mass m(i), relative to the deck / pitch angle of the trailer, relative to the deck t yaw angle of the trailer, relative to the deck j damping constant of the suspension (15kNs/m) k s coe$cient of friction between the deck and the suspension (0.6) A typical trailer and lashing arrangement is shown in Fig. 1. It shows the characteristically high centre of mass and the low level at which the lashings are attached to the trailer. Previous work by the authors has investigated the dynamic behaviour of the much more rigid &&box van'' trailer [1, 2] and the cases of a torsionally #exible &&#at-bed'' trailer with a single [1, 3] and twin concentrated loads [4]. Also, the case of a &&live'' load (i.e. suspended carcasses and tankers with a free liquid surface) has been considered and is described in Ref. [5]", + " A section of the chassis under the suspension mass is assumed rigid due to the sti!ening e!ects of the suspension members and this e!ectively increases the width of the spar at this point. The spars, which protrude from the spine and which carry the lashings, are assumed rigid and they rotate independently about the centre of the spine. The lashings are numbered from the \"fth wheel (i.e. the trestle end) to the back of the trailer in ascending order and the lashings are identi\"ed by their number and their position to the left or right of the spine, as shown in Fig. 1. The masses move with the vertical and horizontal de#ections of the spine but they are assumed not to take up the local slope of the spine. This is based on the assumption that the load is a distributed one and that it does not contribute to the bending or torsional sti!ness of the trailer, thus it will not rotate locally about the oy and oz axes. Because the trailer is assumed to be \"xed in length, there exists a simple relationship between horizontal positions of the centres of the masses along the ox-axis" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000095_0957-4158(94)e0025-l-Figure6-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000095_0957-4158(94)e0025-l-Figure6-1.png", + "caption": "Fig. 6. Linking with a single chain stitch.", + "texts": [ + " The highest quality method of joining these components together is a process termed \"linking\", in which the knitted loops of one component are matched one-for-one with those of the component with which it is to be joined and a chain stitch is used to sew them together. This provides a flexible joint, with minimum bulkiness in the seam, although care must be taken that no loops are missed in the process or the garment can unravel. The process is therefore used for the more expensive products and is performed by hand with the aid of simple machines. These comprise a series of grooved points onto which the loops are loaded and which serve to guide the needle of the machine's sewing head, as illustrated in Fig. 6. The operators of these machines need skill, concentration and good eye-sight. A mechatronic system has been developed to help to automate the linking process 104 T.G. KING [10-14]. Collars and other ribbed garment components are knitted in bulk; each joined to the next by waste courses of knitting which enable them to be separated before linking. The row of loops on each trim to be used for linking is termed the \"slack course\" since it is knitted with slightly larger loops than the other rows. The most arduous part of the manual linking process is identifying these loops and loading them onto the points" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003918_j.jmatprotec.2005.02.163-Figure17-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003918_j.jmatprotec.2005.02.163-Figure17-1.png", + "caption": "Fig. 17. Swivel bar concept with replacing element.", + "texts": [ + " As soon as the plug reaches its final position, the spring presses the bar into of the inlet, the plug returns to the initial position and the plug is fixed by the swivel bar. The reference plug and two swivel bars to fix it are shown in Fig. 16. If the plug reaches the slanted surface, the swivel bars would move around their rotation axis which is symbolised by the drilling. As mentioned earlier, the plug has to be underneath the surface and this could cause a problem for removing the plug. The simplest solution would be to produce two openings, which allows the operator to grab it. Another way to deal with this problem is to attach a replacement part as shown in Fig. 17. By operating the opening and a little further movement, the slanted surface causes the plug to be lifted leading to better accessibility. llows the usage of the same element for all different kinds of lugs; only the position of the rotation axis has to be modified o the certain plug in order to make sure that the right part of he swivel bar is entering the inlet. The second realisation to achieve the normally closed etup is shown in Fig. 16. Similar to the old system, the plug Fig. 14. Reference plug" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002866_095440603322769938-Figure10-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002866_095440603322769938-Figure10-1.png", + "caption": "Fig. 10 (a) The Von Mises distribution for one tooth model of (a) a traditional spur gear and (b) the proposed gear ring-involute tooth", + "texts": [], + "surrounding_texts": [ + "To manufacture the geometric surface of the proposed gear mechanism, this study considers the geometric parameters as shown in Table 1. In actual cases selection of the gear ratio will depend on the application of the gear mechanism. The gear ratio 1.5 was adopted. In the cutting simulation process, the gear and the pinion are assumed to be made of steel. This type of steel is described in section 6. The contour of the gear and the pinion with ring-involute teeth are shown in F ig. 6. According to the analytical equation of the mathematical model for the proposed gear (convex) and pinion (concave), a cutting simulation for manufacturing the gear and the pinion were performed by a computer aided manufacturing program. In section 3, the mathematical models of the gear and pinion were obtained. Therefore, the cutting simulation process is important for the manufacturing of the gear and the pinion with ring-involute teeth. By using the results obtained earlier and using a computer aided manufacturing program, the cutting processes for machining the gear and the pinion with ring-involute teeth are shown in Figs 11 and 12 respectively. Here, only one ringinvolute tooth for the gear and pinion is simulated. The other teeth of the gear and pinion can be manufactured by a computer numerically controlled (CNC) machine and an index head. The index head will be repeated for N 1 times with a rotation of 2p=N 1 when the gear (convex) is manufactured. Similarly, the index head will be repeated for N 2 times with a rotation of 2p=N 2 when the pinion (concave) is manufactured." + ] + }, + { + "image_filename": "designv11_65_0001863_s0022-0728(84)80344-7-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001863_s0022-0728(84)80344-7-Figure1-1.png", + "caption": "Fig. 1. Dependence of the normal ized PP current response of the E ~ E mechan i sm on the degree of a t t rac t ion between adsorbed ions of the in termediate . F r u m k i n ' s i so therm: flA~,_ o = 0 e x p ( a 0 ) / ( 1 - 0); a = 2 g F J R T . 4~ = i (qr tp) l /2(FqD1/2A*)- l ; O = 1 0 . 5 c m 2 / s ; E l \u00b0 = E2\u00b0 ; tp = 50 ms; A * / F s = 1 . 3 X 1 0 3 c m - a ; f lF s = 1.0 cm; a = 0.0 (1), - 0 . 2 5 (2), - 0 . 5 (3), - 1 . 5 (4), - 2 . 2 5 (5), - 3 . 0 (6) and - 4 . 0 (7).", + "texts": [ + " 36 According to the literature data [11,37,38], the dc polarographic wave of a reversible redox couple could be influenced by the adsorption of the reactant and the product, if it follows Frumkin's isotherm, in two ways: the slope of the wave will decrease if the repulsion prevails and, conversely, the slope will increase if the attraction forces between ions in the adsorbed state are larger than the repulsive ones [11,37,38a]. A very high intensity of attraction may cause the sudden jump of the faradaic current from its minimal to its maximal value [37]. Also, the half-width of the linear scan voltammetric peak will decrease in the case of attraction between adsorbed ions [38b]. The repulsion would cause an increase in the half-width of this peak [38b1. Figure 1 shows the theoretically predicted influence of the variation of Frumkin's parameter \" a \" on the faradaic pulse polarographic current of the E ~ E mechanism in the case of attraction between adsorbed ions. The values of the parameters are given in the figure caption. If a = 0, a well-developed pre-wave and post-wave can be seen. Increase of the absolute value of parameter \" a \" causes the change of the pre-wave to the \"pre-peak\" and the post-wave to the \"post-hole\". The inadequacy of the terms \"pre-wave\" and \"post-wave\" has already been discussed [24], but we will use these terms throughout the present paper in order to be better understood. The changes from Fig. 1 are the following: the rising portion of the pre-wave 37 becomes steeper, the maximum of the pre-wave shifts in the positive direction, and the minimum between the pre-wave and the main wave becomes deeper. The postwave is the symmetrical picture of the pre-wave. The physical reason for these changes is the gradual change of the adsorption constant fl* which is proportional to the degree of coverage 0. When parameter \" a \" has a certain negative value, the coverage of the electrode surface, 0, increases faster and at more positive potentials than if a = 0", + " It has been mentioned earlier in the text that for of the variation of Frumkin's parameter \" a \" is negligible. This is logical because 0 is 2\u00a3 8 1.5 1.0 0.5 ,E-E;'dV I I o 0.2 o.1 o!o -oi, -o.2 Fig. 3. Dependence of the normalized PP current response of the E ~ E mechanism on the normalized concentration of the reactant. Frumkin's isotherm. Attraction between adsorbed ions of the intermediate: a = - 2 ; ( A * / F s ) / c m -1 =102 (1), 103 (2), 1.1 x l 0 3 (3), 1.2\u00d7103 (4), 1.3\u00d7103 (5), 1.5 x l 0 3 (6), 2\u00d7103 (7) and >/1 \u00d7 104 (8). All other parameters as in Fig. 1. 39 always very low. Also, if A * / F s >~ 10 4 cm -a, when only a main wave appears, the influence of a is not notable. The reason for this is the very high ratio between the amplitudes of the main wave and the pre-wave. Besides, for such a high reactant concentration and for an electrolysis time of 50 ms, the total coverage of the electrode surface is obtained at the very beginning of the pulse, so that the reduction occurs mainly on the covered electrode surface. Now the results of calculations of the possible influence of Frumkin 's parameter \" a \" on faradaic differential normal pulse polarographic (DNPP) current responses of the E , E mechanism will be described", + " The influence of the attraction between the adsorbed ions can be seen in Fig. 4. As in pulse polarography, the separation between these three peaks and the significant increase of both the pre-peak and post-peak can be seen. The main peak also increases, but not as much as the other two. From the formal point of view, we can treat NDPP responses as being similar to the first derivative of the PP response and then the changes of current which can be observed in Fig. 4 can be understood if they are compared with the changes of current from Fig. 1. The pre-peak corresponds to the rising portion of the pre-wave, the main peak corresponds to the main wave, and the post-peak corresponds to the post-wave. The half-widths of the pre-peak and post-peak decrease if parameter \" a \" becomes more negative, which is in agreement with literature data [38b]. The increase of the amplitude of the main peak reflects the increase of the slope of the main wave in PP, which is also in agreement with literature data [11,37,38a]. The needle-shaped peaks near the summit of the post-peak, which appear for a < 4, are probably the consequence of the instability of the mathematical model, rather than representative of a certain physical process. Note that in PP (Fig. 1) there is no corresponding sudden jump of the rising portion of the post-wave (curve 7, a = - 4). D N P P is probably more sensitive to instabilities than PP. In Fig. 5 one can see the influence of the repulsion between the ions from the adsorbed layer on the DNPP current response of the E~ E mechanism. All the parameters are equal to those in Fig. 4. Very simple changes can be observed: the pre-peak decreases, the main peak remains virtually unchanged and the post-peak vanishes. The half-widths of the pre-peak and the main peak increases, which is the reflection of the decrease of the slopes of the pre-wave and main wave (see Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002765_s0045-7949(02)00029-9-Figure3-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002765_s0045-7949(02)00029-9-Figure3-1.png", + "caption": "Fig. 3. Wheel\u2013ground contact geometry.", + "texts": [ + " The excavator bucket trajectory is C q; t\u00f0 \u00de \u00bc Rxi a cos p 4 t \u00fe b Ryi c1 hi c2 2 4 3 5 \u00bc 0; \u00f04:12\u00de where a is the bucket reach in the x-direction, and b is the half way point of reach for the bucket. The arbitrary constants c1 and c2 specify the height and angular orientation of the bucket. The wheel loader bucket trajectory is similarly defined as follows: C q; t\u00f0 \u00de \u00bc Rxi c1 Ryi \u00fe a cos p 4 t \u00fe b hi c2 2 4 3 5 \u00bc 0: \u00f04:13\u00de The contact force model used for the wheel loader system defines the contact dynamics of the wheel-axis sets with the ground as shown in Fig. 3. It is similar to that used for the caterpillar trackground model studied in [6], and employs a springdamper element between points pi and pj, the former residing on the most penetrated part of the wheel and the latter on the ground. The magnitude of this force, see [7] for further details, is Fcij \u00bc kdij \u00fe cdij _dij; \u00f04:14\u00de where k and c are the spring and damping coefficients respectively and the amount of penetration is given by dij \u00bc Ry;i Ry;j r: \u00f04:15\u00de The virtual work due to a force with magnitude Fcij is dW \u00bc Fcij rpij rpij orpij oqi orpij oqj dqi dqj ; \u00f04:16a\u00de dW \u00bc QT i dqi \u00feQT j dqj; \u00f04:16b\u00de where Qi \u00bc QR;i Qh;i \u00bc Fcij I uTpiA T h;i rpij rpij \u00f04:16c\u00de and Qj \u00bc QR;j Qh;j \u00bc Fcij I uTpjA T h;j rpij rpij : \u00f04:16d\u00de The position vector written in the global coordinate system is rpij \u00bc rpi rpj \u00bc Ri \u00fe Ai upi Rj Aj upj ; \u00f04:17\u00de where upi \u00bc r cos 3p 2 hi r sin 3p 2 hi \" # and upj \u00bc Ri;x Rj;x 0 : \u00f04:18\u00de The generalized force vectors Qi and Qj are included in the external force vector Qe, and one may then complete the construction of Eq" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002338_cca.1993.348311-Figure2-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002338_cca.1993.348311-Figure2-1.png", + "caption": "Figure 2. Schematic of Angular Relationship.", + "texts": [ + " INTEGRATED GUIDANCE LAW/AUTOPILOT DESIGN As indicated earlier, an avenue to improve the homing missile performance is to consider an integrated design of the guidance law and autopilot instead of separate designs. In this manner, the kinematics of the engagement geometry used in the guidance law development and the dynamics of the airframe as reflected in the linearized equations of motion and used in the autopilot design can be brought together. Consider the geometry of the engagement in Figure 1 and the relationship between the flight path angle and the pitch angle and angle of attack as shown in Figure 2. One way to formulate the integrated approach is to relate the guidance law to the dynamics of the airframe directly and solve for the commanded control surface deflection. For this purpose, let us assume a conventional proportional navigation guidance law given by where k - 3e0r, and t, - t/tr This transverse acceleration can be approximated to be perpendicular to the flight path of the missile so that (75) V,? - k( l - tJ where V, = missile velocity and y = flight path angle. The flight path angle is related to the pitch angle and angle of attack as (76) (77) where q = pitch rate" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001555_0301-679x(83)90057-9-Figure3-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001555_0301-679x(83)90057-9-Figure3-1.png", + "caption": "Fig 3 Close up photograph o f the hub piece on the gear\" box side o f the first set o f failed gear couplings", + "texts": [ + " Inspection of the failed gear couplings Photographs of the spool and hub pieces on the turbine side of the second set of failed gear coupling parts are TRIBOLOGY international 0301 679X/83/03014! 06 $03.00 \u00a9 1983 Butterworth & Co (Publishers) Ltd 141 Chander ~nd B/swas -- Abnormal weer of gear couplings shown in Figs ! and 2. The coupl ing had 52 teeth , -which were no~ crowned, wi th a p i tch circle d iameter o f 130 m m The wear o f teeth of the hub and spool pleces on the turbine side was iess than that on the gearbox side (Fig 3). Figs 4 and 5 show more details o f one t o o t h cu~ from_ the hub piece, Here, wear wi th some fretti~:g can be observed In atl the coupling pieces, the t ee th were covered wi th a lacquer-Eke substance. This layer, adhering to the t ee th could no t be removed by scraping wi th a fingernail. I n v e s t i ~ a t [ o r ~ Based on the case his tory and_ observations of the failed gear coup;ing pieces, the fol lowing course af invest igation w~s under taken ~:o analyse the o reb!em in its ent i re ty: s Metallurg~cal analysis of th~ c\u00a2,:~pnng \"~o identif j n-aer:~ lurgica] ~eficiences if' any, m :no m a t e r i \u00a3 o f t;~e coup l ing <~ Vib raucn spec : rum s" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000153_1.2802446-Figure3-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000153_1.2802446-Figure3-1.png", + "caption": "Fig. 3 Top view of the Initial conditions at the ball-rim encounter (B = I). The components in the A frame of the ball inertlal angular velocity are determined by the values of the impact parameter S and ball speed v.", + "texts": [ + "org/ on 01/23/2016 Terms of Use: http://www.asme.org/about-asme/terms-of-use a right-handed orthogonal reference frame A lie in the plane determined by O, S*, and B, with ai perpendicular to the line through B * and B and directed toward the outside of the hole when the tilt into the hole, , is zero. The unit vector a2 points from B to B* and is vertical when is zero. The roll-around angle 9 is the angle between a fixed inertial line (the inertial X axis perpendicular to the initial velocity vector) and the line OB (see also Fig. 3). The assumption of constant ball-rim contact reduces the num ber of generalized coordinates required to specify the system configuration from six to five; 0 and (which determine the position of the ball center) and three orientations of the ball. For example, a 1, 2, 3 Euler angle set might be chosen to characterize ball orientation. In this case the derivatives of the three Euler angles would be linearly related to the three compo nents of the angular velocity of the ball through a transformation matrix which depends on the Euler angles themselves (Kane and Levinson, 1985)", + "org/about-asme/terms-of-use m Note that the only angle involved in Eqs. (8) - (10) is . These equations were also independently verified using a NewtonEuler approach and eliminating the contact tangent plane com ponents of the contact force. The motion of the ball around the rim {0 and 4>) can be calculated by solving only Eqs. ( 8 ) - (10) and (5) and (6), providing further justification for ne glecting the ball orientation coordinates in the formulation of the problem. The model was evaluated by solving the differential equations numerically. Figure 3 gives a top view of the ball-hole encounter and shows the ball rolling toward the hole with its velocity vector V offset an impact distance 6. Note that, unlike Holmes (1991), here our datum for 6 is the lateral hole centerline so that the initial value for 6 is always negative. This corresponds to the ball rolling around the right side of the rim, as seen by the putter. Rolling around the left rim is not dealt with explicitly here. Just as the ball contacts the rim, the initial conditions are determined by 6 and v according to = 0 cos 8o UJ2\u201e = 0 u)i\u201e = - \u2014 sm f\u201e Rh (11) The function of the impact parameter 6 is to determine 9\u201e and thus to apportion the initial horizontally rolling angular velocity iOo = vIRh into uj\\ and LO^ components" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000046_cbo9780511628863.029-Figure3-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000046_cbo9780511628863.029-Figure3-1.png", + "caption": "Figure 3. (a) A wormhole in Euclidean space, (b) A wormhole with both mouths at rest in Minkowski spacetime, and with synchronous time connection.", + "texts": [ + " If you give the walls of Misner space vanishing relative velocity (3 = 0 and change them from flat planes to spheres, you will obtain the simplest example of a traversable Cambridge Books Online \u00a9 Cambridge University Press, 2010https://www.cambridge.org/core/terms. https://doi.org/10.1017/CB 9780511628863.029 Downloaded from https://www.cambridge.org/core. University of Warwick, on 18 Sep 2018 at 06:12:08, subject to the Cambridge Core terms of use, available at wormhole. More specifically: extract two balls of radius b from 3-dimensional Euclidean space, and identify their surfaces so when you enter the surface of the right ball, you find yourself emerging from the surface of the left ball; the result is the wormhole of Figure 3a. For concreteness and simplicity, make the identification via reflection of the balls in the plane that is half way between them, so for example the point P appears at the locations shown in the figure. In Minkowski spacetime, this wormhole is obtained by identifying the two world tubes swept out by the two balls, as in Figure 3b, with events at the same Lorentz time t identified (\"synchronous identification\"), for example the events P shown in the figure. There is spacetime curvature on the wormhole's coinciding spherical walls (which are called its mouths). A key feature of this curvature can be deduced by noticing that a bundle of light rays that enter the right mouth radially converging must emerge from the left mouth radially diverging (because of spherical symmetry). This means that the mouths act like a diverging lens with focal length equal to 26" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001426_robot.1995.525615-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001426_robot.1995.525615-Figure1-1.png", + "caption": "Figure 1 .l: Place on Table", + "texts": [ + " More recently, new robot programming approaches have promoted the use of artificial] user-defined constraints (programmed constraints), as an executable specification language for the desired behavior of a robot (e.g. [211 22, 26, 111). These constraint-based approaches are intermediate levell languages] promising a higher, more declarative level of programming than trajectory-based approaches, and being computationally more tractable than motion planning. Consider an example from [21] of programming a pick-and-place task (as depicted in Figure 1.1) of placing an object in the robot\u2019s hand on the table. In a conventional robot programming language, this must be expressed by arbitrarily commanding a particular motion to some point on the table. In the Least Constraint approach, the i,ask is programmed by placing constraints in the hand position domain. The hand is constrained to stay within the boundaries of the table by the constraints fi and fi. The constraint f3 moves down at constant speed v from height h, so that after a time of h / w , the end-effector is on the table top. Note that the exact location on the table top is unspecified since this is not important for the success of the task. An advantage of this approach is tlhat a program can be easily modified by adding new constraints. For example, if we now wish - 2373 - IEEE Internatlonal Conference on Robotlcs arid Automation 0-7803-1965-6/95 $4.0001995 IEEE our laboratory [l]. testbed and Figure 1.3 shows a schematic. Figure 1.2 shows the actual to place the object at a particular location on the table, this can be achieved by adding the cone-shaped constraint marked f4 in Figure 1.1. In this paper we consider the numerical solution of the system of differential equations and constraints arising from such robot programming methods. Specifically, we consider the numerical solution of initial-value problems of the form q = f ( t , q ) + B(t,q)u (l . la) 4 0 ) = 4 0 , ( l . l b ) subject to c(t,q) 2 0, (1.k) (+tax ,Ymax ) Y f where q E 8\" are the state variables, c E SRncnstr are ( X m n JYmin I inequality constraints to be satisfied by q for all t 2 0, and U E PcntrZ are the controls. Since in practice controls cannot take on arbitrary values, they must also satisfy a feasibility condition, (1.2) Figure 1.3: The mobile robot (not t o scale). The equations of motion of the robot may be writlujl < umaz,j for j = 1 , 2 , . . . , nent,.i. ten as The aim of our algorithm is to prescribe feasible values of U for the equations of motion (l.la), which allows (1.1~) to be satisfied. We note that despite being similar in appearance, inequalities (1.2) are treated differently than inequalities (1,lc). The variables U in (1.2) are 'independent' parameters of the problem to be determined by the algorithm, while the relations (1" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003916_1068009.1068246-Figure3-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003916_1068009.1068246-Figure3-1.png", + "caption": "Fig. 3 Geometry of estimating final time tfi", + "texts": [ + " 1 The force analysis of the missile X Y Z D L T v mg nonlinear trajectory optimization problems. As for the ith problem, we can take use of GA to search the optimal trajectory, and the singular perturbation technique is applied to approximate the final time ( 1, , )fit i N of each problem. The flight trajectory of the target is guaranteed by the ground support system through filtering and identification and transmitted to the missile during the midcourse guidance. A given target trajectory is divided into N elements which are represented as 1 ( 1, 2, , )i iE E i N (as shown in Fig. 3). As for the target position iE , we can compute the orientation angle i of the target and the slant range iR between the target and the missile through the Eqs.9 and 10. arctan( )ti mi i ti mi h h x x (9) 2 2( ) ( )i ti mi ti miR h h x x (10) According to i and iR , we can get the approximate expression of the final time from the missile position iP to the target position iE , where ti is the course angle of the target, tiv is the target velocity, the subscripts m and t denote missile and target respectively. The final time fit is derived from zero-order singular perturbation technique. 8 We assume the target flies at constant speed in the interval from iE to 1iE , ( 1, 2,..., )i N . So, there must be a shortest course for interception of the target, and the final time is the time used by the interceptor. The derivation process of fit is as follows. The symbols employed in this section are shown in Fig.3 By projecting iR and tiv to 1i iPE , we can obtain the expression of the final flight time, cos( ) cos( ) i i i fi mi ti i ti R t v v (11) In order to insure the validity of the homing guidance, we must conserve sufficient energy for terminal engagement of intelligent targets. In other words, the energy consumed in the midcourse phase should be minimal. We represent the consumed energy by the terms of the power of missile iE , ( cos( ) )i i i mi i i T D v E m g (12) where, iD is the drag force on the missile, which is represented in Eq" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000362_0379-6779(92)90286-r-Figure3-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000362_0379-6779(92)90286-r-Figure3-1.png", + "caption": "Fig. 3. Cyclic voltammograms of a film of poly[3-(3,6-dioxaheptyl) pyrrole] adsorbed on platinum electrode (0.07cm 2) (polymerisation in propylene carbonate, LiC1Q 0.1 M, monomer 0.1 M, consumed charge 7 mC); ref. Ag/AgNO~ (0.1 M MeCN); sweep rates from 10 to 100 mV s -1.", + "texts": [ + " Consequently, an original chemical pathway has been suggested to synthesize a subst i tuted pyrrole-bearing polyether chain at the 3-position by means of an ethyl spacer-arm [4]. The new pyrrole derivative obtained was polymerised (see Fig. 2 for the polymer molecular structure) under the electrochemical conditions used for unsubst i tuted pyrrole (NB., at the same potential), indicating no change in the aromatic nucleus activity. The oxidized thin films obtained are smooth, flexible, and exhibit good conductivity: 2 S cm -1. The quite good reversibility of its redox system in organic electrolyte appears as the major feature of this new organic material (Fig. 3). Such behaviour is rarely encountered in the conduct ing polymer family. From its formal redox potential and studies in other electrolytes, it has been concluded that this electroactive system is l inked to an anionic t ransport within the polymer matrix, like polypyrrole itself [4]. This enhanced reversibility can be explained by two concurr ing effects of the polyether substituents. The hysteresis (and sometimes multistep systems) current ly observed on the vol tammograms of conduct ing polymers which arise from a phase t ransi t ion between a rigid conduct ing insert ion compound and more flexible, ion-free neutra l chains, is nearly suppressed here by the flexible adjacent arms" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003052_ias.2000.882088-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003052_ias.2000.882088-Figure1-1.png", + "caption": "Fig. 1. Three-phase Switched Reluctance Machine.", + "texts": [ + " We also will see the influence of the boost capacitor value on the performance of the SRM as a motor and a generator. Finally, we will present the maximum output power obtained by choosing the appropriate turn-on angle for variable speeds, in motor and generator working modes. 11. SYSTEM DEFINITION In this section we first present the system components and then establish its model. The system is composed of a SRM and its converter. A. Structure of the studied Switched Reluctance Machine The SRM considered here is a three-phase doubly salient machine, with six stator poles and four rotor teeth (Fig. 1). 0-7803-6401-5/00/$10.00 Q 2000 IEEE We suppose that the SRM is not saturated; so each phase inductance (L) depends only on the rotor position (Fig. 2). The stator teeth are larger than the rotor ones, so that the inductance variation with rotor position (L( 8 )) can be approximated by a trapezoidal function, as shown on Fig. 2. So, the inductance is constant and maximal over one interval, constant and minimal over a second interval, and varies linearly elsewhere. The current waveform must be rectangular" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001776_0094-114x(89)90016-5-Figure4-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001776_0094-114x(89)90016-5-Figure4-1.png", + "caption": "Fig. 4. The linkage added a dyad.", + "texts": [ + "max~)l \"= ~ /C --lrlmin (24) G a r e min (~re, min ~ ] m a \u00d7 ' = ff)l - - ~[ ]r ( 2 5 ) Then C - Imm Ir G2 (26) re, max C --/max L - ~ . (27) a re, mill Substituting (26) into (27), yields 1 C - 2 : (G;,. . . . . l ' , . .x- : ' G . . . . . . - - G . . . . . in ' G . . . . in /rain)\" (28) Constant C can be calculated by eqn. (28) after selecting appropriate G . . . . . . and Gr~, m~o- This is easier. 5. A N U M E R I C A L E X A M P L E To take a seven-bar linkage for a numerical example. The linkage is shown in Fig. 3 and its dimensions and parameters are listed in Table 1. The linkage added a dyad is shown in Fig. 4, The added dyad group is HIJ and the roller is IJ. The curves of the input torques of the linkage, running at 80 r.p.m., are shown in Fig. 5. Curve 1 and curve 2 are the input torque before and after balancing respectively. The parameters of the dyad and the values of the input torques are listed in Table 2. The fluctuation value of the input torque is reduced by more than 80% and its maximum value is decreased by more than 54% after balancing. Input torque balancing of linkages In Table 2, the type is - 1 when the order of the connection of HIJ is clockwise, otherwise it is + 1" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003073_cp:20020168-Figure2-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003073_cp:20020168-Figure2-1.png", + "caption": "Figure 2. Relation of flux linkage vectors and torque angle 6.", + "texts": [ + " The rotor position is then given by i 6&), which, after differentiation and some filtering gives the speed signal. The proposed speed estimator is thus based on the following steps and inputs: 1. 2. 3. ,4. 5. the stator flux linkage As ; since the stator flux linkage is controlled within the DTC, its position is known, the actual developed torque from equation (6); the torque angle 6 can be calculated from the torque equation (7), [or it may be obtained from a look-up table representing equation (7)]. the rotor position; this is obtained from the stator flux position and the torque angle (figure 2), change of the rotor position io certain time interval gives the rotor speed; this must be filtered to get a smooth speed signal, the initial starting position of the rotor. It is obtained from a simple test at start [ll]. 506 IV. MODELING AND EXPERIMENTAL RESULTS AND DISCUSSIONS To validate the proposed sensorless speed estimator, modeling and experimental studies have been performed on 230 V, 3 phase, &pole, 1 kW IPM synchronous motor indicated in Table I. The sensorless DTC drive system is shown in figure 3" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002286_robot.2001.933085-Figure3-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002286_robot.2001.933085-Figure3-1.png", + "caption": "Figure 3: A 6R manipulator.", + "texts": [ + " Calculate matrices K,,, and K,. 5. Substitute matrices M, . Jc , K O , K,, and K, into the conservative congruence transformation (CCT) equation (14) to obtain the Cartesian stiffness matrix, K. 4.3.1 An example First of all, we must define local coordinates for the end-effector configuration including translation and rotation in order to find the Cartesian stiffness with respect to the twist basis. The configuration of the end-effector will be of interests at T = I. For a 6R manipulator shown in Figure 3, we can choose the local coordinates (51, ( 2 , &, [4, &,, 5 6 ) to parameterize the motion of the end-effector including the translational and rotational in mements. Thus , these six independent parameters are determined by the forward kinematics as follows (1 (2 = = LZ cos 81 cos 02 + L3 cos 01 cos(6'2 + 6%) - H sin 6'1 LZ sin 6'1 cos 8 2 + L3 sin 6'1 cos(& + 03) + H COS 81 <3 = \u20ac4 = 0 4 ( 5 = 05 (b = 06 (19) where L1, L2, La, and H are the link lengths shown in Figure 3. The Jacobian of the manipulator, J t , J c : (81, . . . , e 6 ) - f R 6 , i s g i v e n b y L1 + LZ sin 0 2 + L3 sin(& + 03) J11 XZ ,713 1 5 2 1 JZZ 523 0 0 0 J c = l o 0 5 3 2 0 533 0 0 1 0 0 ~ 0 0 0 0 0 0 1 0 where 511 = -Laslca - Lsslcas - .Hcl, 5 1 2 = - L 2 ~ 1 ~ 2 - L3c1s23, 513 = - L ~ c I s ~ ~ , 5 2 1 = L ~ C I C Z + L ~ c ~ c z ~ - Hsl, 5 2 2 = -L2Sisa - L3S1s23, 5 2 3 = -L3slsas, J32 = Lac2 + L3c23, and 533 = L3ca3. Note that the Jacobian matrix in equation (59) happens to bc the samc as the conventional Jacobian matrix dcfincd by Jo = $$j due to thc choice of local coordinatcs" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000759_004051759106101210-Figure8-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000759_004051759106101210-Figure8-1.png", + "caption": "FIGURE 8. Filling yam crimp.", + "texts": [ + " vertical distance between the two centers of the cross sections of filling yam 1 ); C = the back rail lift above the cloth line, 11 = the distance between the cloth fell and the harness, and /2 = the distance between the harness and the center of the back rail. The expression of dl&dquo;n is where The extension dli of the lower shed warp yams consists of the extension dl, due to shedding and extension dlib due to beat-up: The expression of dl, s ( 12 ] is The expression of dill, is Filling yarn crimp (Figure 8) may be approximately calculated using the following method: The force F acting on the filling yarn by the warp yarn is balanced by the tension S of the filling yam as follows: where and PM, = warp spacing. The tension S of the filling yarn is equal to its extension dlp times its elastic constant Cp: at UCSF LIBRARY & CKM on March 8, 2015trj.sagepub.comDownloaded from 766 Substituting Equation 33 into Equation 32 gives With F, Pw, Cp, the crimp H&dquo; of the filling yarn can be calculated with Equation 34 using a numerical method" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003054_6.2002-533-Figure7-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003054_6.2002-533-Figure7-1.png", + "caption": "Figure 7: Area Ratio Definitions", + "texts": [ + " In cases where a distributed shear-calibrating device was not available, it was necessary to establish a correlation to enable accurate data interpretation with the use of a point-load calibration method. To achieve this, the area ratio concept was formulated. The concept is has two sub-sets, an effective area ratio and a geometric area ratio. The geometric area ratio is based on dimensions measured from the gage as shown in Equation 3: AGAP + AHEAD ARATIOGEO ~~ ' (3) AHEAD where AGAP is the gap area and AHEAD is the sensing head area as illustrated in Figure 7. The geometric area ratio of the gages used here was 2.00. Since the rubber sheet has some flexibility, one cannot assume that the shear acting on the rubber sheet has the same effect as the shear acting on the solid head. Further, shear acting on the rubber sheet that covers the rim of the housing might also contribute slightly to the net force on the flexure. Thus, one must determine what we have called the effective area ratio. The effective area ratio quantifies the area contribution of the rubber sheet as a percentage of the floating head area" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002872_1.1456456-Figure2-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002872_1.1456456-Figure2-1.png", + "caption": "Fig. 2 Disk\u00d5spindle modes; \u201ea\u2026 \u201e0,1\u2026 unbalanced mode and \u201eb\u2026 \u201e0,0\u2026 unbalanced modes", + "texts": [ + " and ~0,1! balanced modes can no longer exist because inertia force and moment from vibration of a single disk are not self-balanced. In contrast, coupled disk/spindle modes are combinations of ~0,1! or ~0,0! disk modes with bearing deformation and spindle rigid-body motion, which includes infinitesimal translation and rocking. The coupled disk/spindle modes are also referred as unbalanced modes in literature. There are two types of unbalanced modes: ~0,1! unbalanced modes and ~0,0! unbalanced modes. Figure 2~a! illustrates the shape of ~0,1! unbalanced modes, which are also known as rocking modes in HDD industry. For this mode, all the disks vibrate with one nodal diameter in the same phase. As a result of the disk deformation, the inertia force from the left half of the disks is entirely equal and opposite to that from the right 2002 by ASME Transactions of the ASME s of Use: http://www.asme.org/about-asme/terms-of-use Downloaded F half of the disks. These two inertia forces create an unbalanced moment about the centroid. Because the disk/spindle assembly is spinning with an angular momentum about the z axis, the presence of the unbalanced moment causes the spindle to undergo a steady precession about the Z axis @1#. Often, ~0,1! unbalanced modes appear in pairs; one has spindle precession in the counterclockwise direction known as forward precession, and the other has spindle precession in the clockwise direction known as backward precession ~e.g., Fig. 2~a!!. Figure 2~b! illustrates the shape of ~0,0! unbalanced modes. For this mode, all the disks undergo identical axisymmetric deformation with the same phase. The disk deformation results in an unbalanced inertia force that has to be compensated for by the axial motion of the disk/spindle assembly @1#. In HDD applications, disk modes appear largely due to flowinduced disk vibration at high spin speed. In contrast, coupled disk/spindle modes are often excited by bearing defects and external shocks ~e.g., operation tests" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003437_0954406042690498-Figure4-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003437_0954406042690498-Figure4-1.png", + "caption": "Fig. 4 (a) Cutting edge of the blade for cutting the pinion tooth surface. (b) Relationship between coordinate systems S3 and S5", + "texts": [ + " 3, where (1) represents the blade supporter disc, (2) the cutting blade, (3) the wedge block, (4) the fastening bolt, (5) the adjusting bolt, (6) the sleeve and (7) the fastening bolt. Several rectangular straight slots are radially distributed over the end face of the blade supporter disc. Each slot holds one blade, which is fixed by wedge blocks and bolts. The radial extending length of the cutting blade is controlled by an adjusting bolt, which is installed in one sleeve of a ring. The ring with radially distributed sleeves is fixed on the blade supporter disc by several fastening bolts. As shown in Fig. 4a, Gp denotes the cutting edge of the blade for the cutting pinion tooth surface. The cutting edge Gp is designed as a parabolic curve with a parabolic parameter k1. Two coordinate systems S7\u00f0o7; x7, y7, z7\u00de and S5\u00f0o5;x5, y5, z5\u00de are applied to connect rigidly to Gp, whose equation is represented in S7 by fx7, y7, z7g \u00bc fu, k1u2, 0g \u00f01\u00de During the generation process, the blade-disc is subjected to pure rotation about a fixed axis and thus the cutting edge Gp forms a surface of revolution. To apply the coordinate transformation method to create the equation of the surface of revolution, the coordinate system S3\u00f0o3; x3, y3, z3\u00de is applied to connect rigidly to the surface of revolution and the coordinate system S5 is provided with pure rotation about the axis of the disc with parameter b. As shown in Fig. 4b, S5 coincides with S3 when b \u00bc 0. The surface of revolution formed by Gp is denoted by S3, whose equation can be obtained by transferring the coordinates of Gp from S7 to S3 according to R3\u00f0u, b\u00de \u00bc M35\u00f0b\u00deM57R7\u00f0u\u00de \u00f02\u00de where R3\u00f0u, b\u00de \u00bc fx3, y3, z3, 1gT, R7\u00f0u\u00de \u00bc fx7, y7, z7, 1gT M35\u00f0b\u00de \u00bc Tx\u00f0 rp\u00deRy\u00f0 b\u00deTx\u00f0rp\u00de M57 \u00bc Rz p 2 a Ty pc cos a 4 The unit normal vector to S3 can be determined by n3\u00f0u, b\u00de \u00bc qbr36qur3 kqbr36qur3k \u00f03\u00de Proc. Instn Mech. Engrs Vol. 218 Part C: J. Mechanical Engineering Science C04304 # IMechE 2004 where r3 \u00bc fx3, y3, z3gT" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001501_robot.2001.932548-Figure2-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001501_robot.2001.932548-Figure2-1.png", + "caption": "Figure 2: The kinematic model of the truc k and trailer of Fig. 1", + "texts": [ + " The crossing of the switching surfaces of the partition and the direc tion of crossing provide the feedback information to v. In synthesis, the switching can be seen as an ex tra feedback loop around the two different closed loop modes. The switching surfaces and the switching logic are designed in such a way that the desired equilibrium inside the backward motion regime is given the mar acter of global attractor from all the initial conditions in a prespecified domain. The differential equations describing the kinematic of the vehicle under exam (see Figure 2 for notation) can be found for example in [1]: Y3 (33 v cos (33 cos [32 1 + -\ufffd tan (32 tan IX cos (h ( M ) \"1 v cos (33 cos (32 1 + _..2:. tan (32 tan Q sin 113 ( M ) \"1 sin f3:l cos (32 ( Ml (3 ) v 1 + -- tan 2 tan IX L3 h (1) The two inputs of the system are the steering angle IX and the longitudinal velocity at the second axle v. Call p = [Y3 03 (33 (32f the configuration state obtained neglecting the longitudinal componen t X3. In a compact w ay, the state equations are written as: p = v(A(p) + B(p, IX)) ( 1) The sign of 1) decides the direction of motion" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001442_bf01573714-Figure7-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001442_bf01573714-Figure7-1.png", + "caption": "Fig. 7. Reciprocal position between the two field coils and the six armature coils at t ime t and with Z = 0 (armature nonoverlapped coils equivalent to an overlapped winding)", + "texts": [ + " 6b is similar to the one analized in 2a, provided guA = Z/2 instead of gyA = Z (and consequently six armature coils - - two for each phase -- are included in the length 2) and provided half the value of the phase current be considered. According to this, the analytical relations of the motor propulsion force developped in 2 a may be also applied to the overlapped winding, when we consider the presence in 2 of six non-overlapped equivalent coils with length LA ~ 2 / 2 - t~ and current i i / 2 . Fig. 7 is the equivalent of Fig. 5 with each phase of the armature corresponding to the configuration of Fig. 6b. Fig. 8 [1]; in this system the stator three-phase winding is made up of non-overlapped coils and the N b - - T i superconducting winding on the vehicle is used both for the motor excitation and for the vehicle levitation. With reference to Figs. 2b, 3b and 4, the following values are assumed for the sizes and pitches of LSM field and armature coils: 2 = g u A = 4 . 2 m p E = 2 . 1 m P A = 1 " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000175_0016-0032(95)00056-9-Figure12-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000175_0016-0032(95)00056-9-Figure12-1.png", + "caption": "FIG 12.", + "texts": [ + " Another approach permits us to view the network as consisting of precisely four six-bar loops of a different type. Each of them is exposed by removing a pair of links having no direct connection from the assemblage and delineating the six-bar with the remaining segments. For instance, by reference to Fig. 8, in which we now imagine that joint axes are spatially separated by the link-lengths specified above, withdrawal of bars 4-5-9 and 2-1- Vol. 332B, No. 6, pp. 657~fi79, 1995 Printed in Great Britain. All rights reserved 6 7 3 12 gives rise to the residual six-bar 6-10-11-3-8-7 , pictured in Fig. 12 with a more convenient scheme of notation for the present purpose. (The reason for the \"s tarred\" symbols is given below.) We see that each of the Bennett loops contributes one member to this composite linkage, so that it appears to be far more sophisticated a synthesis than any previously described in the literature. Examination of its closure equations, however, yields a surprising outcome. By comparison of Figs. 11 and 12, we can write 4~2 = 0~A-rc Journal of the Frankl in Institute 674 Elsevier Science Ltd Bricard's Doubly Collapsible Octahedron 1 \u2022 3 ~ 0 3 C (]~4 = 0 4 C 4~ = O~B--rc q~* = 03B + ~", + " 6 and 7 that an apparently inconsistent effect of our scheme was the variant ordering of edges in the plates surrounding vertex F. The sense of a common normal to a pair of consecutive joint axes in a spherical linkage is merely notional, and we have a reason at this stage, soon to be evident, for countermanding the property just referred to. We can do so by simply reversing the directions of the four common normals (while retaining the sequence of joint axes) circumjacent to vertex F. The result is a change in the signs of all skew angles in the four-bar loop and, with reference to Fig. 12, to introduce, say, q~, = 4 ' , * + ~ ~b6 = ~ t - ~ - (It is pointed out in this context that, although the representations of Figs. 8, 10 and 11 exhibit crossed configurations for all six four-bar loops, the physical octahedron typically features only two crossed four-bars, as in Fig. 9.) Directly from Eqs (1), (S-A), (2) and similar relationships, we see that ~ b l - ~ ' ~ 4 = 3 ~ (Y-l) q~2 + ~5 = x (Y-2) q~3 + q~6 = 2re. (Y-3) By substitution into Eqs (S-A) for 0~A and 04A, we find, after a little manipulation, that 4)1 4)2 fl,-~2 ,/sfl, +c~2 t ~ - t ~ - = s - - 2 / 2 (Y-4) In a similar manner, from equations governing loop B, we obtain , t = s - - 5 - / Other relationships could be written, but Eqs (Y-1)-(Y-5) provide a set of independent closure equations for the six-bar", + " Closure equations for the non-hybrid octahedral linkages of all three kinds are developed in (12), but are based in part there on standard sets of displacement relationships which are difficult to manipulate, and so some of the resulting equations are rather unwieldy. The recent advances made (24, 25) in treating the symmetric octahedra allow for a more direct determination of angular relationships, whence the governing equations of the relevant six-bars can be expressed in simpler formats, as in (26). We are now in a position to do the same for the doubly collapsible case. Taking the last of the loops specified above as the paradigm, it can be schematically depicted by Fig. 12. In this application, however, instead of zero offsets and link-lengths defined by the Bennett condition, we have zero link-lengths and offsets given by the structural requirements of the octahedron, as presented early in the paper, signs having to be taken into account. Remarkable as it appears, the displacement-closure equations governing the quite dissimilar types of linkage must, clearly, be identical; the common attribute of central importance is the set of relative orientations of the joint axes" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003864_50002-2-Figure2.14-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003864_50002-2-Figure2.14-1.png", + "caption": "Fig. 2.14. Computer simulation tyre model with flexible carcass, arbitrary pressure distribution and friction coefficient functions. Forces acting on a single tread element mass during one passage through the contact length are integrated to obtain the total forces and moment F x , F y and M z .", + "texts": [ + " A simpler representation of carcass complance that is experienced in the lower part of the tyre near the contact patch considerably speeds up the computation. Also the way in which the tread elements are handled is crucial. The computer simulation tread-element-following method is attractive and allows considerable freedom to choose pressure distribution and friction coefficient functions of sliding velocity and local contact pressure. The physical model that forms the basis of the latter method has been depicted in Fig.2.14. Influence (Green) functions may be used to describe the carcass horizontal compliance in the contact zone and possibly several rows of tread elements may be considered to move through the contact patch. One element per row is followed while it travels through the length of contact (or several elements through respective sub-zones). During such a passage the carcass deflection is kept constant, the motion of the single mass-spring (tread element) system that is dragged over the ground is computed, the frictional forces are integrated, the total forces and moment determined and the carcass deflection is up-dated" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002898_bf02441311-Figure6-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002898_bf02441311-Figure6-1.png", + "caption": "Fig. 6", + "texts": [ + " However, several authors such as PUSCHETT and ZURBACH (1974), QUEHENBERGER (1977) and CAFLISCH et aL (1978) have claimed that the ionic strength does not influence the antimony potential. Typical pCO 2 response of a highly purified (6N) electrode in lO mM NaHCO 3 . pO 2 = 12.8 kPa 5.2 pCO 2 sensitivity The antimony pCO2-electrode showed a linear mV response to log pCO 2 in all bicarbonate concentrations tested. A typical response obtained for five gases with different pCO 2 values, but the same pO 2 can be seen in Fig. 6. The electrode sensitivity towards pCO 2 defined as A E / A l o g p C 0 2 , for a specific electrolyte concentration did not change significantly (at the 5% level) with time, except during an initial period. This initial deviation is probably caused by the more pronounced drift in potential level, which was observed during the beginning of a measurement (see Fig. 3). The pCO 2 sensitivity against time for a bicarbonate concentration of 10mM is shown in Fig. 7. Electrodes of the same purity showed no significant difference in pCO 2 sensitivity, whereas electrodes made of highly purified (6N) antimony showed a significantly higher sensitivity than that of electrodes made of (4N) pure antimony" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001298_43.85756-Figure4-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001298_43.85756-Figure4-1.png", + "caption": "Fig. 4. Example of an interface requiring four etch vectors and generating three new segments. The etch vectors result from the two existing segments, the etch along the interface, and the etch vector of the segment being created on the other side of the interface from the dominant interface etch.", + "texts": [ + " However, this generates many nodes, is inefficient in CPU time, and gives a stair-step representation of a smooth surface. A much better approach is to calculate the segment direction that will result because of the new exposure and etch and to use the etch vector associated with that segment. The segment direction that will result is such that its intersection with the boundary moves at the same rate as the interface node. Of course, this segment grows with time as the interface exposes more material. An interface node could require four movement vectors as illustrated in Fig. 4: one movement vector is associated with each of the two original segments, a third for the movement of the interface node along the boundary, and a fourth for the resultant segment in the new exposure of material. U-M-I Due to a lack of contrast between text and background, this page did not reproduce well. THURGATE: SEGMENT-BASED ETCH ALGORITHM An interface node is allowed to be coincident with one or more other interface nodes to allow for the intersection of two or more boundaries separating different materials" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000727_s0167-8922(08)70486-3-Figure3-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000727_s0167-8922(08)70486-3-Figure3-1.png", + "caption": "Figure 3 Motion of surfaces", + "texts": [ + "3. Film thickness be written as; The film thickness in a n EHL contact can where ho represents the nominal distance of a rigid ellipsoid and a plane on the z-axis. In EHL problems the value of ho is generally negative, due to the large elastic deformation compared with the film thickness. 2.4. Kinematics Once a spin motion exists on the contact surface, the velocity distribution is a function of the position in the field. Suppose that a n ellipsoid A and a plane B come into contact as shown in Figure 3. There is a rolling/sliding motion in the x direction, but none in the y direction (no side slip). Spin motion exists as well. The motion of the contacting surfaces is defined as shown in Figure 3, i.e., on body A and B respectively; 1. The velocities of the surfaces parallel to the axis are uoA and uOs in the x direction and zero in the y direction. 2. The angular velocities of the spin motion are and COB. Counterclockwise spin motion is determined to have a positive value of w, from a view point of z = + 00. The mean velocities can be defined as follows; UOA + UOB 2 uo = w.4 + w , 2 w = 602 v, = v,, = x u , Now, the mean velocities between the two surfaces at point C are, in the x and y directions respectively; (10) At a point C(xy), the spin motion gives rise to velocity components in the x and y directions" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002302_iros.1997.655123-Figure3-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002302_iros.1997.655123-Figure3-1.png", + "caption": "Figure 3: Distance D to the origin of { W ) of the projection on the xy-plane of the line that supports the ruling segment \u20acor a given orientation corresponding to a type-A basic contact $a) and to a type-B basic contact (b).", + "texts": [ + " For type-B basic contacts, it is constant, and for type-A basic contacts it depends on the orientation of the mobile object ($Jw = $T + 4 + r , where $T is the orientation with respect to {T} of the normal to the contact edge). B: the distance from the origin of ( W ) to the projection on the xy-plane of the line that supports the ruling segment Corresponding to a given mobile object orientation 4. For a type-A basic contact D is given by (figuie 3a): D = x,cos+w + ymsin4w + d~ where (x,,ym) are the coordinates of the contact vertex of the fixed object measured in {W} . For type-B basic contacts D is given by (figure 3b): (1) D = hcos($w + 1~ - y - 4) + dw ( 2 ) where h and y are respectively the module and orientation of the vector defining the contact vertex of the mobile object with respect to {T}. Therefore, for a basic contact i the straigtht line that contains the ruling segment of the C-face for a given orientation q5 is: 2 cos($wi) + ysin($wi) = Di (3) 4 = P 4 D, being defined by equation (1) or (2), depending on the type of basic contact. 2.2 Two basic contacts Definition 4: The C-edge E;j corresponding to a contact situation involving two basic contacts i and j is the set of configurations in C where the contact situation takes place, only considering the constraints 948 imposed by the edges and vertices involved in both contacts" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003190_cca.2000.897544-Figure2-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003190_cca.2000.897544-Figure2-1.png", + "caption": "Figure 2: Unstable remaining dynamics (e = 0.5, g = 1).", + "texts": [ + "2) exhibits an uns4able (saddle) equilibrium point at the origin 8 = 0, 8 = 0 and a center around 8 = A, 8 = 0. For initial conditions with zero angular velocity, the periodic nature of the solutions of 2.2, imply a \u201crockingn motion of the aircraft around its longitudinal axis. For zero initial conditions of the roll angle and nonzero initial angular velocity, the system (2.2) is unstable and hence, as time increases, the aircraft rotates about its longitudinal axis while its center of gravity remains fixed at a constant position in the 2-2 plane (see Figure 2). 2.2 Ditferential flatness of the PVTOL model It has been shown in [4] that the PVTOL model is differentially flat, with flat output given by the horizontal and vertical coordinates (F, L) of the Huygens center of oscillation when the aircraft dynamics is reinterpreted as the dynamics of a pendulum of length e. Such outputs are given by, F = z - E s i n 8 ; L = z + e c o s e (2.3) The PVTOL aircraft system model requires a second order dynamic extension on the control input u1. Instead of taking ul and til as additional state variables, the following auxiliary variable c = u1 - e (6) is introduced as a new state variable" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003864_50002-2-Figure2.7-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003864_50002-2-Figure2.7-1.png", + "caption": "Fig. 2.7. Top-view of vehicle (Exercise 2.1).", + "texts": [ + " / The longitudinal slip spee~ V~, is obtained in a similar way\" Vsx- Vs.l The longitudinal slip x now becomes\" v V . l S X S K , ~ m J Vcx V c ' l The turn slip according to definition (2.18) is derived as follows (u_ i.t ~ 0 t ~ ~ ~ ~ ~ ~ Vcx V c ' l (2.34) (2.35) (2.36) (2.37) (2.38) with in the numerator the time derivative of the unit vector I. The wheel camber angle is obtained as indicated before: siny - - n . s (2.39) Exercise 2.1. Slip and rolling speed of a wheel steered about a vertical axis The vehicle depicted in Fig.2.7 runs over a flat level road. The rear frame moves with velocities u, v and r with respect to an inertial triad (choose (O ~ x ~ yO, z o) which at the instant considered is positioned parallel to the moving triad (B, x, y, z) attached to the rear frame). The front frame can be turned with a rate 3 (=dO/d t) with respect to the rear frame. At the instant considered the front frame is steered over an angle c~. It is assumed that the effective rolling radius is equal to the loaded radius (re = r, C* = C)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002425_cbo9780511529627.012-Figure9.19-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002425_cbo9780511529627.012-Figure9.19-1.png", + "caption": "Figure 9.19. Pairwise configuration parameters coordinates: (a) absolute and (b) relative coordinates.", + "texts": [ + " The configuration space of a system of rigid parts is a parameter space whose points specify the spatial configurations (positions and orientations) of the parts. The parameters usually represent part translations and rotations, but they can be arbitrary generalized coordinates. The configuration space dimension equals the number of independent part motions, called degrees of freedom of the system. We begin by studying a mechanical system consisting of a pair of planar parts. We attach reference frames to the parts and define the configuration of a part to be the position and orientation of its reference frame with respect to a fixed global frame. Figure 9.19(a) shows planar parts A and B, their reference frames, and their configurations (xa, ya,0a) and (xb, yb,6b). The configuration space of the pair is the Cartesian product, (xa,ya,0a,xb,yb,6b),of the part configurations. The configuration space coordinates represent three degrees of freedom of each part. An alternative representation is to describe the relative position and orientation of part A with respect to B, which is fixed and whose reference frame is at the origin of the axes, as illustrated in Figure 9.19(b). In this case, three relative parameters (u,v,is) uniquely describe the relative position of A with respect to B. The relation between the absolute and relative coordinate systems is u = (xa - xb) cos 0b + (ya - yb) sin 0b, v = (ya- yb) cosOb - (xa - (9.48) = ea - eb. The configuration space dimension for planar pairs is six for absolute coordinates, and three for relative coordinates. For spatial pairs, it is 12 for absolute coordinates Cambridge Books Online \u00a9 Cambridge University Press, 2009https://www" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002855_s0261-3069(01)00038-3-Figure3-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002855_s0261-3069(01)00038-3-Figure3-1.png", + "caption": "Fig. 3. The origin of various potential FDC defects. A cylindrical disc being fabricated (left) could possess inter-laminar, inter-road and subperimeter voids. The SEM picture on the right shows an actual FDC GS44 green part with defects, all intentionally built in.", + "texts": [ + " SEM and X-ray radiography, along with fourpoint bend tests were used to determine if WIP\u2019ing could eliminate intentionally introduced FDC build defects. The material used to build FDC parts was RU955 which had a 55 vol.% solids loading of GS44 Si N3 4 powder (Ceramic Components, Torrance, CA) in RU9 binder. The three essential types of defects i.e. interroad defects (road-to-road bonding), inter-laminar defects (layer-to-layer bonding) and sub-perimeter defects (voids in between the perimeter and the rest of the layer) are schematically illustrated in Fig. 3a,b. It should be mentioned that these defects can be eliminated using proper build strategies. Two different specimen geometries were used to evaluate the effect of WIP\u2019ing on the elimination of FDC build related defects: (1) rectangular bars; and (2) square plates. The rectangular bars were used to assess the ability of WIP\u2019ing to remove FDC inter-road defects, such as missing roads or incompletely bonded layers. To ensure a sufficient population of defects in the FDC bars, missing road defects were intentionally introduced in the center section of the bars in a staggered position at every other layer. (note: roads are running at \"458 to the specimen edges). Fig. 4a shows a schematic of the defects. One of the typical defects identified in the FDC process is the sub-perimeter void that was frequently encountered early on in the development of the process. These voids, as shown in Fig. 3b, develop in the area between the perimeter (specimen contour) and the vec- tor paths due to an inadequate impingement of the vector paths on the perimeter. Square plates (25.4=25.4=6.35 mm) with a centrally located hole (12.7 mm in diameter) and sub-perimeter voids surrounding the center hole were used to test whether WIP\u2019ing could effectively remove these defects. Build conditions for all these FDC parts (including control rectangular bars without FDC build defects) are given in Table 1. WIP\u2019ing experiments on the FDC GS44 Si N parts3 4 were performed at Long Peak Engineering (LPE, Draper, Utah)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000464_1.2048488-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000464_1.2048488-Figure1-1.png", + "caption": "Fig. 1. Schematic diagram of the spectroelectrochemical channel cell for absorption UV-visible measurements. Dimensions are not to scale.", + "texts": [ + " Measurements were performed both at steady-state and under transient conditions following the generation of a concentration pulse at a gold electrode surface using the ferro-ferricyanide system in aqueous electrolytes as a probe system. The first type of experiments showed that the predictions of the standard theory 3'4 are fulfilled quantitatively without introducing any adjustable parameters. The transient-type measurements, on the other hand, enabled the faradaic efficiency for ferricyanide production to be determined without any assumptions regarding the time evolution of the concentration profiles. Experimental The channel cell designed for these studies (see Fig. 1) was very similar to that described in Ref. 1, except that the optically transparent segments were enlarged to allow for the beam to probe a large section of the fluid past the electrode surface. The electrochemical and pumping equipment was the same as that specified in Ref. 2. Spectroscopic measurements were performed using a fiber-optics-based system, consisting of a 75 W xenon arc lamp (XM75HS, Optical Radiation Corporation) powered by an Oriel 68806 power supply and mounted in a lamp housing (Oriel 60000 Q series) equipped with a condensing/imaging lens assembly (UV F/1", + " The grating (and its mounting bracket) was removed from the monochromator housing, mounted on an optical scanner (Cambridge Technology Incorporated, Model 6650), and reintroduced into the housing. This modification was made to drive the grating directly, bypassing the original monochromator gear arrangement. A photomultiplier tube (PMT) detector (R636-10 Hamamatsu, Incorporated) powered by a high voltage supply (Ortec 556, EG&G Instrumental Princeton Applied Research) was installed in the monochromator housing. For the experiments described herein only one of the light paths shown in Fig. 1 was utilized. The data acquisition system was the same as that described in Ref. 2, except that the lock in amplifier was upgraded to a Model SR830. For the transient experiments, the average of 30 to 50 acquisitions, depending on the signal quality, was stored in the SR830 memory and then transferred to the computer for further analysis. The channel-type spectroelectrochemical cell was mounted on a precision XYZ stage to enable the fixed optical system to measure the absorbance along y, Ay, (normal to the electrode surface) at specific points along axes parallel (x) and normal to the direction of fluid flow (z) (see 4225 ) unless CC License in place (see abstract).\u00a0 ecsdl.org/site/terms_use address. Redistribution subject to ECS terms of use (see 35.8.11.2Downloaded on 2015-02-25 to IP Fig. 1), past the electrode surface. Measurements of Ay were performed at steady state along each of these latter two axes. Transient Ay curves, following a single (oxidation) current pulse applied to the electrode surface, however, were recorded mostly, although not exclusively, along the center line of the fluid flow at different distances from the electrode edge, x2. The results of these latter experiments were used to construct instantaneous or snapshot images of Ay within the channel, which served as a basis for determining spectroscopically the total amount of ferricyanide produced at the electrode surface", + " The wavelength was set at 420 nm, i.e., the maximum in the absorption of ferricyanide in this electrolyte. Steady-s ta te m e a s u r e m e n t s . - - According to the framework of approximations specified elsewhere, 3'4 which includes the neglect of transverse diffusion, the absorbance along an axis normal to the electrode surface, Ay(\u2022 due to an optically absorbing product, O, generated under steadystate, diffusion-limited conditions at the surface of an electrode in a channel-type cell (see Fig. 1), maybe shown to be given by (X2) = eoC~(IDoh /3 U) ~/s fo Clim(~'x2)d~ As = (2/3) ~/3 eoC~(IDoh2a/V) ~/3 fo ~ Cl~r~ (~,x2)d~ [1] The symbols in this equation have the following meaning X2 = x2/l: dimensionless distance, where x2 is the actual distance along the direction of fluid flow measured from the downstream edge of the electrode and l is the electrode length, cm. %: molar absorptivity of the product, cm2/mol C~: bulk concentration of the reactant mol/cm 3 Do: diffusion coefficient of the product, cm2/s h: half cell thickness cm _U: average flow velocity cm/s Clim(~;X2 ) [ : Co,]ira(~;X2)/cb]: dimensionless product concen- trat ion at the diffusion-limited current w: electrode width, cm V: flow rate, ml/s a: channel width, cm = (3U/hDol)l/3y: dimensionless distance along an axis normal to the electrode surface, where y is the actual distance along the same axis :C~m(~,\u2022 I(~) \u2022 - g(\u2022 [2] is the integrated concentration profile of the product along the dimensionless axis normal to the electrode surface over the full channel height, defined hereafter as I(~)", + " Assume further that the pulse is sufficiently short so that, after a certain period of time, t, the entire concentration pulse is confined within the downstream edge of the electrode (or more precisely, the edge of the electrode Kel-F cast piece) and the downstream end of the transparent segment of the cell, denoted as Lc~ in Fig. i. Under these conditions, the total number of moles of 0 produced at the electrode, denoted hereafter as N~ will be contained within the volume of the channel accessible to the optical beam, Vbe~. Hence N~ = f f f Co(X, Y, z; t) dxdydz Vbeam [4] where t represents the time at which the conditions prescribed above are fulfilled. If the concentration profile of O along y past the electrode edge is independent of z (see Fig. 1) across the full width of the electrode w and zero otherwise, (as the experiments described later in this work indicate), N~ will be given by s N~~ = W J J Co(X, y; t) d x d y [5] A beam where Abeam is 2hLeeu. Hence, for a fixed x = x' , the integrated concentration profile along y can be expressed in terms of the transient absorbance along that axis, Ay (x'; t), as follows ~ h Co(X', y; t )dy = A~(x'; t)/eo [6] where Co, as before, is the absorptivity of the product, and Eq. 5 can then be rewritten as N~~ = W/eo ~ Ay(x; t ) dx J L cell [7] As is shown below, the function Ay (x; t) can be constructed from plots of Ay as a function of time measured at various fixed distances from the electrode edge over the entire Lceu", + " The theoretical curve (see solid line) was obtained from Eq. 2, whereas the experimental absorbances (solid squares) were converted to dimensionless values using Eq. 1 and the physical dimensions of the electrode, l = 0.5 cm; w = 0.6 cm, the height of the channel h = 0.160 _+ 0.003 cm, c~ = 1.0 9 10 -2 M, Do = 7.3 9 10 -6 cm2/s, 5 V = 0.104 ml/s, and eo (O = [Fe(CN)8] 3-) in this electrolyte, 1.02 - 106 cm2/mol. Essentially identical measurements aimed at probing the flow along the z- axis, i.e., parallel to the downstream electrode edge (see Fig. 1) were performed at two fixed values of \u2022 namely, near the edge of the working electrode Kel-F piece and near the end of the optically transparent section downstream. The results obtained indicated that Ay, and thereby I(~) is independent of z along the width of the electrode, w, decaying rapidly to zero away from the electrode side edges? This observation supports the assumption made in the theory regarding the neglect of transverse diffusion. It can therefore be concluded that, within experimental error, the channel cell employed in these studies reproduces quantitatively the steady-state, diffusion-limited behavior predicted by the standard theory" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003189_cdc.2000.912260-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003189_cdc.2000.912260-Figure1-1.png", + "caption": "Figure 1: Scheme of the telescope", + "texts": [ + " It will be installed in a Boeing 747 for operation in the stratosphere and will enable scientist to observe infrared sources inaccessible to ground based observatories. In opposition to space based telescopes free access to the science instrument (camera), demanded by the scientist, will be possible. The primary mirror diameter (aperture) will be 2.5 m and the focal length 49 m. Detailed information on the science objectives and capabilities of the observatory are given by Krabbe and Roser (1999). To enable an unhampered entering of the infrared light, the primary mirror is situated in a cavity of the telescope directly exposed to the outside air (fig. 1). A bulkhead separates the cabin from the cavity and supports the telescope. The fundamental idea of the design is the bearing of the telescope in its centre of mass for insulation from aircraft rotational excursions. The attitude control loop, which is in the focus of this paper, is required for the stabilization of the telescopes inertial attitude. It consists of a gyro, sensing the inertial attitude of the telescope, a torque drive and the controller. The extreme environmental conditions occurring in the aircraft are not comparable to earthbound or space telescopes" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001240_6144.774758-Figure2-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001240_6144.774758-Figure2-1.png", + "caption": "Fig. 2. Solid drug delivery device [1] (courtesy of IEEE).", + "texts": [ + " Solid drug delivery device\u2014Solid drugs are normally delivered by implanting a highly concentrated dose of a drug subcutaneously so that it is slowly dissolved by the body thereby releasing small amounts of the drug. Unfortunately, not all patients dissolve the drug at the same rate, and the traditional method requires permanent exposure to the drug. If a reliable delivery device that opened and closed in response to external stimuli (either from sensor input or on the patient\u2019s command) could be created, many of the problems associated with solid drug delivery would be solved (see Fig. 2 for an example). Such a device requires sensors, actuators, electronic logic, a power source, and must be compatible with the drug it contains. 1521\u20133331/99$10.00 1999 IEEE II. WHAT IS In2m? Over the last several years various industries have been developing nano, micro, and millimeter scale technologies, which have resulted in components ranging from quantum transistors, to widely commercialized integrated circuits, to microelectromechanical sensors. A common emphasis of these fabrication industries has been on the integration of different functions in miniaturized systems [2]; however, the technology currently used to realize these systems is monolithic" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003887_acc.1995.529310-Figure2-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003887_acc.1995.529310-Figure2-1.png", + "caption": "Figure 2. Coordinate Definition", + "texts": [ + " &nce most students learn control systems analysis, desi n and simulation techni ues w t h the aid of MATfAB or MATRIXx, we w i i be converting our hardware to run with the IS1 AClOo/C30 real time prototyping system 1161. Experience with other experiments in our laboratory has shown us that this choice works best with a large number of students. 3. Mathematical Models ke\" functions. This was required for the results re Ere. Or, one can pro ram the DSA cards diredy. Po\"& The 3.1 Nonlinear Equations The choice of coordinates and labelin convention is shown on figure 2. First, the equations o$ motion for the ivot arm and pendulum are easily derived using the Lagran ian formulation. A suitable motor model as well as modgrling of the hd ion effects of the motor and drive shaft are then added to define the complete system model. The kinetic energy of the system is: 1. 1 . . . 1 . 2 2 T = -m,,0: + m,,0,0, + -m& m,, = I, + I , +m&: +m,L2, +m,Li +2m,L, codz m,, = m,, = I , + mzL$ + m,Li + m,L, cos0, m,, = I, + m,L', +mtL: m , = m , + m , + m , m, = mzLd + mtLz where The potential energy is: From the Lagrangian equations: v = g[m&] e o s ~ , + m, cod^, + e,)] 2 d if(T-V) O(T-V) E( )- (7) = 3" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003675_j.triboint.2003.11.004-Figure6-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003675_j.triboint.2003.11.004-Figure6-1.png", + "caption": "Fig. 6. Cam profile details, R1 = 6 mm, R2(a = 28 mm, b = 38 mm, c = 46 mm), r1 = 74 mm, r = 80 mm.", + "texts": [ + " During the test procedure on the experimental machine tool (Fig. 4), the cutting conditions were: Machine tool: Capstan lathe Cutting tool: Turning tool (cutter) Speed: 100\u2013300 (m/min) Feed: 0.1\u20130.4 (mm/revolution) Depth: 0.2\u20131.5 (mm) Coolant: Microemulsion (4%) Workpiece material: Medium carbon steel (YU) Contact conditions in a cam and roller follower contact for machine tools have been studied. Fig. 5 shows experimental cam and roller follower. The cam profile details and roller follower geometry are shown in Fig. 6. The contact load between cam and roller follower was set 930 N in maximum value. Working conditions of these cams (Fig. 4) and other influential factors are shown in Table 2. The results of the experimental investigation are pre- sented for two viscosity grades and three types of lubri- cating oils: . Hydraulic oils, category symbol HM, ISO VG 46 and 68 (first period of time) . Hydraulic oils, category symbol HG, ISO VG 46 and 68 (second period of time) . Gear oils, category symbol CKB (C), ISO VG 46 and 68 (third period of time)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002689_a:1014805931636-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002689_a:1014805931636-Figure1-1.png", + "caption": "Fig. 1", + "texts": [ + " The emphasis was on the neighborhood of the double zero eigenvalue of the linearization matrix. Here we will study a domain in the parameter space such that fork bifurcations may occur in the phase space of an inverted double-link mathematical pendulum with an elastically fixed upper end and an asymmetric follower force. These bifurcations lead to fold and cusp catastrophes in the manifold of stationary states. 1. Formulation of the Problem. Denote by P \u2192 , \u03b1 = \u03b4 + k \u03d52, \u03b4 = const, the angle between the follower force P and the vertical (Fig. 1). Let m1 and m2 be the masses of the material points A1 ( x1 , y1 ) and A2 ( x2 , y2 ), the links O A1 = l1 and A1 A2 = l2 be imponderable, \u00b51 be the coefficient of viscosity in the lower joint O taking into account the external friction, \u00b52 be the coefficient of viscosity in the intermediate joint A1 taking into account the internal friction in the system, c be the stiffness of the elastic fixation of the upper end, and c1 and c2 be the spring forces in the joints O and A1, respectively. We assume that when the links are in the vertical position (\u03d51 = 0 and \u03d52 = 0), the horizontal spring is undeformed" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001007_jsvi.1997.0998-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001007_jsvi.1997.0998-Figure1-1.png", + "caption": "Figure 1. The belt model.", + "texts": [ + " The transient torque of the motor is calculated with the presented program, and is used in calculating the responses of belt systems. Secondly, to verify the usefulness of the presented program, the calculation results obtained with it are compared with experimental results of a test rig driven by an actual induction motor. The following assumptions are made for deriving the equations of motion for belt systems. (1) A belt is replaced by a spring with a stiffness coefficient kij and a damping coefficient cij , as shown in Figure 1. The stiffness coefficient of a belt between a pulley i and a pulley j, kij , is expressed by (a list of nomenclature is given in the Appendix) kij =AE/Lij , (1) where A is the cross-sectional area of the belt, E is the elastic modulus of the belt, and Lij is the distance between a pulley i and a pulley j. (2) The mass and volume of the belt are ignored. (3) There is no slip between the belt and pulley. (4) The bending stiffness of the belt is negligible. Complicated belt systems consisting of multiple belts and pulleys are regarded as combinations of single belt systems consisting of a single belt and several pulleys" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003702_gt2004-53860-Figure3-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003702_gt2004-53860-Figure3-1.png", + "caption": "Fig. 3 Experimental test rig", + "texts": [ + " Mathematical models and approaches, such as perturbation and linearization, which attempt to describe the behavior of a system with dry friction, are not generally successful due to high nonlinearity, especially in the vicinity of equilibrium point. The outcomes of experimental studies also show that the results are affected by time, humidity, interface properties and history of motion [5-7]. EXPERIMENTAl.. TEST RIG An experimental test rig was designed and fabricated in order to conduct dynamic tests on a complete compliant bump foil bearing damper ring assembly. The experimental test rig, shown in Fig. 3, included the following main components: 2 Copyright \u00a9 2004 by ASME url=/data/conferences/gt2004/71222/ on 07/07/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use Do a. Electrodynamic Shaker, with controller and stinger assembly to transfer motion to the bearing damper assembly through an inner ring and disk b. Large base plate on which the test rig assembly was mounted c. Weldment housing, inside which the compliant test elements were mounted d. Inner ring/disk assembly with stinger attachment e", + " As such a built-in air load support with a maximum payload capacity of 100 kg (220 Ib) and full relative displacement was included in the shaker capabilities. Detailed specifications of the shaker are given below: Peak force - sine wave: 4,100 N (922 Ibf) Frequency range: 5-5,000 Hz Maximum acceleration: 951 m/s2 (97 g) Maximum velocity: 1.0 m/s (39.4 in/s) Maximum displacement: 25 mm (0.984 in) Resonant frequency: 3,150 Hz Weldment (Foil bearing housing) The weldment, made of hard rolled steel, was used for holding the compliant foil bearing damper assembly. The weldment hardware, shown in Fig. 3, has overall dimensions of 51\"25\"38 (L*W*H) cm and was secured to the base table via 16 high strength screws. A passage hole was made at the top dead center of the weldment, which allowed for connection of the stinger to the central disk, as well as accommodating the reciprocating motion of the stinger. In order to increase the accuracy of the displacement measurement during the dynamic tests, any detrimental movement by the foil bearing housing had to be minimized. In particular, it was essential that the test system be free of resonances within the desired frequency test range as much as possible", + " The stinger was a solid rod made of PH13-8Mo with a 25.4 mm in diameter and 23.038 cm in length. Both ends of the stinger had inside threaded holes for connection to the inner ring/disk assembly and shaker. Stinger Hydrostatic Guide Bearing In order to minimize undesired off axis motions being introduced, due to potential misalignment between the stinger and the surface of the passage hole at the top dead center of the weldment, an air bearing was used to guide the stinger motion. The air bearing, as shown in Fig. 3, was made of 1010 steel, designed with a diametral clearance of 0.05 mm and used shop air with a maximum pressure of 150 psi. The air bearing housing was mounted on top of the weldment and aligned with the center of the hole on the top of the weldment. Bump Foil Assembly An intermediate ring, made of 17-4 PH, was placed between the weldment and the central disk. The intermediate ring was kept inside the weldment via interference fit. A hole was made at the top of the ring to allow the shaker stinger rod to pass through and be connected to the central disk assembly" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003670_j.jappmathmech.2004.11.006-Figure3-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003670_j.jappmathmech.2004.11.006-Figure3-1.png", + "caption": "Fig. 3", + "texts": [ + " However, if the friction forces at the point where the wheels are in contact with the surface exceed the limit value of the Coulomb dry-friction forces, the wheels begin to slide and the motion of the system will not be described by Eqs (1.3). Hence, we will determine the friction forces at the points where the wheels are in contact with the surface, which enables us to estimate the characteristics of the construction and the parameters of the \"accelerating\" moment M (its amplitude and frequency), for which sliding begins. In deriving the required relations we will use the general theorems of dynamics, which are written for two subsystems (the platformAB and the front wheeled pair B), shown in Fig. 3. Here, for simplicity, we will confine ourselves to the case of weightless wheels. To estimate the realizability of non-holonomic constraints we will use Coulomb's axiom IRAI < f N A, IgsI -< f N B (5.1) wheref is the coefficient of dry friction, and NA and NB are the reactions of the support to the surface. After appropriate calculations, we obtain the following system of inequalities I mVct I< (mca + m s b ) ~ + mc (b - b2 I(,,,ca I < f g ( m c a + m B b ) + mBb)b2cos~ s-~n~ - b \" (5.2) which defines the domain of variation of the variables F, 12, I~, ~, in which the constraint equations (1" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002014_elan.1140010606-Figure8-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002014_elan.1140010606-Figure8-1.png", + "caption": "FIGURE 8. CV curves at the L/L interface with a mixed solvent of CB plus BN as the organic phase. W, 0.01 M LiCI; 0, 0.01 M TBATPB. CB : BN: 1, 8 : 2; 2, 6 : 4; 3, 4 : 6; 4, 2 : 8. Scan rate, 40 mV/s; A: vs. TBA+ISE.", + "texts": [ + " Table 4 shows the physical properties of mixed solvents and the values of potential window. WIo-DCB-BC. Figure 7 shows the CV behavior of the WIo-DCB plus BC mixed-solvent system with 0.01 M LiCl in the aqueous phase and 0.01 M TBATPB in the organic phase. Good transfer behavior can be obtained for systems with o-DCB-BC volume ratios of ratios X : 2, 6 : 4 , 4 : 6 , and 2:8. W/CB-BN. The CB plus HN mixture was also found to be a good system for electrochemical measurements at the L/L interface. Figure 8 shows the CV curves of the CB plus BN mixed solvent with different volume ratios; these suggest that reasonable CR : BN volume ratios are 8 : 2-4 : 6. The physical properties of the mixed solvents o-DCB plus BC and CB plus BN are shown in Table 5. CV Behavior for New Electrolytes Used in the Organic Phase The potential window depends to a great extent on the hydrophobicity of the organic phase\u2019s supporting electrolyte ions. In order to extend the potential range in the negative direction, CPTPB, EVTPB, and THDATPB are introduced as the electrolytes in the organic phase" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003585_s0219878904000069-Figure6-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003585_s0219878904000069-Figure6-1.png", + "caption": "Fig. 6. Unleaned posture about contact point h.", + "texts": [ + " Since the positions or orientations of the unleaned posture about contact point h may differ from those about contact point j, the touching point position xL Rnj will change to xLh Rnj with respect to the base frame O \u2212 XY Z. In general, the change is uncertain for an unknown object. To know the changes, we will get a new distance vector \u2206Ehj by performing Step 6 of the operation procedure to the unleaned posture about contact point h. Then, a new frame Chj\u2212UV W from Cj\u2212UV W can be set on the fingertip touching point xLh Rnj (see Fig. 6). The transform matrix ORChj from Chj \u2212 UV W to O \u2212 XY Z is ORChj = [ \u2206Ehj \u00d7 (\u2212g) \u2016\u2206Ehj \u00d7 (\u2212g)\u2016 \u2206Ehj \u2016\u2206Ehj\u2016 \u2212g \u2016g\u2016 ] \u2208 R3\u00d73. (46) Accordingly, in the unleaned posture about contact point h, the passing-C.M. line LTj is represented with respect to O \u2212 XY Z by xOj = phj + kjT h0j, (47) phj = xLh Rnj + ORChjpCj, (48) T h0j = ORChj ORT CjT 0j . (49) By computing the intersection point of the different lines LTj (j = 1, 2, . . . , h), the center of mass can be estimated. Under the unleaned posture about contact point j, let us consider estimating the center of mass with the plural (j = 1, 2, " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003795_s1554-4516(05)02009-0-Figure7-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003795_s1554-4516(05)02009-0-Figure7-1.png", + "caption": "Fig. 7. Cyclic voltammograms of 1 mM Fe\u00f0CN\u00de3K=4K 6 at s-BLM-coated GC electrode in solutions containing (a) 0.2 mM Eu3C, (b) 0.4 mM Eu3C, (c) 0.6 mM Eu3C, (d) 0.8 mM Eu3C, (e) 1.0 mM Eu3C, (f) 1.2 mM Eu3C, (g) 1.4 mM Eu3C, (h) 1.6 mM Eu3C, (i)1.8 mM Eu3C, (j) 2.0 mM Eu3C. Scan rate, 50 mV sK1.", + "texts": [ + " The binding of cationic Eu3C to a lipid polar head group rendered the lipid surface (partially) positively charged, which might then facilitate the transport of anionic Fe\u00f0CN\u00de3K=4K 6 to the surface of the electrode. On the other 6 at bare GC electrode with 0.1 M KCl as supporting electrolyte. Scan rates are all 50 mV sK1. hand, the lipid charge state might affect the electron-tunneling kinetics of anionic Fe\u00f0CN\u00de3K=4K 6 due to the electrostatic interactions. The cyclic voltammograms of the GCEmodified by BLM in Fe\u00f0CN\u00de3K=4K 6 solution containing different concentration of Eu3C were recorded in Fig. 7. Each peak current was obtained after an interaction between Eu3C and lipid reaching equilibrium. With increasing concentration of Eu3C, the redox peaks were increasingly obvious. Simultaneously, peak separation (Fig. 8(a)) decreased and peak currents (Fig. 8(b)) of the redox increased. From Fig. 8(a), we can see that, from 0.2 mM (473 mV) to 1.0 mM (150 mV), the peak separation decreased rapidly, and from 1.0 mM (150 mV) to 2.0 mM (130 mV) it decreased slowly. The decrease of peak separation meant the electro-transfer velocity increased" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001309_6.1998-4212-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001309_6.1998-4212-Figure1-1.png", + "caption": "Fig. 1 Planar pursuit geometry", + "texts": [], + "surrounding_texts": [ + "In this section, we exploit the numerical simulations to justify the feasibility of the above theoretical derivations. To make the analysis independent of the physical D ow nl oa de d by U N IV E R SI T Y O F N E W S O U T H W A L E S on J ul y 30 , 2 01 5 | h ttp :// ar c. ai aa .o rg | D O I: 1 0. 25 14 /6 .1 99 8- 42 12 units, we normalize the governing equations to their dimensionless forms in terms of the following dimensionless variables. The relative distance r is normalized to r = r/rfa the relative velocities VT and Vg are normalized to Vr = VT/Vo and Vg = Vg/Vp, respectively, where V0 = ^/V?0 + V* = ^ f 2 + (r0e0)2; the acceleration components ur and ug are normalized to ur = ur/(V0 2/r0) and ug \u2014 ug/(Vg/r0}, respectively; the time t is normalized to T = t/(r0/Vo). In the following, the variables with over-line symbol denote dimensionless variables. The performance and robustness of four missile guidance laws will be tested and compared. Their acceleration commands are listed below. (Ml) HOO guidancejaw with 7 > 1: ur = \u2014p2Vg/(2fo), ug = -p2VTVg/r0. _ 2 (M2) ffoo_guidance law with 7 < 1: UT = XVe/fo, ug = 2\\VTVg/ro, where A is constrained by Eq.(32). (M3) Pure proportional navigation10 (PPN) guidance law: ur \u2014 \u2014V]tf(d6/dT)sm(k6_+ ao), ug = V M(d6 / dr] su\\(kd + ao), where VM is the normalized missile's velocity; k = A \u2014 1 with A being the navigation gain; QO is the initial aspect angle of VM with respect to the inertial reference line (see Fig.l). (M4) Realistic true proportional navigation10 (RTPN): ur = 0,ue = -X(d0/dr)VT. \u2022 Maneuvering targets: In this part, we employ seven maneuvering strategies for targets to test the robust capability of the HOO guidance laws. These seven different targets which are widely discussed in the literature include: (Tl) Smart target: wr = Q,wg = XT/(rdO/dr). (T2) Modified smart target12: wr = 0, wg = XT/(dr/dr-de/dT). (T3) Sinusoidal target11: wr = wg = Arsin(rd0 / 'dr) (T4) Ramp target11: wr \u2014 wg = \\TT (T5) Step target11: WT = Q,wg = \\T (T6) RTPN target11: wr = Q,ws = -XT(dB/dr)VT. (T7) PPN target11: wr = -VT(d8/dT)sm(k6 + /30), we = VT(d9/dT)sm(k6 + (30), where VT is the normalized target's velocity; k = XT \u2014 I with AT being the target's navigation gain; 0o is the initial aspect angle of VT with respect to an inertial reference line. Here, we choose the target's navigation gains XT and the initial engagement conditions as the uncertainty disturbances to test the robust property of the Hx guidance laws. From the definition of system's L^-gain (see Eq.(12)), we recognize that when 7 is small, it indicates that good interceptive performance can be preserved in the presence of large target maneuvers (i.e.large || w \\\\z). In the following, the simulation will show that using Hx robust guidance law the system's I/2-gain can be kept below 7 under the variations of target's maneuvering strategy, under the variations of target's navigation gains, and under the variations of the initial engagement conditions in homing phase. (A) Robustness with respect to varying target's maneuvering In Fig.4, we show the variations of the target's maneuvering energy with respect to the target's navigation gain under the action of the HOO robust guidance law with 7 > 1. There are two abscissas in Fig.4, where the one with scale larger than 1 is fitted for step target, ramp target, sinusoidal target, RTPN target, and PPN target; the other one with scale less than 1 is fitted to smart target, and modified smart target. This figure is very helpful for us to select the reasonable range of each target's navigation gam. For example, for RTPN target, the gain can be selected as 10, but for smart target the gain's reasonable range is between 0.6 and 0.8 as can be seen from Fig.3. Since in practical implementation the target's maneuvering energy must be finite, and if the gain does not fall within the appropriate range, the target's energy will approach to infinite. Although the Hx robust guidance law can maintain admissible performance for almost arbitrary maneuvering targets, it is restricted to pursue the targets whose acceleration energy is belong to the set of 1/2, i.e., || w \\\\2 is finite. Indeed, the target with infinite acceleration energy does not exist in the real world. Now we demonstrate the robust property of the Hex guidance law with respect to the target's varying navigation gains as selected from Fig. 3 for different missile guidance laws. In Fig.4 and Fig.5, the two Hx guidance laws demonstrate the robustness ability against the seven aforementioned maneuvering targets, being able to kept the L,2-gain smaller 7 for different target's navigation gain and for different target's maneuvering strategies. A comparison between the results of Fig.4 and Fig.5 reveals that the robustness of the HOO guidance law with 7 < 1 is better than that of the Hx guidance law with 7 > 1. More precisely, the Hx guidance law with 7 < 1 can maintain the disturbance attenuation level below one, but the HOO guidance law with 7 > 1 only can attain some value which is larger than one. On the other hand, if the missile's guidance law employs the conventional proportional navigations such as RTPN and PPN (see Fig.6 and Fig.7, respectively), the disturbance attenuation level may diverge for some specific maneuvering targets. We can find from Fig.6 that the RTPN missile guidance law has poor ability to pursue step target and sinusoidal target; while from Fig.7 we see that the PPN missile guidance law has poor ability to pursue step target. In summary, among the four missile guidance laws, the HOO guidance laws, es- D ow nl oa de d by U N IV E R SI T Y O F N E W S O U T H W A L E S on J ul y 30 , 2 01 5 | h ttp :// ar c. ai aa .o rg | D O I: 1 0. 25 14 /6 .1 99 8- 42 12 pecially for 7 < 1, exhibit the most robust performance with respect to the variation of the target's navigation gains. (B) Robustness with respect to varying initial engagement conditions In this part, we wish to investigate the robust ability of HOC guidance laws with respect to the variations of the initial engagement conditions. For convenience, we introduce the initial angular momentum ho = r^do as an index to reflect the impact of initial conditions. Fig.8 to Fig.11 express the robustness of the four missile guidance laws with respect to changing ho for the seven different target's maneuvers. The abscissa in these figures is the normalized initial angular mom entum ho = ho/(roVo). The magnitude of hqjs between 0 and 1. For a tail-chase initial condition, ho = 0, while for the worst_case where the target is escaping fully tangentially, ho is equal to one. Hence, when we increase ho from 0 to 1, we impose increasingly stringent engagement conditions on the missile. It is observed that for any ho between 0 and 1, the two kinds of H^, guidance laws can maintain excellent disturbance attenuation ability, while the performance of RTPN and PPN fluctuates dramatically when changing ho. As expected, the robustness against initial condition variation for H^ guidance law with 7 < 1 is better than that with 7 > 1, being able to maintain the I/2-gain below 1. It is emphasized that the ability of maintaining the Z^-gain below 1 for HQQ guidance law with 7 < 1 is not only valid for the seven targets considered here, but also valid for all the targets with finite accelerations. The Hx guidance law with 7 > 1 only can maintain the \u00a3.3 -gain below some value about 25. On the other hand, the performance of RTPN and PPN guidance laws is divergent for some specific targets and under some specific range of ho (see Fig. 10 and Fig.ll). VI. Conclusions The nonlinear Hx control theory has been exploited in this paper to design guidance law for homing missiles. By regarding target's maneuvers as disturbances inputs, we reformulate the missile guidance problem as a nonlinear disturbance attenuation control problem. After solving the associated Hamilton-Jacobi Partial differential inequality, we obtain three kinds of Hx guidance laws which possess excellent performance robustness against maneuvering targets and against variations of initial engagement conditions. As compared with the proportional navigation schemes, the performance of the proposed HOC robust guidance laws is shown to be more insensitive to variations of target's maneuvers. References 1. Hyde, R.A., Hx Aerospace Control Design, Springer, London, 1995. 2. Gilles, F., and Mohammed M., \"Parametric Robustness Evaluation of a HOC Missile Autopilot\", Journal of Guidance, Control, and Dynamics, Vol.19, pp.621-627, May-June 1996. 3. Jeff, B.B., and David, L.K., \"Parametric Uncertainty Reduction in Robust Missile Autopilot Design\" , Journal of Guidance, Control, and Dynamics, Vol.19, pp.733-736, Feb.-March 1996. 4. Miele, A., and Wang, T., \"Near-Optimal Highly Robust guidance for Aeroassisted Orbital Transfer\" , Journal of Guidance, Control, and Dynamics, Vol.19, pp.549-556, May-June 1996. 5. Kang, W., \"Nonlinear Hx Control and Its Applications to /rigid Spacecraft\" , IEEE Trans. Automatic Control, Vol.40, pp.1281-1285, 1995. 6. Yaesh, I., and Ben Asher J.Z., \"Optimum Guidance with a Single Uncertain Time Lag\", Journal of Guidance, Control, and Dynamics, Vol.18, pp.981-988, Sep.-Oct., 1995. 7. Van der Schaft, A. J., \"I/2-gain analysis of nonlinear systems and nonlinear state feedback HOS control\", IEEE Trans. Automatic Control, Vol. 37, pp. 770-784, 1992. 8. Dalsmo, M., and Egeland, 0., \"State feedback Hx control of a Rigid Spacecraft\", Proceedings of the 34th Conference on Decision and Control, December 1995. 9. Patpong, L., Sampei, M., Koga, M., and Shimizu, E., \"A Numerical Computational Approach of Hamilton-Jacobi-Isaacs Equation in Nonlinear H Control Problems,\" Proceedings of the 35th Conference on Decision and Control, pp.37743779, Kobe, Japan, Dec. 1996. 10. Yang, C.D., and Yang, C.C., \"A Unified Approach to Proportional Navigation,\" IEEE Trans. Aerospace and Electronic Systems, Vol.AES-33, No.2, pp.557-567, 1997. 11. Ohlmeyer, E.J., \"Root-mean-square miss distance of proportional navigation missile against sinusoidal target\", Journal of Guidance, Control, and Dynamics, Vol.19, pp.563-568, 1996. 12. Rao, S.K., \"Comments on \"True Proportional Navigation with Maneuvering Targets\", Journal of Guidance, Control, and Dynamics, Vol.33, pp.273274, 1997. D ow nl oa de d by U N IV E R SI T Y O F N E W S O U T H W A L E S on J ul y 30 , 2 01 5 | h ttp :// ar c. ai aa .o rg | D O I: 1 0. 25 14 /6 .1 99 8- 42 12 Fig. 2 Relation of X. and y for various p Fig. 3 Relation between target navigation gain and target acceleration magnitude PPN target . modified smart target RTPN .smarttarget target /sjnusoid3| target ramp target target navigation gair Rg. 4 Robustness of (.j-gain for H, guidance law with llzl 1.2 target navigation gain ; Fig. 5 Robustness of 1.,-gain for H. guidance law with y < 1 initial condition Jft ToVo Fig. 9 System Lj-gain v.s. h~, for H. guidance law with y < 1 D ow nl oa de d by U N IV E R SI T Y O F N E W S O U T H W A L E S on J ul y 30 , 2 01 5 | h ttp :// ar c. ai aa .o rg | D O I: 1 0. 25 14 /6 .1 99 8- 42 12 0.1 0-2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 initial condition JT = -ii\u00a3- r0Vo Fig. 10 System Lj-gain v.s. hi for RTPN guidance law D ow nl oa de d by U N IV E R SI T Y O F N E W S O U T H W A L E S on J ul y 30 , 2 01 5 | h ttp :// ar c. ai aa .o rg | D O I: 1 0. 25 14 /6 .1 99 8- 42 12" + ] + }, + { + "image_filename": "designv11_65_0003159_12.470666-Figure11-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003159_12.470666-Figure11-1.png", + "caption": "Fig. 11 Simple geometry for axis/wheel combination", + "texts": [], + "surrounding_texts": [ + "As mentioned above the liquidus point of brass (CuZn37) is at about 920 \u00b0C whereas the stainless steel S 20AP melts at a much higher temperature (about 1400\u00b0C). Hence, the process has to be controlled in a way that the steel part reaches first the melting temperature and the energy is transferred by means if heat conduction to the brass. This is shown in Fig. 1. Due to the lower absorption coefficient of brass this effect is achieved by positioning the laser spot slightly on the steel axis. Most of the energy is absorbed in the steel and the brass is heated by heat conduction only. Material Axis S20AP Wheel CuZn37 \u2205 0.3 mm Material Inner wheel S20AP Wheel CuZn37 \u2205 1.9 mm Proc. SPIE Vol. 4637 579 Downloaded From: http://proceedings.spiedigitallibrary.org/ on 06/16/2016 Terms of Use: http://spiedigitallibrary.org/ss/TermsOfUse.aspx" + ] + }, + { + "image_filename": "designv11_65_0002181_iecon.1996.570948-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002181_iecon.1996.570948-Figure1-1.png", + "caption": "Fig. 1. Coordinates", + "texts": [ + " c, : CO : Cft : Cbi : coi : C f i : xo : Ro : z f i : Rf i : inertial base frame a coordinate frame attached to the center of mass of the object a coordinate frame attached to each finger the point of contact in Cb the point of contact in CO the point of contact in Cf> the position of the origins of CO in Cb the orientation of the origins of CO in Cb the position of the origins of Cf , in Cb the orientation of the origins of Cf , in Cb For the analysis, consider a finger in contact with an object as shown in Fig.1. Let cb be an inertial base frame. Denote by C, a coordinate frame attached to the center of mass of the object, and by Cf , attached to each finger. Further, denote the point of contact in C,, C,, and C f * , as Cb, , co%, and c f , . Define z, zf% to be the position of the origins of CO, Cf, in the base frame Cb, and R,, R f s l are the rotation matrices giving the orientations of CO, Cft in the base frame Cb, respectively. Using the above definitions, the point of the contact is described in two ways, one for the finger arid the other for the object" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000167_jsvi.1997.1078-Figure2-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000167_jsvi.1997.1078-Figure2-1.png", + "caption": "Figure 2. Elastic domain evolution in the load\u2013internal hardening plane as a consequence of isotropic hardening.", + "texts": [ + " Experimental observations establish that plastic displacement has the following evolution equation [2]: x\u0307p = g sign (P\u2212 b), (4) where g represents the rate at which plastic displacements take place. The yield surface, which defines the elastic domain limit, is expressed by h(P, a, b)= =P\u2212 b =\u2212(Py +Ga)=0, (5) where G is the plastic modulus. The form of this function shows that kinematic hardening causes the elastic domain translation, while isotropic hardening causes the expansion of this domain. Figure 2 shows the expansion caused by the isotropic hardening in the load\u2013internal hardening plane. The irreversible nature of plastic flow is represented by means of the Kuhn\u2013Tucker conditions [3]. Another constraint must be satisfied when h(P, a, b)=0. It is referred to as the consistency condition and corresponds to the physical requirement that a load point on the yield surface must persist on it [2]. These conditions are presented as follows: ge 0, gh(P, a, b)=0, gh (P, a, b)=0 if h(P, a, b)=0. (6) Considering a single-degree-of-freedom oscillator with mass m and an external linear viscous dissipation parameter c, the balance of linear momentum is expressed by the following equation: mx\u0308+ cx\u0307+P(x, xp, a, b)=F(t), (7) where P(x, xP, a, b) is the elasto\u2013plastic restitution force of the oscillator and F(t) is an external force" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003321_978-1-4613-4145-1_6-Figure6.23-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003321_978-1-4613-4145-1_6-Figure6.23-1.png", + "caption": "Figure 6.23 Theoretical volt ammo grams for a reversible charge transfer reaction involving adsorption of either the reactant or product.(48) QJ = (E - El/2) n. The dashed curves correspond to no adsorption.", + "texts": [ + " Of particular interest in this study was the effect of Ko and KR on the form of the cyclic volt ammo gram, since a small value for K; indicates strong adsorption of species i, while a large value indicates weak adsorption [see equations (6.141) and (6.142)]. For ease of presentation we identify the following adsorption cases of interest: W(O), W{R), S(O), S{R), W(O) - W(R), and W{O) - S{R), where Wand S designate weak and strong, respectively. The theoretical voltammograms for the first four cases have been derived by W opschall and Shain(48) and are presented in Figure 6.23. If the reactant or product is weakly adsorbed onto the surface, then the voltammograms generally exhibit an enhancement of peak currents in (al Weak adsorption of reactant, W(O). poQJo = 0.01. 1, Po = 5.0; 2, Po = 1.0; 3, Po = 0.1. The values for Po correspond to relative scan rates of 2500, 100, and 1. (b) Weak adsorption of product, W(R). PRQJR = 0.01. 1, PR = 20; 2, PR = 5.0; 3, PR = 0.1. The values for PR cor respond to relative scan rates of 40,000, 2,500, and 1. (c) Strong adsorption of reactant, S(O). (d) Strong adsorption of product, S(Rl. 222 Chap. 6 \u2022 Linear Potential Sweep. Cyclic Voltammetry 2 c 3 ____ s== 0'----'---'--L-......L--'.-----1 -2 0 log\u00a2o comparison with the uncomplicated charge transfer case. Thus, if the reactant is weakly adsorbed (Figure 6.23a) the reduction peak on the for ward scan is increased relative to the nonadsorbed case, since a greater amount of 0 is available at the surface. On reversal of the scan a greater amount of product R is also present at the interface (although it is not adsorbed), and hence the anodic peak is also enhanced by weak reactant adsorption. Note, however, that the enhancement of the anodic peak is less than that for the reduction wave since some of the excess R will have diffused into the bulk solution. In the case of weakly adsorbed product (Figure 6.23b) little enhancement of the reduction peak is observed. How ever, the reduction peak is shifted to more positive potentials with increasing scan rate, and a large enhancement of the anodic peak is predicted since more of R is retained at the surface than in the case where no adsorption is present. If the reactant (Figure 6.23c) or product (Figure 6.23d) is strongly adsorbed, then separate adsorption peaks (note that Ki is potential de pendent) may occur after or prior to the reduction peak on the forward scan, and vice versa for the reverse scan. The exact natures of the adsorption peaks are discussed further by Wopschall and Shain(4B). The potential sweep rate has a marked effect upon the form of the voltammograms for reversible charge transfer reactions involving adsorp tion (see Figure 6.24a, b). In the case of weak adsorption of the reactant the ratio of the reduction peak current to that for the uncomplicated charge transfer increases rapidly above a certain sweep rate (Figure 6" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000217_s0007-8506(07)62254-9-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000217_s0007-8506(07)62254-9-Figure1-1.png", + "caption": "Figure 1 : Prototypes of the bevel gear gauge", + "texts": [], + "surrounding_texts": [ + "The increasing complexity of measuring tasks to be carried out by coordinate measuring machines (CMMs) requires an intensification of the existing demands on the periodical supervision of the instruments. In particular that applies to measuring tasks at straight-toothed and helical cylindrical and bevel gears, whose functhmal surfaces cannot be taken by probing parallel to the coordinate axis - as it is done at prismatic and rotationally symmetrical workpieces - but exclusive have to be measured by three dimensional probing processes. Many of the modem coordinate measuring machines are not prepared for this measuring tasks, because among other difficulties e.g. the construction of the probing head is obsolete. Therefore a task specific inspection technique has been developed, which enables the user to obtain an objective and manufacturer-indepenht information about the quality of the used CMMs and software systems by using a task specific gauge [ I ] . In principle for the analysis of measuring deviations a highprecision master toothing could be used, which however neither can be manufactured nor calibrated with sufficient precision today. Annals of rhe ClRP Vol. 43/1/1994 For the task specific inspection of coordinate measuring machines a geometrical stable \"synthetic\" gauge is better suited, which can be built up with basic geometric elements (Fia. 1) 121. This gauge can be manufactured with conventional production processes in a high precision and can be calibrated reliable with known methods. The nominal 'points on the surface of the cylinders are generated by a special software unit, which as well analyses the measured points and generates a graphical evaluation of the measuring deviations. In adcition an objective analysis can be calculated: Comparing the detected geometrical differences with the limit values given by DIN and I S 0 standards the user of the bevel gear gauge can decide immediately, which quality of toothing can be measured with the tested CMM. 2. The Bevel Gear Gauge After finishing extensive investigations on the practical advantages and possibilities for manufacturing >f basic geometrical elements it was seen, that especially spheres and cylinders are suited for the use in a synthetic gauge. These elements can be manufactured with an acceptable effort in an enormous precision (deviation of shape <0.15 pm, surface roughness Ra < 0.02 pm and surface hardness 63-65 HRC), thus are available low-priced and guarantee small wear. The latest model of the bevel gear gauge is equipped with eight high-precision probing cylinders. Four of them are positioned at an angle of 30\" to the rotational axis of the master body, while the second grour nf four cylinders is attached at an angle of 15\" to the radial direction. As the probing cylinders are mounted so as to measure their shape on a roundness measuring machine with rotating spindle, all constructive requirements to the demands of calibration are fulfilled. The first officiaf calibration of the manufactured bevel gear gauge has been realised at the laboratories of the Physikalisch-Technische Bundesanstalt (PTB). On the surface of each probing cylinders more than 23.000 calibration points were recorded, whereas the relati c'e position of the probing cylinders to the master body 113s been taken in a 32-fold swing-round calibration (in analogy to the established four-position method of the calibration of a ball-plate) (Fig. 3). The overall uncertainty of calibration for each point on the surface of the probing cylinders, estimated for the measurement of a real bevel gear, amounts to approximately 2.0 pm, whereas the repetitive accuracy of measurements does not exceed 1.0 pm as can be proved. 3. Application of the Bevel Gear Gauge In order to investigate the capability of a coordinate measuring machine for the measurement of a bevel gear by means of the bevel gear gauge first of all the same operational steps have to be carried out as measuring a conventional bevel gear wheel or pinion. The gauge will be fixed in a suited manner and will be aligned at the grinded reference surfaces according to the type of machine in a manually or computer controlled way. Then a probe will be chosen, which is able to scan the surface of the nearest probing cylinder without collision. I .-I 1 . Cylinder 1 2. Cylinder 3. Cylinder W I Figure 3: Calibration of the relative position of the probing cylinders During the next working steps the nominal poims on the surface of the probing cylinders have to be generated and transmitted to the coordinate measuring machine under test. While the shape of real bevel gear flanks consist of three-dimensional free-form geometries, the shape of the bevel gear gauge to be probed is part of the area of a cylindrical surface. Therefore the PC-based program KEGNORM has been developed, which calculates the mathematical exact nominal points on the surface of the probing cylinders out of preconditioned base data of a bevel gear toothing. The flank geometry can be generated according to the calculation methods of Klingelnberg, Oerlikon or Gleason, where the spatial dimensions and the realised radius of curvature of the probing grids are similar to those on the surface of the corresponding real bevel gear flanks. The operator feeds the base data of the toothing into the computer, upon which the coarse geometry of the two flanks of a bevel gear tooth according to the kinematics of the gearing process is calculated. During the computing it is ensured, that all of the manufacturing characteristics of the gear cutting machine stay within their limits. After best fitting of a cylinder by Gaussian approximation, scaling of the probing grids and spatial shifting onto the real probing cylinders of the bevel gear gauge mathematical exact nominal points arc available, which can be transmitted online or by disk to the coordinate measuring machine under investigation e.g. in the form of a VDAFS-File (Fin.). 4. Evaluation of the actual Points with KEGNORIM After finishing the measuring process the actual points can be assessed by performing a conventional evaluation with the software of the manufacturer of the coordinate measuring machine. Nevertheless for a real judgement of the geometrical deviations an exclusive evaluation with best fit of nominal and actual points has to le done, which minimises errors of alignment and software included in the results. The external PC-based program system KEGNORM is prepared to perform such an independent evaluation. One condition for the application of this strategy is, that the used manufacturer-specific measuring program is able to store the probed actual points on disk in form of a VDAFS-File. After reading this file into the eialuation unit of KEGNORM, a best fit of nominal and actual points is performed in order to minimize the error components enclosed in the measuring data. Afterwards the geometrical deviations can be plotted relative to the nominal grid or in the length measuring uncertainty template in order to obtain a three-dimensional impression of the measuring uncertainty of the CMM." + ] + }, + { + "image_filename": "designv11_65_0001007_jsvi.1997.0998-Figure8-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001007_jsvi.1997.0998-Figure8-1.png", + "caption": "Figure 8. A belt system consisting of two loops connected by two pulleys.", + "texts": [ + " The stiffness matrix of the belt system is (k41 + k13)r2 1 0 \u2212k13r1r3 \u2212k41r3r4 0 (k62 + k23)r2 2 \u2212k23r2r3 0 \u2212k13r1r3 \u2212k23r2r3 (k13 + k34 + k23 + k35)r2 3 \u2212k34r3r4 K= G G G G G G G K k \u2212k41r4r1 0 \u2212k34r3r4 (k34 + k41)r2 4 0 0 \u2212k35r3r5 0 0 \u2212k62r6r2 0 0 0 0 0 \u2212k62r6r2 \u2212k35r3r5 0 0 0 G G G G G G G L l , (9) (k35 + k56)r2 5 \u2212k56r5r6 \u2212k34r3r4 (k56 + k62)r2 6 where the element with respect to the common pulley 3, (3, 3), is the summation of the elements of the stiffness matrix in each loop. The location of the elements of the stiffness matrices KL1 and KL2 in the stiffness matrix of the belt system are shown in Figure 7. As another example, consider the belt system shown in Figure 8. This belt system consists of two loops, L1 and L2, which are connected by the common pulleys 2 and 3. The equation of motion for the belt system is Iu +Cu +Ku=T, (10) where u= {u1, u2, . . . , u5}T, I=diag[I1, I2, . . . , I5] and T= {T1, T2, . . . , T5}T. The stiffness matrix of the belt system is (k51 + k12)r2 1 \u2212k12r1r2 0 \u2212k12r1r2 (k12 + k23 + k42 + k'23)r2 2 (\u2212k23 \u2212 k'23)r2 3 K=G G G G G K k 0 (\u2212k23 \u2212 k'23)r2 2 (k23 + k35 + k'23 + k34)r2 3 0 \u2212k42r4r2 \u2212k34r3r4 \u2212k51r5r1 0 \u2212k35r3r5 0 \u2212k51r5r1 \u2212k42r4r2 0 \u2212k34r3r4 \u2212k35r3r5 G G G G G G G L l , (11) (k34 + k42)r2 4 0 0 (k35 + k51)r2 5 where the elements with respect to the common pulleys 2 and 3, (2, 2), (2, 3), (3, 2) and (3,3), are the summation of the elements of the stiffness matrix in each loop" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000991_s0022-0728(96)05057-7-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000991_s0022-0728(96)05057-7-Figure1-1.png", + "caption": "Fig. 1. Cyclic voltammograms at a Pt disc electrode (diameter 5 mm) in CH.~CN +0.1 M TBAP of: (A) 2 mM [Mn(dmbpy)3]2+; (B) solution (A) after exhaustive oxidation at 0.72V; (C) 2mM [Mn(L)3] 2+ ((a) scan between 0 and - 2 . 3 V; (b) first scan between 0.05 and 1.2 V; (c) second scan between - 0 . 3 and 0.5V; sweep rate u = 100mVs-I) .", + "texts": [], + "surrounding_texts": [ + "\u2022 in poly[Mn(L)3] 2+ films prepared by complexation of Mn 2 + cations in polyLH + [ 11] obtained by electropolymerization of the pyrrole-substituted 2,2'-bipyridine ligand L. In order to have a better understanding of the electro-\nchemical behaviour of these films, we have firstly reinvestigated the conversion of the model complex [Mn(dmbpy)3] 2+ into [Mn202(dmbpy)4] 3+ (dmbpy = 4,4'- dimethyl-2,2'-bipyridine) in CH3CN solution [12]. The reversibility of this reaction was also studied.\n2. Experimental\n2.1. Electrochemistry\nAll electrochemical experiments were run under an argon atmosphere in a glove-box, using a standard threeelectrode electrochemical cell. All potentials are referred to an Agll0mM Ag + reference electrode in acetonitrile + tetra-n-butylammonium perchlorate (TBAP) electrolyte. The working electrode was either a platinum disc (0.2 cm 2) or a vitreous carbon disc (0.2 or 0.07 cm 2) polished with 1 Ixm diamond paste. Exhaustive electrolyses were carried out with a 5cm 2 platinum cylinder or a 10 x 10 \u00d7 4mm 3 carbon felt electrode (RCV 2000, 65 mgcm -3, from Le Carbone Lorraine). The electrolyte was a 0.1 M solution of supporting electrolyte in CH3CN (Rathburn, HPLC grade S) or dimethyl sulphoxide (DMSO) (SDS, pure on 4A molecular sieves). TBAP (recrystallized from dichloromethane + cyclohexane) was obtained from Fluka (puriss), dried under vacuum at 80\u00b0C before use, and stored under argon. Electrochemical measurements were carried out using an EG&G PAR model 173 potentiostat equipped with a model 179 digital coulometer and a model 175 programmer with output recorded on a Sefram TGM 164 X- Y recorder.\nThe apparent surface coverage F/molcm -2 of electroactive species immobilized were determined from charge under their cyclic voltammetric peaks, when the electrode is transferred into clean electrolyte.\n2.2. Spectroscopies\nElectronic absorption spectra were recorded on a Hewlett-Packard 8452A diode array spectrophotometer equipped with a Compaq 286 computer and a Citizen 120D printer.\n2.3. Ligands and complexes\nThe ligands 4,4'-dimethyl-2,2'-bipyridine (dmbpy) was purchased from Aldrich and 4-(4-pyrrol-l-yl-butyi)-4'methyl-2,2'-bipyridine (L) was prepared as described elsewhere [13]. The complex [Mn202(dmbpy)4] 3+ was synthesized following the procedure described for the corre-\nsponding 2,2'-bipyridine derivative [14] and characterized by EPR spectroscopy.\n2.3.1. [Mn(dmbpy) 3 I[BF 4 ]2 A solution of 0.245g of Mn(OOCCH3) 2 -4H20 (lmmol) in 12.5ml of H20 was added to 0.553g of dmbpy (3 mmol) in 12.5 ml of acetone. The resulting yellow solution was stirred for 15 min and then filtered. The complex was precipitated by addition of an aqueous saturated solution of NaBF 4, washed with water and precipitated twice from CH2CI 2 + diethyl ether and dried under vacuum (yield 80%).\n2.3.2. IMn(L) 3 IIBF412 This complex was prepared by the same procedure as above using L instead of dmbpy. Yield 60%.\n3. Results and discussion\n3.1. Electrochemical behaviour of [Mn(dmbpy) 3 ]e + in acetonitrile solution\nMorrison and Sawyer [12] have previously demonstrated that the electrochemical oxidation of tris(bipyridine)- and phenanthroline-substituted or not manganese(II) complexes leads to the corresponding di-/z-oxo dimanganese(IV,IV) complexes in CHaCN + 0.1 M TBAP.\nFor instance, the cyclic voltammogram of [Mn(dmbpy)3] 2+ in this solvent presents in the positive region a large irreversible peak at Epa = 1.07V corresponding to the metal oxidation process Mn(II) ~ Mn(IIl) (Fig. I(A)). The Mn(llI) species is unstable and reacts rapidly with residual water in the solvent via a disproportionation reaction to form the binuclear manganese(Ill,IV) species, which is oxidized to its (IV,IV) form as a consequence of its easier oxidizability than the starting complex. This subsequent chemical reaction is indicated on the reverse scan, as a weak irreversible cathodic peak detected at Ep, = -0.06 V, corresponding to the second step of the reduction of the di-/z-oxo binuclear species formed. A larger cathodic peak is observed at a more positive potential ( Ep~ -- 0.63 V). It does not correspond to the first reduction step of the di-/z-oxo since this latter is located at Ell 2 = 0.85 V (see below). This could be due to the reduction of a transient species such as a mono/z-oxo binuclear manganese complex. Eq. (1) summarizes the overall process:\n2[Mnn(dmbpy)3] 2+ + 2H20\nIV IV 4+ [Mn2' O2(dmbpy)4] + 4H \u00f7+ 4 e - + 2dmbpy\n(l)\nIn sweeps in the negative region, the regular three successive ligand-based reversible reductions at E~/2 =", + "- 1.77, - 1.95 and -2 .16V appear on the voltammogram [14] (Fig. l(B)).\nAs observed previously [12], an exhaustive electrolysis carried at a potential more positive than that of the anodic peak (E = 1.3 V) consumes two electrons per manganese and the binuclear complex is then obtained in its oxidized red form (IV,IV) ([MnEOE(dmbpy)4 ]4+ ). Controlled potential reduction at 0.5V furnished its green (III,IV) form ([Mn202(dmbpy)4] 3+) after 0.50 electron per manganese has been passed. It should be noted that the total yield of the transformation is only around 75%. Some undefined by-products are also formed.\nHowever, we found that the direct electrochemical generation of the binuclear complex in its mixed-valence state (Ill,IV) can be achieved if the oxidation potential is carefully controlled. As we will see in Section 3.2, this is a crucial point for the modification of the surface electrode by oxidation of [Mn(L)3] 2+. An exhaustive oxidation of\nthe solution conducted at the foot of the anodic peak (E = 0.72 V) consumes 1.5 electrons per manganese and furnishes with a 75% yield a green solution which presents electrochemical characteristics of [Mn202(dmbpy)4] 3+. The reversible one-electron wave at E~/2 = 0.85V (AEp = 0.06 V) corresponds to tht: (III,IV)/(IV,IV) redox couple, while the irreversible one (Ep~ = -0 .06 V) (Fig. I(B)) corresponds to the formation of the mononuclear complex [Mn(dmbpy)3 ]2+ via the (IH,III) transient species [15].\nThe oxidation of the [Mn(dmbpy)3] 2+ complex and the subsequent dimerization of the product species require release of two ligands and four protons to yield the di-p,-oxo dimanganese(lll,lV) complex following the reaction\n2[Mnn(dmbpy)3] 2+ + 2H20\n[Mnm'lVO2(dmbpy)4] 3+ + 4H + + 3e- + 2dmbpy\n(2) In acetonitrile solution, it has been suggested [12] that the free ligands released are protonated by the electrogencrated protons. We have confirmed this point by cyclic voitammetry analysis of the solution conducted between 0 and -2 .3 V (Fig. I(B)). The irreversible peak at -0 .8 V at a Pt electrode is typical of the reduction of dmbpyH \u00f7 into dmbpy [11] and this reduction is shifted to - 1.36 V at a C electrode.\nThe transformation of [Mn(dmbpy)3] 2\u00f7 to [Mn202(dmbpy)4] 3+ (Eq. (1)) is a reversible process, as attested by the observation of the three typical reversible redox systems of [Mn(dmbpy)3] 2+ on the cyclic voltammogram, past the reduction of the dmbpyH + peak. This is confirmed by a controlled-potential reduction at -0 .5 V which consumes three electrons per binuclear complex and returns to a solution of [Mn(dmbpy)3] 2\u00f7 with a 90% yield. The presence of free ligands in solution is necessary to ensure the completion of the reaction [ 12]. Indeed we have observed that a controlled-potential reduction at -0 .2 V of an authentic sample of the binuclear complex in CH3CN + 0.1 M TBAP furnishes only half the amount of [Mn(dmbpy)3] 2+ and consumes only one electron per binuclear complex. A brown side-product which is partially soluble is also formed and probably corresponds to manganese oxide. Eq. (3) summarizes the proposed overall\nprocess:\n[Mnm.'VO2(dmbpy)4] 3+ + e-\n[Mnn(dmbpy)3]2+ + Mn'VO2 + 2dmbpy (3)\nAn exhaustive reoxidation of this solution at 1.3 V followed by a reduction at 0.5 V restores almost entirely the initial amount of the binuclear /z-oxodimanganese(Ill,lV) complex. This demonstrates that the overall chemical process (Eq. (3)) is reversible. Furthermore, if the precipitate is removed by filtration before the electrolysis only 40% of [Mn202(dmbpy)4] 3+ is then restored by the same electrochemical sequence.", + "68 M.-N. Collomb Dunand-Sauthier, A. Deronzier / Journal of Electroanalytical Chemistry 428 (1997) 65-71\n3.2. Electrodes modified by electropolymerization of [Mn(L) 312 + in acetonitrile solution\n3.2.1. Redox behaviour of the monomer [Mn(L) 312 \u00f7 As expected, the electrochemical behaviour of [Mn(L)3] 2+ in CHaCN + 0.1 M TBAP is basically identical to that of the model complex [Mn(dmbpy)3] 2+. The voltammogram shows, in the negative region, the three reversible reduction waves (E~/2 = - 1.77, -1 .95 and - 2 . 1 5 V ) corresponding to the successive one-electron ligand-centred reductions (Fig. I(C), curve (a)).\nOn the positive scan, an intense irreversible peak at Ell 2 = 1.5V (not shown here) flanked by a shoulder at 0.9 V is observed, depicting the oxidation of the metal centre of the complex M n ( l l ) ~ Mn(lll) in [Mn(L)3] 2+, that of the pyrrole group and the overoxidation of the resulting polypyrrole film. If the anodic potential is restricted to 1.2V, the electrochemical response of the polypyrrole on the reverse scan, is then detected around 0.3 V (Fig. I(C), curves (b) and (c)).\nFunctionalized polypyrrole-modified electrodes can be\neasily obtained on Pt or C surfaces by repeated cyclic voltammetfic scans between - 0 . 2 V and 0.75 to 0.8V. This is depicted in Fig. 2(A), where is shown the steady growth of the typical quasi-reversible peak system of the polypyrrole around 0.3 V upon scanning over this potential domain. The growth of the film is regular, showing an excellent efficiency of the electropolymerization process in contrast to what is observed for electropolymerization of [Mn202(L)4] 3+ monomer. After the transfer of this electrode to pure electrolyte, the polypyrrole response persists and is stable if the potential is restricted to the - 0 . 2 to 0.5 V range (Fig. 2(B), curve (a)). Scanning up to 1 V induces the appearance of a second quasi-reversible system located at El~ 2 = 0.86V (Fig. 2(B), curves (b) and (c)). The shape of the voltammogram is similar to that obtained from electropolymerization of [Mn202(L)4] 3+ [10] and the half-wave potential is closed to that of the model complex [Mn202(dmbpy)4] 3+ in bulk solution (El~ 2 =0 .85V; (HI,IV)/(IV,IV) couple). These features demonstrate that the complex immobilized in this polypyrrole film is in its di-/.t-oxo form. The dimerization of [Mn(L)3] 2+ to give [Mn202(L)4] 3+ occurs probably prior to electropolymerization since this process is efficient from 0.7 V for the model complex [Mn(dmbpy)3] 2+ in solution. However, additional dimerization occurring from some preformed films of poly[Mn(L) 312 + cannot be rejected. This transformarion can be mediated in the film by the oxidized polypyrrole during the first oxidation scan. Similar modified electrodes can be obtained by controlled-potential oxidation at 0.7V. Thin films up to FMn2O 2 = 1.2 \u00d7 10 -8 molcm -2 were easily obtained from this potentiostatic procedure.\nCycling this Ptlpoly[Mn202(L)4] 3+ film several times\nbetween 0 and 1 V induces both a decrease of the (III,IV)/(IV,IV) and polypyrrole redox peak systems. After 70 successive cycles the electroactivity of the modified electrode vanishes. At this stage, a polymer film remains at the electrode surface. The modification of the electroactivity of the film is due to the overoxidation of the polypyrrole framework. Protons liberated as a consequence of this overoxidation [16], interact with the [Mn202(L)4] 3+ core leading to protonation, demetallation and disproportionation phenomena [10].\nAs observed in homogeneous solution for [Mn202(dmbpy)4] 3+, the reduction of the immobilized binuclear [Mn202(L)4] 3+ sites leads to the mononuclear" + ] + }, + { + "image_filename": "designv11_65_0001881_0032-3861(84)90380-x-Figure4-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001881_0032-3861(84)90380-x-Figure4-1.png", + "caption": "Figure 4 Cyclic voltammograms for the reduction of oxygen molecules by the graphite electrode coated with polyviologens in oxygen saturated 0.2 M phosphate buffer solution (pH 7.0) at 25\"C. Scan speed was 50 mV s-1. Amount of polyviologens were 5.1 x l 0 -8 unit molcm -2", + "texts": [ + " Poly(viologen)s usually became insoluble in aqueous perchlorate solutions and therefore the impervious films formed become inaccessable to ferricyanide penetration and perchlorate migration. The Koutecky-Levich plots for i t were non-linear both in KCI and NaCIO4 solutions. This could also be caused by the slow electron-propagation through the polymer layers. As for the PCV/Fe(CN)6 3 - system, the results were very similar to those obtained for the PEV/Fe(CN)6 3 - system. Heterogeneous cross-reaction with molecular oxygen PXV + has been known to be one of the most potent catalysts for molecular oxygen reduction 29. Figure 4 illustrates the cyclic voltammograms recorded at poly(viologen)s coated electrodes in oxygen saturated 0.2 M phosphate buffer solution (pH 7.0). The much enhanced cathodic peak currents, the absence of return (anodic) waves and with peak potentials being more positive than those measured in oxygen free solution represent the rapid heterogeneous cross-reaction. From the slopes of the Tafel plots for the reduction of molecular oxygen, 85, 79 and 68 mV/decade were obtained for PXV-PSS, PEV and PCV coated electrodes, respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003725_s1474-6670(17)31398-8-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003725_s1474-6670(17)31398-8-Figure1-1.png", + "caption": "Fig. 1. Kinematic chain model of a limb.", + "texts": [ + " Once again we obtain a linear relation between the output y and the new input v and thus previous results can be extended. The only difference is in the additional input Uo that can be used to stabilize the zero dynamics. This second approach is strictly related to (Khatib, 1987) where At = JT(JM-1JT)-1 and where the reader can also find details on how to choose Uo that stabilizes the zero dynamics. u = A(x)t[v - b(x)] + [Im - A(x)tA(x)]uo. With the intent of imitating human reaching motion, we considered a 2DOF (two degrees of freedom) model of a limb (see Fig. 1) as in (Morasso, 1981) and (Giszter et al., 1993). The dynamics of this model can be expressed in the form (4) with p = m = 2. Let the input be the torques applied at the joints, u = [Ul U2], and the output be the cartesian position of the extremity P, i.e. y = [xp yp]T. The input to output feedback linearization of this model leads Case 2: p < m. A first possibility consists in reducing back to the case p = m adding a secondary output y' such that [yT, y'T]T E lRm . Details on how to choose y' so as to minimize a given cost can be found in (Samson C" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002100_pime_proc_1986_200_139_02-Figure2-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002100_pime_proc_1986_200_139_02-Figure2-1.png", + "caption": "Fig. 2 Ball contact geometries and velocities", + "texts": [ + " = - ' z sin ai F A Po = - z sin a, ( 3 ) (4) where 6* is obtained from standard tables [for example Harris (6)] - (D/2)cf; - 1 + 6,) cos a, I (D/2Ui - 1 ) + 6i ai = cos-' Equations ( 1 ) to (6) are solved by assuming an initial contact angle and working through the equations in the order that they are presented above. The value of ai obtained from equation (6) is then used to start a new iteration. The process is repeated until convergence is achieved. 2.2 Cage speed Consider the geometry shown in Fig. 2: d m D ~i = - - - cos ai = 3(dm - D cos ai) 2 2 ro = $d, + D cos a,) v . = r . w. 1 1 1 u, = row, %ui + v,) = %ri wi + r, w,) = %ro + ri)w, Therefore, 0, = ri 0, + r, w, r, + ri (7) 0 IMechE 1986 at Purdue University on March 13, 2015pic.sagepub.comDownloaded from A1 Original inner race centre A2 Deflected inner race centre B Fixed centre of outer race C Deflected ball centre or (1 3) wi(d, - D cos ai) + w0(d, + D cos a,) w, = (dmD/2)(cos a. - cos ai) The epicyclic ratio is always defined with respect to a stationary outer race; thus, (14) w -6" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003674_j.mechmachtheory.2004.05.002-Figure5-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003674_j.mechmachtheory.2004.05.002-Figure5-1.png", + "caption": "Fig. 5. Rigid body displacement associated with a revolute joint.", + "texts": [ + " It can be noted that such a partial derivative has been obtained using physical and geometrical reasoning. Moreover, this was formulated in terms of the Lie product of the involved screws. Thus, a systematic computational procedure can be devised by directly utilizing easily understood vector quantities. Simultaneously, the vector nature also provides direct insight into the physical interpretation of the Lie product [18]. In this case, the screw axis coincides with the motion axis of a revolute joint. A geometrical description of the case under study is shown in Fig. 5. Thus, for this particular case, the screws associated to the corresponding motion axis, the reference configuration and the current configuration are, respectively, as follows: $ \u00bc e r e ; $i \u00bc ei ri ei \u00fe pei ; $j \u00bc ej rj ej \u00fe pej \u00f029\u00de Then, resorting to the definition of the Lie product, Eq. (5), and for the case under study, we obtain that $ $i\u00bd \u00bc e ei e \u00f0ri ei\u00de ei \u00f0r e\u00de \u00fe e \u00f0pei\u00de \u00f030\u00de On the other hand, considering the geometry shown in Fig. 5, we can observe that the translational vector parallel to the screw axis, dk does not exist for this case, i.e., dk \u00bc 0. Then, we have that ej \u00bc Qei \u00f031\u00de rj ej \u00bc r Qei \u00feQ\u00f0p ei\u00de \u00f032\u00de pej \u00bc p\u00f0Qei\u00de \u00f033\u00de In analogy with the procedure presented in Section 4.1, the corresponding partial derivatives of the above expressions take now the following form: oej o/ \u00bc e ei \u00f034\u00de o\u00f0rj ej\u00de o/ \u00bc e \u00f0ri ei\u00de ei \u00f0r e\u00de \u00f035\u00de o\u00f0pej\u00de o/ \u00bc e \u00f0pei\u00de \u00f036\u00de Thus, the partial derivative of screw $j can be arranged as follows: o$j o/ oej=o/ o\u00f0rj ej\u00de=o/ \u00fe o\u00f0pej\u00de=o/ \u00bc e ei e \u00f0ri ei\u00de ei \u00f0r e\u00de \u00fe e \u00f0pei\u00de \u00f037\u00de Analyzing the foregoing expression, it can be noted that Eq" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001437_bfb0061422-Figure8-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001437_bfb0061422-Figure8-1.png", + "caption": "Fig. 8a Fig. 8b", + "texts": [ + " Let us suppose that (Mi,@i) has been defined and proceed inductively to define (Mi+l,@i+l) in one of the three following ways: (1) If #i has a trajectory, say y, connecting two different multisaddles points p and q, we obtain (Mi+l,~i+l) by blowing down \u00a5 to a point; that is, we define Mi+ 1 = = Mi/[p } U {q] U \u00a5 and a flow @i+l on Mi+ 1 in such a way thai the trajectories of @i+l coincide with those of \u00a2i in the set M i - [[P} U {q} U Y] , and moreover [p] U [q} U Y is a fixed point of ~i+l\" See Fig. 7a before blowing down y and Fig. 7b after blowing down y. 325 (2) If \u00a2i has no trajectories connecting two distinct multisaddle points but it has two source-fixed-points, say p and p' , which are connected to the same multisaddle p~int q by means of trajectories 8 and @' respectively, and moreover q has a separatrix which is adjaeent to both 8 and @' (see Fig. 8a). Then, we make M i = Mi+ I and define @i+l on Mi+ I in such a way that: (A) The trajectories of \u00a2i+l coincide with those of @i in M i - [k U W u\u00a2i(p) U wU@i (q)} (wU@i (p) denote the unstable manifold of p for the flow @i ) and (B) @i+l has a source-fixed-point r connected to q be a unique trajectory and W u (r) = k D @i+l U W u (p) U W u ~i(q ) (See Figs. 8a and 8b). \u00a2i 326 (3) If @i does not satisfy the conditions required in (i) or (2) above to define (Mi+l, ~i+l) , but it has a (non-fixed point) trajectory, say y, such that both ~(y) and w(Y) are the same multisaddle point q" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003585_s0219878904000069-Figure7-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003585_s0219878904000069-Figure7-1.png", + "caption": "Fig. 7. Experimental setup and object.", + "texts": [ + " , h, h \u2265 2), there are Qu = s \u2208 R3h, h \u2265 2, (51) Q = [ [\u2212T h01I3]T \u00b7 \u00b7 \u00b7 [\u2212T h0hI3]T ]T \u2208 R3h\u00d74, (52) u = [ kg P T g ]T \u2208 R4\u00d71, (53) s = [ pT h1 \u00b7 \u00b7 \u00b7 pT hh ]T \u2208 R3j\u00d71. (54) When h \u2265 2 and rankQ = 4, the optimal approximate solution of u, which contains the object mass and the position of center of mass, can be obtained from Eq. (51) since 3h > 4, u = Q+s, (55) where Q+ \u2208 R4\u00d73h is the pseudoinverse of Q. From the obtained u, we can obtain the center position P g of mass with respect to frame O \u2212 XY Z, and the object mass m m = 1/kg. (56) The experiments were performed using a robot arm PUMA260 shown in outline by Fig. 7(a). The unknown object is a plastic frustum of a cone shown in Fig. 7(b), where a weight is added in the frustum. The experimental environment is an unknown plane and there is sufficient friction to prevent object slipping. In t. J. I nf . A cq ui si tio n 20 04 .0 1: 47 -5 5. D ow nl oa de d fr om w w w .w or ld sc ie nt if ic .c om by N A N Y A N G T E C H N O L O G IC A L U N IV E R SI T Y o n 04 /2 7/ 15 . F or p er so na l u se o nl y. Table 1. Results of experiment. pgu [cm] pgv [cm] pgw [cm] m [kg] Estimated Value \u221210.46 \u22120.28 8.21 0.746 Reference Value \u221210" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003067_iros.1998.724635-Figure7-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003067_iros.1998.724635-Figure7-1.png", + "caption": "Fig. 7: Several configurations of multiple objects", + "texts": [ + " , c j ) and fcot (t = l , . . . , ~ ) denote the contact forces applied by the kth contact of finger j and applied at the tth contact between two objects, respectively, and m g denotes the gravity vector. When rankDz. = rank[Dz m], fc satisfying the equilibrium condition( f = 0 ) exists. Especially, when rankD: < cy=l 3cj + 3r, fc is not uniquely determined. Here, we define an index as follows: n Id = 3cj + 3r - rankDz, (20) j = 1 where Id denotes the degrees of freedom of the internal force. Fig.7 shows four grasp configurations, where the direction of gravity is shown by an arrow, and each object is a cylindrical object whose radius and weight are unity. We assume that the coefficient of static friction at each contact point is large enough to avoid any slip. (a) and (b) show examples for two objects grasped by a two-fingered hand. (c) and (d) show examples for three objects grasped by twofingered hand and by three-fingered hand, respectively. dimDB, rankDB and rank[Dg m] for each grasp configuration are shown in Table 1" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001984_jsvi.2001.3643-Figure3-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001984_jsvi.2001.3643-Figure3-1.png", + "caption": "Figure 3. (a) Slotted isotropic shaft with inertia slots: (b) front view; (c) side view.", + "texts": [ + " sequences and materials have also been considered in the present study. The results are compared with those of the isotropic slotted shaft. The present analysis of slotted shaft has been done using FEMAP 97 as pre- and post-processor and NASTRAN as the solver. The slotted steel shaft as shown in Figure 1 is analyzed using tetrahedral element, CTETRA (Figure 2) with 10 nodes. The PSOLID and MAT1 cards are used to input the element and the material properties respectively. In NASTRAN, eigenvalue analysis is performed using SOL3 and SOL103. Figure 3(a}c) shows the solid model of an isotropic slotted shaft. The slots in the shaft are generated by using bullion operations. Six longitudinal slots are considered for this model with 303 angle between the two adjacent slots. Repeating bullion operations with circular disk generates inertia slots, which are also shown in Figure 3(a}c). Four numbers of inertia slots (compensatory) are considered in this model. As shown in Figure 4, the rotor has been discretized by using 1600 ten-noded tetrahedral elements. Simply supported boundary conditions are assumed for the shaft ends. Composite materials have many advantages, which made them applicable to a wide variety of products and attempts have been made to replace isotropic shafts by composite shafts. There have been very few works reported on composite shafts even though, Zinberg and Symmonds [10] studied these way back", + " By split eigenfrequencies is meant, throughout this paper, the eigenfrequencies in #exure associated with the direction of maximum and minimum TABLE 2 Comparison of NAS\u00b9RAN1s solution and previous results Properties of di+erent composite materials E E G \"G G Material (GPa) (GPa) (GPa) (GPa) (kg/m ) Boron/epoxy 211 24)1 0)36 6)9 6)9 1967 Graphite/epoxy 139 11 0)313 6)05 3)78 1578 Carbon/epoxy 130 10 0)25 7)0 7)0 1500 rigidities. It can be seen from Figure 8 that the in#uence of slot depth to increase the regime between the split eigenfrequencies is more compared to that of the width. 4.1.2. Inertia slots Four inertia slots are created on pole face as explained in section 2 (see Figure 3), there by reducing the bending sti!ness in the x}x plane to a value closer to that for the y}y plane. The results with the inertia slots are shown in Table 1. From Table 1 it is clear that due to these slots the rotor is made almost symmetric. For the purpose of comparing the results with the previous ones, a composite hollow shaft of the following data [12] was considered: length of the shaft\"2)47 m, radius of the rotor\"0)0635 m, thickness\"0)132E-3 m, density\"1965 kg/m , E1\"211 GPa, E2\"24)1 GPa, G \"6)9 GPa, \"0)36, bearing sti" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000776_s0957-4158(97)00027-5-Figure11-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000776_s0957-4158(97)00027-5-Figure11-1.png", + "caption": "Fig. 11. Schematic diagram of 3-1ink robot used in numerical experiment 2.", + "texts": [ + " The Time delay control of nonlinear systems with neural network modeling 625 control result using this filter with P = - 1 is shown in Fig. 9. The result shows that the fluctuation is suppressed and the control error is also within reasonable range. 5.2. Numerical experiment 2 (inverse kinematics of 3-link robots) In this experiment, the control of robot end effector position through inverse kinematics are presented. The robots (ll = 1, 12 = 1, 13 = 0.5) and the related coordinate systems are shown in Fig. 11. The forward kinematics of the robot are trained with the following conditions. \u2022 Three input nodes (joint angles) in the input layer, 20 nodes (tangent hyperbolic function as an activation function) in the single hidden layer and three linear output nodes (cartesian coordinates) in the output layer. \u2022 30 000 individual learnings with randomly generated samples in the working space. \u2022 Simple error backpropagation learning with learning rate = 0.02. The problem is to track the semi-circle (0~ = - 1 / 2 n to 1/2n, 02 = n/4, 03 = n/4) without considering the robot dynamics" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003674_j.mechmachtheory.2004.05.002-Figure2-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003674_j.mechmachtheory.2004.05.002-Figure2-1.png", + "caption": "Fig. 2. Graphical representation of the displacement associated with a helical joint.", + "texts": [ + " It should be noted that, for a general displacement of a rigid body, the screw axis does not necessarily pass through the origin O of the fixed frame, i.e., in general, krQk 6\u00bc 0. From the geometry shown in Fig. 1, we can formulate the following equations: e \u00bc Qe0; r \u00bc rQ \u00feQ\u00f0r0 rQ\u00de \u00fe dk \u00f01\u00de where Q is a proper orthogonal matrix [25] which represents a rotation about axis eQ. Moreover, it should be noted that the translation vector d of an arbitrarily selected point, fixed on the body, can be conveniently decomposed [24] into vectors parallel dk and perpendicular d? to the rotation axis eQ, as shown in Figs. 1 and 2. In fact, Fig. 2 may be considered as a particular representation of the displacement corresponding to a helical joint because of it was assumed a screw displacement, and, for that reason, vector dk depends on the rotation angle /. By using Eq. (1), we can now state the following result: r e \u00bc rQ Qe0 \u00feQ\u00f0r0 rQ\u00de Qe0 \u00fe dk Qe0 \u00f02\u00de The above equation can be simplified by noting that the cross product of two rotated vectors is identical to the rotated cross product, thereby obtaining r e \u00bc Q \u00f0r0 rQ\u00de e0 \u00fe \u00f0rQ \u00fe dk\u00de Qe0 \u00f03\u00de Thus, Eqs" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001810_0094-114x(84)90051-x-FigureI-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001810_0094-114x(84)90051-x-FigureI-1.png", + "caption": "Fig. I. Chain load distribution test machine.", + "texts": [], + "surrounding_texts": [ + "This paper presents an experimental and analytical evaluation of roller chain load distribution. No known experimental studies have been performed to determine the load distribution in a sprocketchain drive. By conducting these experiments, it was possible to determine the performance of steel chains on steel sprockets and plot chain load distribution curves for the steel sprocket drive under a variety of loading conditions. Other factors that affect the load distribution were investigated, such as lubrication, misalignment, sprocket speed of rotation and slack-side load. Because of the equal importance of the sprocket when it functions either as a driver or a driven sprocket, this paper includes the experimental results of the sprocket load distribution for both the driver and the driven sprocket, depending on the sprocket direction of rotation with respect to the load. The results show that there is a substantial difference that no known analysis has focused on until now. In the sprocket chain drive, each tooth of the sprocket is subjected to a certain percentage of the total transmitted load which in turn is proportional to the total transmitted torque. The distribution of the tooth load and also the chain tension depend on the sprocket size, the number of sprocket teeth in contact, the pressure angle and the elastic properties of both the chain and sprocket materials [1]. Binder [1] gives fundamental force relationships that apply to chain joints and link plates, with and without centrifugal force due to rotation. Binder presents a geometric progression load distribution that is applied when centrifugal forces are neglected. Koyama et al. [2-5] discuss the toothed belt drive system, which is similar to the roller chain drive. In a second report [2], Koyama discusses the effect of pitch difference on load distribution. In a third report [3], Koyama presents an experimentally derived relationship between the tight-side tension, the belt life and the fracture features of toothed belts. In a fourth report [4], Koyama gives the load distribution in case of incomplete meshing, and in a fifth report [5], Koyama gives the effect of pitch difference on the fatigue strength of toothed belts. TEST CHAINS AND SPROCKETS One-inch, double-pitch, number 2040 riveted steel roller chain was selected for the experimental and analytical investigation of the roller chain load distribution. The chain length used in conducting these experiments was 1.5 m. In selecting the chain, there were two main requirements: (1) the chain pin link plate had to be large enough to accommodate a strain gage used to measure the chain load distribution; (2) the chain had to give a reasonable extension under a moderate load. Size 2040B20, ASA standard steel sprockets (1.0 in. pitch; 3.236 in. pitch diameter; and 3.46 in. outside diameter) were used to represent the high modulus of elasticity class of sprocket materials. The steel sprockets were machined from C1144 carbon steel bar stock. The sprocket bore was cut to 1.25 -+ 0.005 in. to fit a mounting shaft of a load distribution test apparatus. Also, a keyway was cut to secure the sprocket on the mounting shaft." + ] + }, + { + "image_filename": "designv11_65_0002201_iemdc.2001.939386-Figure15-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002201_iemdc.2001.939386-Figure15-1.png", + "caption": "Fig. 15. Flow in the Stator Cooling Gaps With the Modifications to the Stator Bars and the Gap Supports", + "texts": [ + " (U32 - U Z 2 ) (10) And for the increase of static pressures follows: into the rotor-stator gap through the slot on the leeward side of the stator bar. This situation is shown in Fig. 14. With the modifications to the stator bar and to the gap supports this reingestion is suppressed. The cooling air is leaving the rotor-stator gap uniformly through both slots 1 1 2 1 P 3 - P 2 ' = - 'P' (W2 - U z 2 ) i- - ' P . C a x 2 - 5 - P - C T ~ ~ (11) 2 2 6 of 8 on the windward and on the leeward side of the stator bars. The flow stays attached to the channel walls (gap supports and wedges) all the way into the plenum. The improved flow is shown in Fig. 15. VI. PROTOTYPE DESCRIPTION Pictures of the modifications applied to the prototype engine are shown in Fig. 16. They are virtually identical to the modifications applied to the experimental test section. A motor of the same type but without any of the modifications was tested as a reference. VII. PROTOTYPE R SULTS Both motors were tested under the same operating conditions. Table I lists the physical parameters recorded under steady-state conditions. The temperature increase of the stator bars with respect to the cooling inlet temperature is reduced from 42" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002780_1521-4109(200206)14:11<741::aid-elan741>3.0.co;2-u-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002780_1521-4109(200206)14:11<741::aid-elan741>3.0.co;2-u-Figure1-1.png", + "caption": "Fig. 1. Schematic representation of front (A) and overhead (B) view of the voltammetric detector of tubular configuration: a, electric cable; b, rectangular silver plate; c, fragment of the master pellet; d, Perspex holder; e, non-conductive epoxy resin; f, contact with the reference electrode; Ew, working electrode; Eaux, auxiliary electrode; Eref, reference electrode.", + "texts": [ + " Graphite-paraffin pellets were prepared by dissolving 0.25 g of paraffin wax in 10 mL of warm n-hexane (40 C) in a beaker placed in a water-bath and adding 4.75 g of graphite powder with stirring. After complete evaporation of the organic solvent, 0.20 g of the dry graphite powder, now containing 5% of paraffin wax (% w/w), was pressed with a 10.0 mmdiameter pellet press at 19000 kg cm 2 for 5 min. In this way, a disk with a 10.0 mm diameter and 1.2 mm thick was obtained. Electrical contact was made through an electrical cable (Fig. 1A,a) that was attached with a electric solder to a small rectangular silver plate (1.0 3.0 mm) (Fig. 1A,b) to which a square-shaped fragmentof thegraphite-paraffinpellet (Fig. 1- A,c) was glued with a conductive silver based epoxy resin. Following this, the fragment of the pellet was housed in a Perspex holder (Fig. 1A,d) with a small cavity filled with a non-conductive epoxy resin (Fig. 1A,e). Finally, after hardening, a channel of 0.8 mm diameter was drilled perpendicular to the opposite sides of the housing, through the center of the Perspex holder. The tubular voltammetric detection system (Fig. 1) was made up of a working (Fig. 1, Ew) and an auxiliary electrode Electroanalysis 2002, 14, No. 11 (Fig. 1, Eaux), both carbon composites, of tubular configuration and constructed following the process previously described. The reference electrode (Fig. 1, Eref), an Orion 90-02-00 double junction AgCl/Ag electrode (inner filling solution, Orion 90-00-02; outer filling solution 0.1 MKNO3), was placed in a previously described support [23]. The connection between the reference electrode and the Perspex structure that housed the other two electrodes (Fig. 1,f) was established through Teflon tubing (0.8 mm internal diameter). The tubular voltammetric detection system was connected to the sliding part of the injector-commutator bymeans of Teflon tubes (0.8 mm i.d.) with the shortest possible length. The total inner volume of the movable detector was determined as being 100 L [25]. All commercial fruit juiceswere also analyzed by theAOAC method, which is based on the reduction of 2,6-dichlorophenolindophenol by ascorbic acid. The method involves titration of acidified sample with the sodium salt of a blue dye (2,6-dichlorophenolindophenol), the endpoint being indicated by appearance of the pink acid form of the dye" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000019_978-1-4684-7767-2_2-Figure2-2-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000019_978-1-4684-7767-2_2-Figure2-2-1.png", + "caption": "Figure 2-2 Three possible models for multilayered structures. This work uses the axisymmetric model.", + "texts": [ + " (2-1) (2-2) (2-3) where (J is the stress, E is the elastic modulus, ex is the coefficient of linear thermal expansion, v is Poisson's ratio and I1T is the temperature change from the temperature where the stresses are zero. (This may be the assembly temperature, or, if relaxation has occurred, it may be near room tempera ture.) Now if the plate is in the x-y plane, we have (Jz equal to zero and this ANALYSIS 81 gives us (2-4) (2-5) (2-6) There are three special cases of some interest here as shown in Fig. 2-2. First, the \"strips.\" A strip is much thinner in the z-direction than in the x and y directions, but its x-dimension is much larger than its y-dimension. In this case, we can assume that (J y is much less than (J)C. This leads to (2-7) This is the equation used for most of the bimetallic strip analyses, including the classic by Timoshenko. 1 Second, if the bending is all in one direction (cylindrical bending), then the only strain in the y-direction is thermal strain, and we can see that (Jy equals V(Jx" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002946_abb-2003-9693532-Figure12-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002946_abb-2003-9693532-Figure12-1.png", + "caption": "Figure 12 Coordinate system.", + "texts": [ + " However, when excited or frightened, apes can plunge through the forest canopy at astonishing speeds, sometimes covering 30 feet (~ 9.5 m) 0 time Trajectory ydi(t) tm stm-1 s tm+1 s b (t) m k y (t) m k y (t) m k ~ yd (t) m i b (t) m k b (t) m ky (t) m k ~ - rm k b (t) m ky (t) m k ~ -( ( Figure 11 Adaptation of desired trajectory. Applied Bionics and Biomechanics 2003:1(1)62 or more in a single jump without a break in \u2018stride\u2019 (\u2018fast brachiation\u2019, ricocheting) (Eimerl and DeVore 1966). In this section, we adapt the proposed learning algorithm to two types of slow brachiation: over-hand brachiation and sidehand brachiation. Figure 12 shows the coordinate system. Motion measurement using real-time tracking system A vision sensor is very useful to measure a dynamical motion without constant constraints, because the constrained points are switched in accordance with the body posture. During brachiation, it is almost impossible to measure the body position, eg the tip of the free arm or the centre of gravity of the robot, because the slip angle at the catching grip is not directly measurable using a potentiometer or rotary encoder" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001696_s0263-8223(01)00168-4-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001696_s0263-8223(01)00168-4-Figure1-1.png", + "caption": "Fig. 1. Geometry of composite circular spring with parallel flat contact surfaces.", + "texts": [ + " A three-dimensional finite element model taking into consideration of transverse shear deformation effects has also been employed to investigate the mechanical characteristics of the candidate springs. The analytical and numerical results are compared with experimental data. The relationships between the spring stiffnesses of woven composite springs with extended flat contact surfaces under different loading configurations and geometry are also illustrated. A schematic diagram of a mid-surface symmetric laminated thin circular spring with parallel flat contact surfaces is shown in Fig. 1. The spring consists of two circular portions with mean radius \u2018\u2018R\u2019\u2019 and two flat surfaces on the top and bottom with length \u2018\u2018w\u2019\u2019. \u2018\u2018a\u2019\u2019 is the semi-included angle. The spring has overall nominal thickness \u2018\u2018t\u2019\u2019 and width \u2018\u2018L\u2019\u2019. The load acts along the centre line Ct\u2013Ct of the flat contact surface on the top while the centre line Cb\u2013Cb of the flat contact surface at the bottom is rigidly clamped for the line-loading condition. In surface-loading, the applied load is uniformly distributed on the top and bottom flat surfaces" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002845_tpas.1981.316842-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002845_tpas.1981.316842-Figure1-1.png", + "caption": "Figure 1. Two-Winding Stepped Air Gap Induction Motor", + "texts": [ + " 80 SM 523-1 A paper recommended and approved by the IEEE Rotating Machinery Committee of the IEEE Power Engineering Society for presentation at the IEEE PES Summer Meeting, Minneapolis, Minnesota, July 13-18, 1980. Manuscript submitted January 22, 1980; made available for printing April 21, 1980. This paper will expand upon that shaded-pole miotor model by adding the effect of a stepped air gap. One will see that it is still possible to formulate a simple equivalent circuit fromthe resultingequations. The coupling between the main and shading windings is, of course, more complex. DESCRIPTION OF MACHINE Figure 1 shows a diagram of the machine to be ana- lyzed. The stator has a main winding located at = 0 and an auxiliary winding located at e = -a. The stator provides a nonuniform air gap whose smallest portion centers at 0 = -S. The rotor cage breaks into two windings, x and y, in quadrature with each other. The self and mutual inductances of the windings are not easily determined by inspection but can be readily calculated through the'use of winding functions [16]. According to the derivation of Appendix A, one may calculate the mutual inductance between two windings, denoted 1 and 2, by the formula (1)L1 2 = p0 rQ f h(X)N1(X)N2(X) dX 0 where Nl,N2= winding functions of windings x = displacement angle along periphery motor h(X) = inverse air gap function r = radius of air gap = stack length Assuming that a sinusoidal analysis is valid for the motor, the winding functions for the stator and rotor are given by N (X) = N cos X N (X) =a N cos(X + c) 'sa N (X) = N cos(X 0) (2) N (X) = N sin(X 0) y r where Ns and Nr are the peak values of the main and rotor winding functions, respectively, and a is the stator auxiliary-to-main turns ratio" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002426_841086-Figure7-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002426_841086-Figure7-1.png", + "caption": "Figure 7 Pressure distribution for asperity contact.", + "texts": [ + "5 x 10 cycles, these cracks are typically 2-10jjm in length and are inclined at an angle less than 90\u00b0 with the crack tip pointing in the direction opposite to the rolling direction. This section summarizes the results in developing an analytical model for predicting the initiation of these surface cracks. In developing the crack initiation model, it is assumed that the material close to the surface is subjected to a combined stress field induced by a contact between two circular cylinders and by a spherical asperity tip within the line contact (Fig. 7). The amplitude of the average shear stress and the maximum average normal stress over a semi-circular area of a radius c inclined at an angle a from the surface are considered to be the main driving forces initi ating a surface crack. These average stresses are found to be a function of the inclination angle a, the frictional coefficients for the b. The qualitative stress behavior falls into two categories: (1) for low friction (f ~ 0.1) the points of maximum shear stress and maximum octahedral shear stress are located beneath the surface, and (2) for high friction (f > 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000005_jiee-1.1922.0066-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000005_jiee-1.1922.0066-Figure1-1.png", + "caption": "FIG. 1.\u2014Brush which tends to come out of contact under the influence of current.", + "texts": [ + " It must not be assumed from the film that the switches which apparently fail were in all cases inadequate to withstand a test equal to their rating. The numerical figures of the tests must be studied in order to arrive at a conclusion in this respect, as in some cases the testing currents were increased above those corresponding to the breaking capacity rating in order to compare different types of construction. It was found that practically all the circuit I reakers tested had brush contacts arranged as in Fig. 1, so that when the current flowed the resultant mechanical force acted in a direction opposed to the brush pressure, thus tending to open the contacts and cause them to burn and weld together. This can be understood by reference to the principle that a closed electric circuit always tends to open out and enclose the maximum area. As a result of the experiments an improved arrangement of the brushes was introduced and is illustrated in Fig. 2, from which it will be seen that the force set up by the current increases the brush pressure", + " Spurr : The experiments shown on these films are doubly interesting as they represent types of switches in common use in this country, and, although the tests resulted in the almost complete destruction of the switches, we must not forget that we are to-day operating switches that may be called upon at any moment to operate under very similar conditions for the first few periods when a fault occurs. Owing to the necessity of a public supply undertaking having to take immediate steps to resume the supply when an interruption occurs, it often happens that time and circumstances will not always permit the damaged apparatus to be thoroughly examined by engineers who can visualize what has occurred and record their observations for future use. A switch of very similar design to Fig. 1 in Dr. Garrard's \" Notes \" failed to open on a heavy cable fault where three feeders were operating in parallel. At the generating station the faulty feeder switch opened quite satisfactorily; at the substation, however, the two sound feeder switches opened and cleared the fault but also interrupted the supply to the substation. Upon an inspection being made of the defective switch it was found that one of the auxiliary sparking fingers had been welded to the bottom block sufficiently long to prevent the switch opening when the reverse relays came into action" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003017_1.1481369-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003017_1.1481369-Figure1-1.png", + "caption": "Fig. 1 Crank-slider mechanism and its bond graph representation", + "texts": [ + " The expressions of efforts and flows of I-type elements in integral causality give the dynamics equations and the flow balances give the constraint equations. Two examples are dealt with in order to illustrate the alternative procedures presented above. The first example is a crank-slider mechanism. LaCAP, lLCAP, HaCAP, and lHCAP are used to derive the corresponding dynamics equations. The second example is a pendulum. BHCAP and lBHCAP are used to write equations for this second example. Crank Slider Mechanism. The system and its bond graph representation are shown in Fig. 1. For the sake of clarity the signal bonds have not been displayed for the modulations of MTF elements. Each modulus is a function of one of the two variables u1 and u2 . The application of LaCAP, lLCAP, HaCAP, and lHCAP is recapitulated in Table 1 where numbers on the different causal bond graphs indicate the step order in which causality has been propagated ~encircled numbers refer to multiple strokes and squared numbers refer to added elements!. The resulting dynamics equations are given in Table 1" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002783_vib-48424-Figure3-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002783_vib-48424-Figure3-1.png", + "caption": "Fig. 3 Cross-section of angular contact ball bearing with reference coordinate", + "texts": [ + "99 40.0 0.0 38 20.684 19 12.03 7.1 5.72 3.99 40.0 1.794 9013 20.684 20 14.08 5.1 4.57 3.81 11.2 2.022 10138 20.684 21 1.207 3.8 8.38 3.81 231.8 0.0 46 20.684 22 0.654 0 0 0 0 0.0 42 0.0 3 Copyright \u00a9 2003 by ASME Terms of Use: http://www.asme.org/about-asme/terms-of-use Dow Table 2 Lumped parameters and cross-sectional properties of the gas generator rotor The bearing components are assumed to have contact only in the elastic region. Races are modeled as rigid except for local contact deformation. Figure 3 shows a cross-section of an angular contact ball bearing with reference coordinates. The inner race is externally loaded by the force vector {F}. The angular motions of the inner race are not considered here. The displacement vector {X} is },,{}{ zyxX T = . As shown in Fig.4, the inner race cross section at a ball is loaded by the contact force vector {Q} at the reference point p (inner race groove center), which has a displacement vector {u}. The parameter \u03c6 is calculated such that the r- z plane passes through the center of the ball" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003913_j.ijnonlinmec.2006.05.001-Figure6-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003913_j.ijnonlinmec.2006.05.001-Figure6-1.png", + "caption": "Fig. 6. Schematic of the 3-2-3 Euler angles ( 1, 2, 3) used to describe the orientation of a symmetric top. The axis of symmetry of the top lies along e3. The locus of the tip of the top describing a Type III rotation where the nutation angle 2 and spin angle 3 are constant is also displayed. The vectors e\u20321 and e\u20322 are the result of rotating E1 and E2, respectively, through 1 about E3.", + "texts": [ + " In other words, to an observer of the motion of a rigid body with a constant angular velocity either of the three types of rotations will be identical. What will distinguish them will be the choice of the basis that is fixed in space. For planar motions of rigid bodies, it is standard to choose E3 to be parallel to p and , thus rotations of Type I are discussed exclusively. For a rotation of Type III to be present, we require = 1e1 + 3e3, (52) where 1,3 are constant. A possible occurrence of such a motion is the steady precessional motion of a symmetric top with one point fixed in a constant gravitational field.5 As in Fig. 6, if we use a 3-2-3 set of Euler angles to parameterize the rotation of the top, then the angular velocity vector for this steady motion is = \u03071 sin ( 2 0 ) ( \u2212 cos( 3)e1 + sin( 3)e2 ) + ( \u03073 + \u03071 cos( 2 0) ) e3, (53) where 1 is the angle of precession, 2 = 2 0 is the constant nutation angle, and 3 is the spin angle. For this motion to be a Type III rotation, we require a constant spin rate: \u03073=0. This is an easy motion to demonstrate using a top: one simply holds the apex of the stationary top with one\u2019s fingers, tilts the top to give it a constant nutation angle 2 0, and then moves the top\u2019s apex to describe a circle. That is, the locus of the axis of symmetry e3 is a circle (see Fig. 6). However, such a steady motion cannot be sustained by the gravitational force alone and is consequently absent in the classical spinning top problem. It is interesting to note that related remarks pertain to the steady motions of a circular disk rolling without slipping on a horizontal plane.6 A further area of application arises in the kinematics of rods. For many rod theories, a rotation tensor P relating the directors d1 and d2 in the deformed configuration of the rod, to the fixed orthonormal directors D1 and D2 in the reference configuration is needed" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000522_hlca.19940770134-Figure5-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000522_hlca.19940770134-Figure5-1.png", + "caption": "Fig. 5. Switching and charge-rrapping prow~se.! ut , .ed~~.~--~~o/~. irrc , . hIkri~i,,lfilt~r-nzodifiL.d glassy carbon electrodes ( A = 0.07 an2, r in mol/cm\u2019) in 0 . 2 ~ Bu,N (CIO,)/MeCN at v = 100 mV/s. a ) Copolymer (6b + 7b) monolayer (r6b % 5 . lo-\u2019, r,, = 4. all redox levels equilibrate. b ) GC/D,/D, (rGb 5 2.4, lo-\u2019, r,, 5 1. lo-\u2019), charge in the outer layer is trapped and untrapped. c ) GC/D3/D2 (f,, 2 8 . lo-\u2019, f,, % 5 . charge in the outer layer is trapped (see text).", + "texts": [ + " An Organic Schrnitt-Trigger: Switching and Charge-Trapping Properties of the Redox-Polymer Bilayer Film. - Electropolymerization from a MeCN solution containing 6b/7b/pyrrole 1:1:0.2 leads to a modified glassy carbon electrode, denoted GC/ D, + D,. Sequential electropolymerization from a solution of 6b and pyrrole in MeCN followed by another electropolymerization onto the modified electrode, from 7b and pyrrole in MeCN, affords the redox-polymer bilayer-film-modified electrode GC/D,/D,. The reverse procedure is used to prepare GC/D,/D, (cf. Exper. Part). The cyclic voltammogram of GC/D, + D, (Fig. 5a ) in pure solvent-electrolyte corresponds essentially to the overlaid responses of the monolayer-modified electrodes GC/D, and GC/D, with the second redox couple being not resolved as EF(D,) - EF(D3) = 60 mV. A slight preference for the incorporation of 6b over 7b into the polymer is consistent with the observations described in Chapt.4. In GC/D, + D,, there exist four redox levels, which extend from the electrode surface to the polymer/solution interface, and these equilibrate with the electrode potential on the time scale of cyclic voltammetry", + " On the oxidative scan, the outer layer is - for the same reason - not discharged at E;(D,), i.e. the charge is trapped in the outer layer, but electrons shoot, if the electrode potential has reached the foot of the first anodic wave (charge untrapping). In terms of electronic devices, such behavior is typical for a bistable rnultivibrator or Schmitt trigger, able to transform the treshold of an analog signal (the electrode potential) into a digital signal (redox state of the outer polymer). For the inverse sequence of redox polymers (Fig. 5c), essentially the inner layer D, response without spike formation is obtained. The bilayer properties show up, if a sequential scan 0+-1.4+0+-1.4-t0 V is performed: The cathodic peak for the D;' reduction drops drastically on the second scan. This is related to the charge that stays trapped in the outer layer as indicated by the activation barriers involved (Fig. 5c). ') At p:&] = 1 . M, the additional current mediated by Df is supposed to be charge-transfer-controlled at the polymer/solute interphase as it reaches only ca. 80% of i, at the bare electrode, but does go to 100% at more negative potentials where D; is involved in catalysis [28]. 362 HELVFTICA HIMICA ACTA ~ Vol 77 (1994) Charge untrapping occurs by chemical oxidation of the outer redox couple or electrochemically, if a very positive electrode potential in the range of polypyrrole oxidation is applied" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002860_1.1427312-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002860_1.1427312-Figure1-1.png", + "caption": "Fig. 1. Optical system for observing the diffraction of light by ultrasonic waves in liquid.", + "texts": [ + " We then transform the data into a position\u2013time graph that represents the object\u2019s motion. We used the vision-based motion sensor in two experiments and compared its performance with that of conventional procedures. In the first experiment we measured the period of a simple pendulum. In the second experiment we determined the terminal velocity of a falling balloon. The results from these experiments provide insight into the accuracy of the position and time measurements obtained using the vision-based motion sensor. Fig. 1. A sequence of images taken over one period of the pendulum. 869 869Am. J. Phys., Vol. 70, No. 8, August 2002 Apparatus and Demonstration Notes This article is copyrighted as indicated in the article. Reuse of AAPT content is subject to the terms at: http://scitation.aip.org/termsconditions. Downloaded to IP: 202.28.191.34 On: Mon, 23 Feb 2015 10:56:13 The simple pendulum consisting of a bob attached to a cord was designed in such a way that the pendulum bob appeared darker than the background. A ruler with centimeter markers was used for calibration purposes. The camera was oriented perpendicular to the plane of motion. We acquired a sequence of 50 images, some of which are shown in Fig. 1. In this experiment, the geometric center of the pendulum bob coincides with its center of mass. In Fig. 2, the horizontal position of the center of mass with respect to its equilibrium position is plotted against time. The data points were fitted using a sinusoidal function, from which the period was determined to be 1.24360.001 s. Using a stopwatch to measure the period gave a value of 1.360.1 s, while with the rotary motion sensor, a period of 1.2560.01 s was obtained. The different experimental methods for determining the period are in good agreement with each other", + " In addition, the depth of each layer can be changed readily and its effects observed, both on the measured resistivity and on the distribution of the current density. Also, analysis of the two-dimensional paper model is somewhat simpler than the three-dimensional field model. Both horizontal and vertical layering are investigated, and results are obtained that are qualitatively equivalent to an actual geophysical survey. For simplicity, we model the resistivity of the earth using the Wenner electrode spread.4 This electrode spread is commonly employed by geophysicists and is characterized by an equal spacing between adjacent electrodes. ~For example, see Fig. 1.! The outer two electrodes are connected to a variable power supply and the current is held constant. The electric potential difference between the inner two electrodes is then measured to determine the resistivity of the ground. To derive the resistivity for a homogeneous subsurface using a Wenner electrode spread, one applies Ohm\u2019s law and the definition of the electric field as the gradient of the electric potential to an infinite hemisphere.4 The result of this derivation is that the electric potential difference (DV) between the inner two electrodes is determined by DV5rI/2pa , where a is the spacing between electrodes, r is the resistivity, and I is the current", + "6 The kit includes 23 cm330 cm conductive paper sheets with a printed grid, regular white paper sheets with a similar grid, and larger 30 cm346 cm conductive sheets without a grid. In addition, the kit includes silver paint for drawing conducting distributions and pushpins to connect the conducting distributions to the wires from the power supply. Horizontal layers. To model a Wenner array over horizontal layers, we use the large conductive paper with pushpins inserted on prepainted silver pads at the points where measurements will be taken, as shown in Fig. 1. The pushpins represent current and potential electrodes and the top of the paper represents the surface of the earth. To simulate a lower resistivity layer at some depth ~for example, water saturated sediment! we use a 1 in.31 in.320 in. brass bar that can be moved to any desired position below the pushpins. The brass bar is first placed at the bottom of the paper, a distance of 28 cm from the pushpins, and a Wenner spread is conducted for each of the possible a-spacings. The current is held constant at 100 mA for the entire experiment. The brass bar is then moved upward to depths of 24, 20, 16, 12, 8, 6, 4, and 2 cm and subsequent Wenner spreads are conducted for each new position. Sample data for the electric potential difference (DV) between the inner electrodes as a function of a-spacing are shown in Fig. 2. The natural logarithm of DV has been plotted to emphasize relative changes. For clarity, data for suc- Fig. 1. Four electrode Wenner spread on paper with adjustable brass bar. The power supply ~represented as a variable battery! is attached to the outer two electrodes and the current is held constant in this section of the circuit. The electrical potential difference ~labeled DV) between the inner two electrodes is measured. To accommodate other a-spacings, all four electrodes are moved, keeping the center of the electrode spread stationary. For example, the curved arrows indicate how the spread would be relocated from its present location ~for an a-spacing of 10 cm", + " An old ultrasonic nebulizer can usually be obtained at a pharmacy; it\u2019s a nebulizer that has a piezoelectric transducer connected to a radio-frequency generator ~in our case a 25 MHz one! that works at ultrasonic frequencies ~as the name indicates!. The advantage of the ultrasonic nebulizer is that no circuit design is needed and it\u2019s inexpensive. A test tube, whose bottom was previously cut out, has been glued around the piezoelectric disk of the nebulizer. To make it possible for us to change the effective length of the tube, a piston of the same diameter of the test tube was used. The optical arrangement of the apparatus is shown in Fig. 1. A plane acoustic wave front is sent through the liquid at a right angle to the direction of propagation of light. After passing through the liquid in the test tube, the laser beam turns into a divergent wave front as shown in Fig. 2. The angular positions of the diffraction maxima for a liquid grating were found by Raman and Nath3 to be given by a sin fm5ml ~m50,1,2,...!, ~1! where fm is the angle corresponding to the mth order diffracted beam, a is the spatial periodicity of the grating, and l is the wavelength of light in air" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003674_j.mechmachtheory.2004.05.002-Figure3-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003674_j.mechmachtheory.2004.05.002-Figure3-1.png", + "caption": "Fig. 3. Geometrical parameters associated with two arbitrary helical joints.", + "texts": [ + " (1), we can now state the following result: r e \u00bc rQ Qe0 \u00feQ\u00f0r0 rQ\u00de Qe0 \u00fe dk Qe0 \u00f02\u00de The above equation can be simplified by noting that the cross product of two rotated vectors is identical to the rotated cross product, thereby obtaining r e \u00bc Q \u00f0r0 rQ\u00de e0 \u00fe \u00f0rQ \u00fe dk\u00de Qe0 \u00f03\u00de Thus, Eqs. (1) and (3) fully characterize the current configuration B of a rigid body after a general displacement from a reference configuration B0. It should be noted that, by using such a purely geometric approach, a great deal of physical insight into the finite-displacement-problem is gained, and as it will be shown later, it is possible to further extend these results to infinitesimal displacements. Consider two screws associated with two arbitrary helical joints, as shown in Fig. 3. In this figure, O denotes a fixed point, r1 and r2 are the position vectors of points on lines L1 and L2, h1 and h2 are the pitches of screws $1 and $2, e1 and e2 are unit vectors parallel to lines L1 and L2, respectively. Thus, the involved screws are given by $1 \u00bc e1 r1 e1 \u00fe h1e1 ; $2 \u00bc e2 r2 e2 \u00fe h2e2 \u00f04\u00de Then, the Lie product, also known as motor product or dual vector product was defined [3,17] as $12 $1 $2\u00bd e1 e2 e1 \u00f0r2 e2 \u00fe h2e2\u00de e2 \u00f0r1 e1 \u00fe h1e1\u00de u v \u00f05\u00de which is apparently another screw", + " Thus, to this end, and, in order to have a more general geometrical interpretation of the Lie product, we have the following. Theorem. Regardless of which points are chosen to specify the screws associated to two arbitrary lines, the Lie product of such screws is always a twist, of magnitude sin a12 around a screw of pitch a12 cot a12 \u00fe \u00f0h1 \u00fe h2\u00de and associated to the line that joins the screw-lines L1 and L2 along the common perpendicular between them. Proof. Consider two arbitrary lines L1 and L2 associated with two arbitrary screws $1 and $2, as shown in Fig. 3. In this figure, points 1 and 2 are arbitrarily selected points, whereas points A and B are particularly selected points, i.e., they are particularly located at the intersections of the screw-lines and their common perpendicular. Thus, the positions of such points can be related as follows: r1 \u00bc rA \u00fe r1=A; r2 \u00bc rB \u00fe r2=B \u00f06\u00de On the other hand, the terms appearing in the Lie product, Eq. (5), can be arranged as follows: e1 e2 \u00bc \u00f0sin a12\u00dee \u00f07\u00de e1 \u00f0r2 e2\u00de \u00bc r2 \u00f0e2 e1\u00de e2 \u00f0e1 r2\u00de \u00f08\u00de e1 \u00f0h2e2\u00de \u00bc \u00f0h2 sin a12\u00dee \u00f09\u00de e2 \u00f0h1e1\u00de \u00bc \u00f0h1 sin a12\u00dee \u00f010\u00de where Eq. (8) is just a triple vector product identity, and e is a unit vector along the common perpendicular to the involved lines, L1 and L2, as shown in Fig. 3. In this way, we have now that v \u00bc r2 \u00f0e2 e1\u00de e2 \u00f0e1 r2\u00de \u00fe e2 \u00f0e1 r1\u00de \u00fe \u00f0h2 sin a12\u00dee\u00fe \u00f0h1 sin a12\u00dee By using expressions (6), (7) and observing that rA rB \u00bc a12e, we can obtain v \u00bc sin a12\u00f0rB e\u00de \u00fe \u00f0a12 cos a12\u00dee\u00fe sin a12\u00f0h1 \u00fe h2\u00dee \u00f011\u00de Thus, the Lie product finally takes the following form: $12 \u00bc sin a12 e rB e\u00fe \u00f0a12 cot a12 \u00fe h1 \u00fe h2\u00dee \u00f012\u00de which is apparently a twist of magnitude sin a12 with pitch \u00f0a12 cot a12 \u00fe h1 \u00fe h2\u00de associated to the line L, thereby completing the proof. h As it was shown previously, the Lie product involves two screws, namely, a driver screw and a driven screw" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003674_j.mechmachtheory.2004.05.002-Figure4-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003674_j.mechmachtheory.2004.05.002-Figure4-1.png", + "caption": "Fig. 4. Rigid body displacement associated with a helical joint.", + "texts": [ + " This transformation is called in mathematics as the adjoint representation of the Euclidean group. Although the derivation of the Lie product for the revolute and prismatic pairs could be considered as particular cases of the helical pair, we will present each derivation separately in order to show the special features and details of each case and also in order to formulate a comprehensive and complete derivation process. In order to derive the Lie product when the screw axis of the displacement coincides with the motion axis of a helical joint, we will refer to the geometry shown in Fig. 4. Then, for the case under study, we have that $ \u00bc e r e\u00fe qe ; $i \u00bc ei ri ei \u00fe pei ; $j \u00bc ej rj ej \u00fe pej \u00f013\u00de being q and p the pitches corresponding to the screws $ and $i, respectively. It should be noted that this particular case rightly corresponds to the definition of the Lie product, Eq. (5). On the other hand, from Fig. 4, we have now the following vector equation: rj \u00bc r\u00fe p0 \u00fe dk \u00f014\u00de It can be shown [25] that the rotated vector p0 can be computed as follows: p0 \u00bc Qp; Q \u00bc eeT \u00fe cos /\u00f01 eeT\u00de \u00fe sin /E \u00f015\u00de being / the rotation angle about unit vector e, the term 1 representing the 3\u00b7 3 identity matrix and E denoting the cross product matrix of vector e, i.e., e p Ep. Moreover, for a helical joint, vector dk is given by dk \u00bc q/e \u00f016\u00de Now, substituting Eqs. (15) and (16) into Eq. (14), we obtain that rj \u00bc r\u00feQp\u00fe q/e \u00f017\u00de Additionally, we also know that ej \u00bc Qei \u00f018\u00de Then, using (17) and (18) we can now perform the following cross product: rj ej \u00bc r Qei \u00feQ\u00f0p ei\u00de \u00fe q/e Qei \u00f019\u00de Thus, the derivatives of ej and rj ej with respect to parameter / may be evaluated [27] by differentiating Eqs. (18) and (19) and letting / ! 0, to get oej o/ \u00bc oQ o/ ei \u00f020\u00de o\u00f0rj ej\u00de o/ \u00bc r oQ o/ ei \u00fe oQ o/ \u00f0p ei\u00de \u00fe qe Qei \u00fe q/e oQ o/ ei \u00f021\u00de where, from Eq. (15), the following partial derivative can be easily calculated: oQ o/ \u00bc sin /\u00f01 eeT\u00de \u00fe cos /E \u00f022\u00de Thus, after letting / ! 0, we have that oej o/ \u00bc e ei \u00f023\u00de o\u00f0rj ej\u00de o/ \u00bc r \u00f0e ei\u00de \u00fe \u00f0e p\u00de ei \u00fe qe ei \u00f024\u00de But, from the geometry shown in Fig. 4, it can be noted that p \u00bc ri r and, by using the following property of the vector triple product, namely, r \u00f0e ei\u00de \u00fe e \u00f0ei r\u00de \u00fe ei \u00f0r e\u00de \u00bc 0, we have that o\u00f0rj ej\u00de o/ \u00bc e \u00f0ri ei\u00de ei \u00f0r e\u00de \u00fe qe ei \u00f025\u00de Finally, the partial derivative of the last term appearing in the expression for $j, see Eq. (13), can be calculated as follows: o\u00f0pej\u00de o/ \u00bc p oQ o/ /!0 \" # ei \u00bc p\u00f0e ei\u00de \u00bc e \u00f0pei\u00de \u00f026\u00de Then, we can now state that o$j o/ oej=o/ o\u00f0rj ej\u00de=o/ \u00fe o\u00f0pej\u00de=o/ \u00bc e ei e \u00f0ri ei\u00de \u00fe e \u00f0pei\u00de ei \u00f0r e\u00de ei \u00f0qe\u00de \u00f027\u00de It should be noted that, according to the foregoing procedure, $j is nothing but the screwen $i after an infinitesimal change of the independent parameter of motion" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003437_0954406042690498-Figure10-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003437_0954406042690498-Figure10-1.png", + "caption": "Fig. 10 Three-dimensional solid models of the pinion and gear at the meshing state", + "texts": [ + " The input angle of the pinion, dimensions of the contact ellipse, specific slidings and the kinematic error for each contact point are listed in Table 2. At the initial and final contact points, kinematic errors are equal to 20 arcsec. The magnitude of the kinematic error function conforms with the designed value for x. The maximum absolute value of specific sliding occurs at the initial contact point on the pinion; thus, the most severe abrasion occurs at this place. Three-dimensional solid models of the pinion and gear at the meshing condition are provided in Fig. 10. It is noted that the teeth are symmetric, the tooth traces on both sides of the teeth of the pinion are concave and the tooth traces on both sides of the teeth of the gear are convex. With the theory proposed in section 4.3, designers can determine the values for parameters of cutting tools satisfying a pre-designed magnitude of a parabolic-like kinematic error function. Here, the pre-designed magnitude is varied from 10 to 30 arcsec and the parabolic parameter of the cutting edge for the pinion is varied from 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002998_esej:20020106-Figure4-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002998_esej:20020106-Figure4-1.png", + "caption": "Fig. 4 Third generation 1999", + "texts": [ + " The adoption of a digital controller and provided interface solution took away the challenges of design with a microcontroller (e.g. which pins/ports to use for which function etc.). Develop mechatronics skills more Illy, i.e. the ability to utilise algorithmic sophistication to enhance mechanical components to deliver superior performance and to integrate these, in some meanin@ way, with better sensing and IT. Provide a means of deepening prior knowledge and improve the aspects of teamwork. This resulted in the next generation kit shown in Fig. 4. This was a major departure from the ENGINEERING SCIENCE AND EDUCATION JOURNAL. FEBRUARY 2002 previous generations. A more advanced processor with analogue input, PWM and communications fimctionality was chosen (PIC 17C756). The processor board was kept general purpose so as to provide systems design challenges in deciding how to partition h c - tionality and which ports/ pins to use. LEDs were provided in the general digital I/O region to help students troubleshoot their designs. The design choice was enhanced by the adoption of self-contained drive modules" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003864_50002-2-Figure2.5-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003864_50002-2-Figure2.5-1.png", + "caption": "Fig. 2.5. Rolling and slipping of a tyre over an undulated road surface.", + "texts": [ + " When the wheel is locked (g2 = 0) we obviously have x = - 1. In the literature, the symbol s (or S) is more commonly used to denote the slip ratio. The angular speed of rolling Or more precisely defined for the case of moving over undulated road surfaces, is the time rate of change of the angle between the radius connecting S and A (this radius is thought to be attached to the wheel) and the radius r defined in Fig.2.3 (always lying in the plane normal to the road through the wheel spin axis). Figure 2.5 illustrates the situation. The linear speed of rolling Vr is defined as the velocity with which an imaginary point C* that is positioned on the line along the radius vector r and coincides with point S at the instant of observation, moves forward (in x direction) with respect to point S that is fixed to the wheel rim: V - re~'-~ r (2.6) For a tyre freely rolling over a flat road we have: ~c~ r --~-~ and with ~,~ = 0 in addition: Vr = Vx. Note, that at wheel lock (f2 = 0) the angular speed of rolling Or is not equal to zero when the wheel moves over a road with a curved vertical profile (then not always the same point of the wheel is in contact with the road)", + " For a given tyre the effective rolling radius re is a function of amongst other things the unloaded radius, the radial deflection, the camber angle and the speeA of travel. The vector for the speed of propagation of the contact centre V~ representing the magnitude and direction of the velocity with which point C moves over the road surface, is obtained by differentiation with respect to time of position vector c (2.21): V c - d - /; +d + t : - V + t : (2.26) With V the velocity vector of the wheel centre A (Fig.2.3). The speed of propagation of point C* represented by the vector ~ becomes (cf. Fig.2.5 and assume re/r constant): r Vc* - V + --Let: (2.27) r The velocity vector of point S that is fixed to the wheel body results from ?. V s - V + --Leto \u2022 (2.28) r with o9 being the angular velocity of the wheel body with respect to the inertial frame. On the other hand, this velocity is equal to the speed of point C* minus the linear speed of rolling V s - V c - V 1 (2.29) from which Vr follows: 1/r - l\" ( V c - V s) (2.30) o r V - Vcx - Vsx (2.31) The linear speed of rolling is according to (2.6) related to the angular speed of rolling: s - 1 V (2" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002353_iros.2000.895310-Figure3-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002353_iros.2000.895310-Figure3-1.png", + "caption": "Fig. 3 . Error function and incremental transformation.", + "texts": [ + "3 Estimation algorithm Let the detected point of the end tip of the conic calibration object be a two dimensional point where j is n for nominal point and r for the real, measured point. The error function is then determined as a 2- dimensional function (5 ) To calculate the nominal points in the image we need to calculate the end points of the calibration tool in the robot frame and transform them into the sensor frame, after which we can apply the pin hole camera model to map the points into the image plane. Let the calibration point in the robot tool h e be a homogeneous point Then the point in the sensor coordinate frame is (fig. 3) (7) where H , , is the homogeneous transformation from the robot world (or base) fi-ame into the sensor frame, H , , is the homogeneous transformation from the robot world (or base) frame into the robot tool frame and p s is the point in the sensor fi-ame. The point in the image plane becomes then f *PS,X f *PSJ T Pim,n - -1 f -Ps,z f - P , z - -[ where f is the focal length of the camera, and psJ ; j= x, y or z is the point in the sensor frame. The state vector consists of the sensor pose parameters s , and the location of the calibration object\u2019s end tip in tool fi-ame p, " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002289_aim.2001.936431-Figure4-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002289_aim.2001.936431-Figure4-1.png", + "caption": "Fig. 4 shows coordinate system adopted for kinematics analysis. Point 0 is the origin and the criteria coordinate is the base coordinate system.", + "texts": [ + " Inverse Kinematic Analysis i a b T b dt Tt 2 -L. li . S posture of end-effector rotation matrix, which represents the posture of the end-effector, chain number, vector from center of base to base joints, length of a b , vector from center of end-plate to endtable joints, length of 4, unit vector from base joint datum point t o actuator joints datum point, length of chain of prismatic joints, unit vector from actuator joints datum point to end-plate joints datum point, From relations between base joints and end-effector joints (Fig. 4), we get p + R&j - a b i = l j . ~ + b s . (3) where Li is used for p + m, - a b j . Both sides of equation are squared and because z2 = 1, s2 = 1, and we get the following equation. This equation is solved for l i , and we get, li = (Li . z) f d ( L j . z ) ~ - Lq + b2. (6) Where, Lj = (&, L,, Lz) , z = (O,O, 1) and using the constraints that the chains sign of square root is negative, we get the following equation. (7) lj = L,i - b2 - L2. - L2. J x1 yzl Thus, the length of the link 11 is determined by the tip position of the end-effector F(z , y, z, 4,8, $)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003162_iros.1996.568991-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003162_iros.1996.568991-Figure1-1.png", + "caption": "Fig. 1. The simulated robot arm environment used in the experiments.", + "texts": [ + " The model is a network that is trained to indicate if a pair of configurations is linked by a straight collisionfree trajectory [13]. This kind of approaches needs a previous stage for the acquisition of the model. On the contrary, our reinforcement-based approach allows the robot to interact with the environment from the very beginning and to improve its performance as learning proceeds. 2 The Simulated Robot Arm Environment We have simulated a simple two-link robot arm with two revolute joints (see Fig. 1). Both joints rotate in the same plane with angle (ql,q2). The angular movements of the joints are constrained as follows: qlmin S q1 I qlmm and q2min < - q2 I q2max where Robot links are straight line segments of the same length. Also, to simplify calculations, obstacles are represented by circles. Each link has three rings of range sensors that are evenly placed. Each ring has two sensors pointing to the right and left side of the link, respectively. The sensors provide distances to the obstacles around the robot" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002700_fie.1997.632672-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002700_fie.1997.632672-Figure1-1.png", + "caption": "Figure 1. Schematic of laboratory gantry crane.", + "texts": [ + " There are many electro-mechanical systems that could be used to satisfy the learning objectives of the courses discussed in this paper. We have chosen the gantry crane for two reasons, 1) it has many characteristics which can be used to help students learn in a variety of different Mechanical Engineering courses, and 2) its dynamics are such that many concepts can easily be visualized in realtime by students. This type of system has been cited as an educational tool in other works, for example, [2-31. The rig, shown in Fig. 1, is a laboratory-scale model of a gantry crane that is typically used to move heavy loads in many manufacturing and other commercial environments. The crane is made up of a carriage (motor and cart) which rolls on rails and which supports a suspended load. The rails are mounted to the ceiling in the laboratory, and provide 14 feet of track on which to maneuver the crane. The motor is a dc servomotor, which is coupled to the drive wheels through a gear train and a belt. Using the current motor, the maximum speed of the crane is approximately 5 #sec, and the bandwidth of the system is approximately 15 Hz" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001966_ijmtm.2002.001439-Figure5-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001966_ijmtm.2002.001439-Figure5-1.png", + "caption": "Figure 5 Various ramp configurations", + "texts": [ + " During a corkscrew test, a vehicle with longitudinal velocity runs over a ramp. Initially, the wheels from one side of the vehicle contact the ramp, while the wheels on the other side are in contact with the ground. As the vehicle continues to move up the ramp, it experiences a high asymmetric acceleration from the vertical direction causing it to rotate along its longitudinal axis. When it leaves the ramp, the vehicle continues to rotate along its longitudinal axis until it contacts the ground. Figure 5 shows various ramp configurations of different height, width and length that have appeared in the literature. One type, shown in Figure 5(d), is a straight-up-flat-surface ramp, while the other, shown in Figure 5(b), is a spiral ramp with a curved surface. These are described in Sections 2.2.1.1 and 2.2.1.2. It is designed in such a way that the side edge of the ramp should not interfere (or contact) the bottom of the test vehicle. This requirement eliminates one source of possible test variation. A ramp that is designed to meet this requirement for testing small size cars to light trucks and SUVs is shown in Figure 5(d). An early corkscrew ramp test was incorporated into SAE Recommended Practice J857 using the ramp shown in Figure 5(a). According to Wilson and Gannon (1972), this procedure, developed by Chrysler, uses a curved guide-rail to direct the test vehicle up the ramp (see Figure 6) to a point high enough to initiate vehicular roll. Sakurai et al. (1991) used the SAE J857 ramp for their rollover study by conducting a series of 12 ramp rollover tests at 50 km hr 1 (31 mph) as shown in Figure 7. The roll motion was developed by driving the right side tires onto the ramp while quickly turning the steering wheel to the right, resulting in a left-side-leading rollover. Source: Sakurai et al. (1991). Issues pertaining to this test mode may include: 1 getting the test vehicle on the ramp in the same manner every time 2 whether the steering wheel on the test vehicle should be locked in position or not during test. The ramps shown in Figure 5(b) and (c) were used by TUV of Europe. Figure 5(b) shows a spiral ramp. The ramp, shown in Figure 5(c), consists of four segments: the first three segments of lower heights are an integral part (referred to as a three-segment ramp), while the fourth segment can be separated from the entire fixture (which is referred to as a four-segment system). Given a forward velocity, the three-segment ramp is used for non-roll tests, while the four-segment ramp is used for roll cases. The sequence of a vehicular motion using TUV ramp is shown in Figure 8 (Wech and Ostmann, 1996). Source: Wech and Ostmann (1996)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002679_ias.1995.530622-Figure4-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002679_ias.1995.530622-Figure4-1.png", + "caption": "Fig. 4: Cross Section of a Two Winding CWT", + "texts": [ + " For example, the short circuit driving point admittance of circuit 2 is the admittance of circuit 2 with all other windings short circuited, (4) Y22 = Y12 + Y23 + Y24 The values of the admittance coefficients in (2) and the branch admittances of Fig. 3 can be determined experimentally by a number of tests where one of the windings is excited with all other windings short circuited [ 11. B ) Mulii- Winding CO-axial Winding Transformer The co-axial winding transformer (CWT) has many unique features compared with conventional winding transformers [3,4]. Fig. 4 shows a cross sectional view of a two winding CWT. Since the inner winding is totally enclosed by the outer winding, all the flux produced by the outer winding will link the inner one. In addition, the leakage field can only exist within the winding space between the inner and the outer windings. This allows the leakage inductance to be both controlled and minimized. Fig. 5 shows the physical construction of a multiwinding co-axial winding transformer. The cross section of the winding shows a bundle of co-axial cables where the parallel connected outer conductors (shields) form the primary winding while each of the inner conductors forms one of the secondary windings" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002360_icec.1995.489164-Figure8-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002360_icec.1995.489164-Figure8-1.png", + "caption": "Fig. 8: Two-link manipulator", + "texts": [ + " MGA was able to discover a good controller that was successful for more than 100,000 time steps in about 200 and about 650 generations for the singlepole and double-pole respectively. Fig. 6 shows the control result of the single pole system using the neurocontroller found by MGA. Initial condition for 6 is 0.1 rad and all other states are zeros. Fig. 7 shows the result of the double pole system. 0, is initially 2.5\u201d and all other states are zeros. f ( x ) = - 1 + 2 / ( l + e - \u201d ) (15) 4.3 Robot manipulator control The nonlinear system considered here is a one-link and a two-link robot manipulator. Fig. 8 shows the two-link manipulator system. A neural network is used as a nonlinear PD-type controller for the robot (17) 3 =[id1 +r, +4 +&+4&c; +I2% .t.Z((z +V&J+4l% 3 =k& +1,I,C;>+l,l?] +@& +Iz% +WgA4 +m$,cl(18) -wgA(%142 d)+nrd,lc;+rris(llc;+&G) where the following notations were used. q, NI, Mass of link i I i 1, Length of link i $, (= 1, 12) Joint angle of joint i Moment of inertia of link i Distance between joint i and the center of mass of link i S, = sinq, , C, = cosg, , Cu = cos(g, +q,) The simulation parameters are vi1 = 10 kg, in2 = 5 kg, I , = I , = 1" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000765_s0378-4754(97)00149-3-Figure4-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000765_s0378-4754(97)00149-3-Figure4-1.png", + "caption": "Fig. 4. Signals u and w.", + "texts": [ + " Case (b) and (c) @g i n @t i n @g i n @Xi n Pi n 0 4.2.1. @gin @tin @gin @Xi n Pi n 0 It can be the case of the discontinuous conduction mode. (A current which is zero on a period is not a discrete state variable). 4.2.2. @gin @tin \u00ff @gin @Xi n Pi n When a switching instant is given by gi n Xi n;Un; t i n u Xi n Xn;Un; t i n \u00ff w ti n we have the property if du dt ti n dw dt ti n (19) Then the curve w is tangent to the signal control u at just after ti n. This case is illustrated in the case b of Fig. 4. In this figure, we see that a perturbation which subsists in the case (a) (Fig. 4(a)) is immediately recaptured in case (b) (Fig. 4(b)): we have an order reduction and a necessary condition for a deadbeat response. We consider here the case of a two-phased switched reluctance motor (SRM) fed by an asymmetrical converter (Figs. 5 and 6). The parameter controls are the two angles noted and p. In this section we are working in open loop. The case considered here contains some important particularities: the motor as well as the inverter are two-phased, and the control chronograms of the two phases are independent (Fig. 7). The inverter provides full wave voltages with variable frequency" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000433_elan.1140040410-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000433_elan.1140040410-Figure1-1.png", + "caption": "FIGURE 1. Amperometric cell with a spe membrane. (A) overall view; (6) expanded view of the detection space. (1,2) two parts of the detector body, (3) spe membrane, (4) auxiliary electrolyte solution, (5,6) reference and auxiliary electrodes, (7) working wire electrode, (8) electrical contact to the working electrode, (9) Plexiglas cylinder, (1 0) fixing screw, (11) inlet capillary, (12) seal, (13) fixing screw, (14) outlet.", + "texts": [ + " Therefore, the present paper describes an amperometric flow-through cell employing a thin and short bright platinum wire working electrode protrudmg into the detection space through a Nafion membrane. The membrane is in contact with a solution of an auxiliary electrolyte outside the detection space and the auxiliary electrolyte contains conventional reference and counter electrodes. The difference between the geometric and microscopic surface areas of this working electrode is small and thus favorable S/N value and lower limits of detection can be expected The detector cell (Figure 1) consists of two cylindrical parts, 1 and 2, made of Plexiglas or PTFE, between which the spe membrane, 3, is pressed and serves both for the electrolytic connection among the electrodes and as the seal between the two parts of the cell (the bolts connecting the two parts of the cell are not shown). The membrane is a 15 x 20 mm piece of Nafion 117 (0.18 nun thick, Aldrich, USA, Cat. No. 29,256-7). Prior to use, Nafion was boiled for ca. half an hour in 1 M sulfuric acid. Part I of the cell contains space 4 for the auxiliary electrolyte solution in which a silver/silver chloride reference electrode, 5, and counter electrode G (a 0", + " Cylinder 9 is placed in the appropriate opening in cell part 1 and pressed with screw 10 to the spe membrane so that the platinum wire penetrates through the membrane (it is advisable to puncture the membrane beforehand with a stainless steel needle with a diameter smaller than that of the wire). The platinum wire then protrudes into the detection space at a length of M . 0.2 mm. The detection space in cell part 2 is 2 mm in diameter and the test liquid is fed into it by PTFE capillary 11 whose end is ca. 0.2 mm from the spe membrane surface, opposite to the working electrode (see the expanded view in Figure 1B). The position of capillary 11 is maintained by silicone rubber seal 12 and screw 23. The liquid leaves the detection space through outlet 24. The geometric volume of the detection space is ca. 0.6 to 0.7 pl. Testing of the Cell. The cell was tested in FIA and HPLC systems. The FLA system consisted of an LKB 2150 high-performance pump (LKB, Bromma, Sweden), a Rheodyne 7125 sampling valve (Cotati, USA) with 10, 20, 50, and 200 pl loops and the studied detection cell, connected to the sampling valve by a PTFE capillary 30 cm long, with an internal and external diameters of 0", + " A strong effect of the analyte diffusion on the response is also demonstrated in Figure 3 depicting the dependence of the current on the mobile phase volume flow rate. Curve A in Figure 3 corresponds to the steady-state current dependence. The exponent of this dependence is only 0.16, much less than expected for convection-diffusional transport (from 0.3 to 1.0 in dependence on the character of the flow 111 ). This suppressed dependence on the flow velocity can be explained as follows: The liquid entering the detection space (see Figure 1B) hits the electroinactive tip of the platinum wire electrode and the lines of flow are diverted from the direction perpendicular to the spe membrane. The electroactive part of the platinum wire close to the membrane is thus shielded from the flow. This effect is probably enhanced by local turbulence at the wire tip. Therefore, the stagnant layer of the liquid at the spe-platinum-mobile phase interface is much thicker than with a cylinder placed freely in a flowing liquid and molecular diffusion predominates in controlling the analyte transport towards the electroactive part of the wire" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000471_bf01178518-Figure2-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000471_bf01178518-Figure2-1.png", + "caption": "Fig. 2. a Elastic friction contact: limit and loading surfaces, b elastic-plastic friction contact: slip and configuration surfaces, e elastic-plastic friction contact: composed superellipsoidal surfaces", + "texts": [ + " The total displacements or strains are decomposed into elastic and slip (or plastic) displacements and strains, so that Ni : Ui e @ uiS~ ~ : ~e ~ vs 7i = 71 e + 71 s, ~2 = 72 e + 72 s, 8 n : 8n e Jr- gn s (2) and a similar decomposition applies to their rates or increments. For an elastic, friction contact problem the sliding rules were associated with the limit friction surface FL(a, %, %) = 0. Slip rules for consecutive loading events were generated by the active loading surfaces F~(a, zl, %) = 0 moving inside the interior domain bounded by the limit surface, c.f. Fig. 2 a for i = 1. In the present model we assume similarly that the overall sliding occurs when the stress state reaches the external configuration smface FL(a, %, %, e ) = 0, and the microslip effects are described by the internal slip (or yield) surface Fo(a, %, %, fl) = 0, where and fl denote the set of internal variables specifying the irreversible changes of the contact layer. The stress states corresponding to the interior of the yield surface correspond to elastic response / o o 00 F;=O %=0 Fig. 2 a 61 of the contact. The configuration surface for an elastic friction contact is represented by an open cone in the stress space, similarly as the limit surface for the case of two contacting spheres, Fig. 2 a. However, assuming plastic deformation of asperities and the contact zone, the configuration surface is assumed as a closed surface. Assuming the associated slip rule, the contact zone dilatancy would be generated in Fig. 2b for the stress regimes CB and CD and contact zone compaction would occur for stress regimes represented by BE and DE. Obviously, for some contact surfaces the normal stress required to reach the compaction regime would be very high and all contact slip would usually proceed in the dilatancy regime, a < eL. To describe the configuration surface, let us use the superelliptical description ([G -- C~L['~\" ( [ r l [ ~ \" ( ]Z2])\" __ 1 = 0 ' FL= \\ C +C~L/ + \\b , ] + \\b2,] (3) or a composed superelliptical form FL('> = + \\ b, / \\ b, / (5 < (~L, o ' > ~ L . (4) For 1 < nl < 2, we obtain the oblong shape with obtuse \"corners\" along principal axes, for n2 > 2, the superellipse is tending to a more \"swollen\" shape, Fig. 2c. By using the composed forms (3), a variety of forms of the configuration surfaces can be generated. As the state variables, we can assume er, bl, b2, a, and C, where eL + C, bl, b2 and a are the semiaxes and C represents the contact tensile strength which may be exhibited in the initial stage of slip. In particular, for n = 2, bl = bz, Eq. (2) provides the rotational ellipsoid ~0\" - - ~LX~ 2 Fr= \\ ~ , / + ( b ) Z + ( b ) 2 - 1 : 0 (5) depending on three parameters C, eL and b. To specify the evolution of contact state variables, assume that the accumulated measure of slip is expressed as follows: T T fi(T) = ~ [fi[ dt = ~ (zil z + fi22) I/2 dr, (6) 0 0 where t is a time-like parameter", + " If after application of a normal compressive stress a the shear stress is monotonically increasing at constant o-, it is assumed that the active loading surface F1 = 0 develops along the loading path so that the stress point remains on this surface 9 The active loading surface remains tangential to the limit surface FL = 0 at the tension point C, o. = - C, and lies within the domain enclosed by the surface FL = 0. The consecutive loading surfaces Fi = 0 and Fl ' = 0 associated with the stress path 1 - 1' are shown in Fig. 2 b. The yield surface F0 = 0 is assumed to be translated with the stress point inside the loading surface, so that they become tangential to each other along the stress path. This assumption is identical to that used in multisurface plasticity formulation [3]. The evolution of the center of elliptical active loading surface (n = 2) can now be easily formulated. In fact, this surface is tangential at C to the limit surface FL = 0 and passes through the actual stress point, so that the equation = - - + - 1 = 0 (9) \\ rl / provides the actual value of ~1 ~", + " (11) The slip surface of constant size parameter ro remains tangential to Fi = 0, so the following proport ion exists: r 0 T - - ~0 r O\" - - (XO a - - ~ , ( 1 2 ) Y 1 T O\" - - ~1 and the evolution of the center of the slip surface is specified by the formulae following from (12): = - - + (O- - - r i + C ) (6\" + C ) -}- m o 2 z 2 (o- + C) mo2z ro + \" - - dz, (o. -- ri + C) (o. + C) + moaz 2 rl (o. -- rl + C) ~ . ro do. dc~~ = (o. - ri + C) (o. + C) + mo2z 2 r~ [(to) mo2 2 + 1 - ~ + ( o - - r ~ + C ) ( o - + C ) + m o 2 z 2 9 r~l do. (13) When the loading path reaches the limit surface, then F1 = FL ---- 0 and the subsequent stress path could lie on FL = 0 or constitute unloading. Figure 2 c shows the evolution of active loading and yield surfaces for superelliptical forms of these surfaces, where 1 < n < 2 for a < eL and n > 2 for a > eL, and eL denotes the stress value at the connecting line between two superellipses. Consider now the unloading process, with the stress path penetrat ing inside the loading surface. It can be assumed that the active loading surface F2 -- 0 develops along the unloading path so that it passes through the actual stress point P and remains tangential to the pr ior Phenomenological model of contact slip 65 loading surface F~ = 0 associated with the stress unloading point, Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000726_cbo9780511530173.007-Figure5.5-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000726_cbo9780511530173.007-Figure5.5-1.png", + "caption": "Figure 5.5 A passive three-parameter spring actuated by a 3R manipulator.", + "texts": [ + " Columbia University - Law Library, on 13 Aug 2019 at 15:09:03, subject to the Cambridge Also, a change in the constraint force 8fn cannot produce motion. As stated earlier, such forces and displacements are reciprocal and >, = 5/[0.707, 0.707] Spt \\-\u00b0J07} = o. [ 0.707 J It follows that the small slider displacements 8dc and 8dt can be used for the simultaneous control of normal force and tangential motion. By superposition \\ = GcSfn\\ \\ + G, 8pt \\. (5.40) Sd2\\ \\ 0.0000 0.707 In (5.40) Gc and Gr are dimensionless scalar gains, and 8pt and 8fn are errors in the wheel position and the normal contact force. Figure 5.5 illustrates a planar three-re volute serial manipulator with a compliant wrist. The workpiece, which is held fixed in the gripper, is in contact with a fixed rigid lamina at a single point P. Hence, there is a single constraint force that acts on the workpiece along the line %a and the workpiece thus has two freedoms: a pure translational displacement along the surface of the lamina and a rotation about the contact point P. Assume that the workpiece is to remain in contact with the rigid lamina", + "42) Cambridge Books Online \u00a9 Cambridge University Press, 2009Core terms of use, available at https://www.cambridge.org/core/terms. https://doi.org/10.1017/CBO9780511530173.007 Downloaded from https://www.cambridge.org/core. Columbia University - Law Library, on 13 Aug 2019 at 15:09:03, subject to the Cambridge Following this, the required joint motions Sif/i, Sfe and 8^ can be computed from a reverse analysis of the 3R manipulator, and using (3.52): = J\"1 8DF. (5.43) Suppose that the tool is in point contact with the ground at P (see Fig. 5.5). The line of action of the normal constraint force is labeled $fl. We need to determine a point F about which the base movable lamina will rotate to control (and even reduce to zero, if required) the constraint or contact force at P. From (5.42): 8DF = -[K]\" 1 (5.44) where 8wa = 8f[ca, sa', ra] T and 8DF = 8(l>\\yF, -xF\\ l]T. The required point of rotation F with coordinates (xF, yF) is determined from (5.44) by the stiffness mapping [K]. A rotation of the movable lamina about F cannot move the tool because the change of force applied to it is acting on the line %a", + " Assume finally that the base movable lamina undergoes an infinitesimal twist about an axis through some point G. This twist can be decomposed into a twist of the workpiece about an axis through point E, together with a twist Cambridge Books Online \u00a9 Cambridge University Press, 2009Core terms of use, available at https://www.cambridge.org/core/terms. https://doi.org/10.1017/CBO9780511530173.007 Downloaded from https://www.cambridge.org/core. Columbia University - Law Library, on 13 Aug 2019 at 15:09:03, subject to the Cambridge about an axis through point F (see Fig. 5.5) which simply alters the contact force along $fl, and 8DG = 8DF + 8DE. (5.45) The line $g that joins points F and G intersects $a at point E. The workpiece twists an amount 8DE (= 8DG \u2014 8DF) relative to the ground. Point G can lie anywhere in the plane. This means that a pencil of lines can be drawn through point F, and there are corresponding points of intersections E on the line $a in the range \u2014 <\u00bb to +o\u00b0 (see Fig. 5.7). If a point G is selected, such that the line FG is parallel to Sa9 then point E lies at infinity and the motion of the workpiece is a pure displacement along Cambridge Books Online \u00a9 Cambridge University Press, 2009Core terms of use, available at https://www" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002881_2013.15630-Figure3-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002881_2013.15630-Figure3-1.png", + "caption": "Figure 3. Test run for evaluation of actual power transmission efficiency.", + "texts": [ + " For evaluation of the actual power transmission efficiency of the test tractor in transport operations, the torques and speeds were measured as previously described. The test tractor towed a trailer carrying a load of 2 t uphill on a farm road inclined 2 to 3 degrees using the high-first (H1) transmission gear. The engine speed was adjusted to the rated speed before starting. Once started, the engine speed was decreased to 2550 rpm, resulting in a forward velocity of 9.71 km/h during the test runs. The engine torque at the rated speed was estimated to be 112 N.m. Figure 3 shows the test tractor ready for torque and speed measurements. Figure 4 shows the time histories of engine speed, forward velocity of the tractor, transmission input shaft, and wheel axle torques measured for 15 s after the start of the engine. As expected, the engine speed decreased rapidly to 1900 rpm as the tractor accelerated and increased again to a constant level of 2550 rpm in 3 s. The forward velocity, which was determined based on the measured speed of the wheel axle, also varied accordingly from a high of 11 km/h to a low of 7 km/h at the beginning of the time period, and then settled at a relatively constant value of 9" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001712_0043-1648(83)90299-5-Figure10-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001712_0043-1648(83)90299-5-Figure10-1.png", + "caption": "Fig. 10. Hysteresis loop under starting rolling friction.", + "texts": [ + " By contrast, when the contact width is large and the pressure is low, the moment is large. Therefore it may be supposed that the change in rolling friction is closely related to the contact width. Furthermore, although it is well known that a large surface roughness gives a large contact width [10], it can be explained in an identical way from the change hi the contact, width that the surface roughness affects the rolling friction. The hysteresis loop in the region of the starting rolling displacement is Lndicated in Fig. 10 and the size of the hysteresis loop depends on the rolling distance. As shown in Fig. 11, the calculated hysteresis loop formed about Fig. 11. Diagram of the hysteresis loop and the contact model. t he cen t re is d i f fe ren t f r o m pa t te rns of hysteresis loops o b t a i n e d f rom an expe r imen t . However , if t h e roll ing d is tance is k e p t cons tan t , the hysteresis loop always shows t he same fo rm regardless of its posi t ion. Thus the behaviour of t he hysteresis loop can be invest igated w i t h o u t the fo rma t ion of a hysteresis loop about the centre" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000175_0016-0032(95)00056-9-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000175_0016-0032(95)00056-9-Figure1-1.png", + "caption": "FIG 1 1.", + "texts": [], + "surrounding_texts": [ + "We re-present Fig. 8 as Fig. 11, now emphasising the joint angles, in conformity with the nomenclature and senses of traverse already defined. (The joint angles could be applied equally to the planar network of Fig. 10. They would take different Vol. 332B, No, 6, pp. 657~79, 1995 Printed in Great Britain, All rights reserved 667" + ] + }, + { + "image_filename": "designv11_65_0003475_elan.200403085-Figure5-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003475_elan.200403085-Figure5-1.png", + "caption": "Fig. 5. Layout of the screen-printed disposable strip sensor (typical dimensions were 0.5 cm by 3 cm).", + "texts": [ + " It could be expected that were such systems employed within an authentic matrix (physiological or pharmaceutical), electrode fouling or passivation at the carbon indicating electrode would impact considerably on the analytical performance. The transfer of the technology to a disposable system could be envisaged as the base electrode substrates have been exploited in a variety of sensing applications. These are invariably used in an amperometric/voltammetric detection mode [15 \u2013 18] and as yet, the suitability of such platforms for potentiometric methods is unproven. The basic schematic of the sensing strips used in this work are shown in Figure 5 and consist of conventional screen printed carbon (indicating) and silver \u2013 silver chloride paste (refer- Electroanalysis 2005, 17, No. 3 2005 WILEY-VCH Verlag GmbH & Co. KGaA, Weinheim ence) electrodes [19]. The response obtained at these substrates to 1 mM NQ before and after the addition of various aliquots of Captopril (a common thiol drug for treating blood pressure and heart failure) are shown in Figure 6A. The profile is similar to that observed with glutathione with sharp response times that quickly stabilize (typically within 20 \u2013 30 s) and allow the extraction of a consistent analytical signal which can be related to the Log of the analyte concentration (Figure 6B)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003455_jmes_jour_1980_022_045_02-Figure14-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003455_jmes_jour_1980_022_045_02-Figure14-1.png", + "caption": "Fig. 14. The general case of folding along a curved line", + "texts": [ + " 6(a), ds= R do, /3 is given by (16) and W I - ~ M , \\: 2 tan (: cos2 8 ) R d0 where (21) When Rlr and 8\": are given, I(%*) requires to be calculated numerically. 4.1 The general case of folding on a cylinder In the general case of folding along a curved line, both R and 0 vary along it. Assume that the equation of the folding line in the initially flat sheet is y = y ( x ) for -X 8 20 ::J c;, ~ 15 CI> 1: \u00ab 10 5 a) o o 0 0 0 o o 30 60 90 120 150 180 210 240 270 300 Time (min) 55 T---------------------------------------------~ 50 45 40 ~ 35 ~ 130 c.i c 8 25 CI> ., o g 20 c;, ., g 15 c CI> > 10 5 o o o O+---~----r---~--~----~--_r--~----~--~--~ o 30 60 90 120 150 180 210 240 270 300 b) Time (min) __ plasma o dialysate AC-100 -- plasma o dialysate AC-100 73 Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000045_1.48033-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000045_1.48033-Figure1-1.png", + "caption": "Figure. 1. Principle of dynamic beam based alignment", + "texts": [ + " Using a modulation of the quadrupole strength was first used in LEP in 1992 and presented on the Third Workshop on LEP Performance (4). The measurements are carried out during normal luminosity runs. The modulation of the magnetic field is done by modulating the current of the magnets at fixed frequencies between 0.8 and 17 Hz. The relative change of the magnetic field is of the order of 10~4. The resulting oscillation amplitude of the beam is measured by calculating the Fourier spectrum of a directional coupler signal (Fig. 1). This allows simultaneous excitation and amplitude measurements of several magnets with different frequencies (Fig. 3). These measurements are recorded for several hours, and the offset is extracted by evaluating the minimum response and comparing it to the nearby position monitor. The magnetic field of the quadrupoles is changed in two different ways. The quadrupoles near to the IPs have their own power supply thus they can be modulated by changing the current in the power supply. All other quadrupoles are powered with one power supply for at least two quadrupoles" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000345_1.580546-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000345_1.580546-Figure1-1.png", + "caption": "FIG. 1. Developed wafer-transfer robot.", + "texts": [ + " As the MBE instrument tends to be large scale, a multichamber type instrument that has several reaction chambers is often used in the most recent manufacturing lines. As a wafer transfer system used in recent MBE instruments should be able to be used in an ultrahigh vacuum, generate few particles, and occupy little space, we developed a polar coordinate center robot for a multichamber, which consists of spring arms made of sheet springs and rigid plates and has magnet coupling as an introduction system into the ultrahigh vacuum. This article describes the construction of the spring arms, of the magnet coupling, and the performance of the developed robot. Figure 1 shows the developed wafer-transfer robot. A pair of spring arms made of sheet springs is attached to each revolving rigid arm on the vacuum fringes, which is set in the ultrahigh vacuum region. A pair of revolving rigid arms can be operated independently from the atmospheric region. A holder attached to the top of both arms can position the wafer. When the spring arms move in the same direction, the holder rotates around the central axis. The spring arms separate ~one moves clockwise, the other moves anticlockwise", + " 2~d!#. ~6! The holder with the wafer is moved into chamber A by the linear forward movement of the arm @Fig. 2~e!#. ~7! The wafer on the holder is moved onto the table in chamber A. ~8! The holder is moved out of chamber A by the linear retraction of the arm @Fig. 2~f!#. By the series of repetitive movements described above, a wafer can be moved to any chamber that is attached around the preparation chamber. The spring arm consists of two sheet springs on the upper and lower side, as shown in Fig. 1. It has the following features: ~1! It has strong rigidity in the upper and lower directions. ~2! The operational revolution torque is lower than that of transfer used by the magnet coupling. ~3! The maximum stress generated by the change in shape of the sheet spring is lower than the maximum permissible stress.a!Electronic mail: Masafumi-crl.Kanetomo@c-net3.crl.hitachi.co.jp 1385 1385J. Vac. Sci. Technol. A 15(3), May/Jun 1997 0734-2101/97/15(3)/1385/4/$10.00 \u00a91997 American Vacuum Society Redistribution subject to AVS license or copyright; see http://scitation" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000722_910121-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000722_910121-Figure1-1.png", + "caption": "Figure 1 Collision Interface Geometry [ I ]", + "texts": [], + "surrounding_texts": [ + "T h e OLDMISS computer program is used to cst lmate the AV's of vehicles in a collision, when the damage to one of t he vehicles is unknown. T h e program estimates the energy dissipated in thc s t ruc ture of t he missing vehicle based on the knc~wledgeof the generics.tructura1 characteris t ics of :~utomobiles , and cer ta in assumptions.\nOLDMISS has been found to be less accura te fo r certain impact configurat ions, especially s ide impacts. Th i s paper examines the possible sources of errors in the OLDMISS algorithm.\nThis paper also detai ls a n a l te rna te algori thm fo r est imating the energy absorbed by the missing vehicle. This proposed method bypasses the need fo r est imating the damage profi le on the missing vehicle. I t also takes in to account the presence of induced damage ( fo r side impacts). T h e new a lgor i thm is val idated by reconstruct ing six vehicle-to-vehicle staged impacts, and twelve RIC'SAC tests.\nTHIS paper describes the suggested changes to the OLDMISS computer program. This program is usecl to est imate the AV'S when damage to one of the vehicles is unknown. T h e program estimates the energy dissipated in the s t ruc ture of the missing vehicle based on the knowledge of the generic s t ruc tura l characteris t ics of automobiles, and certain assumptions. The program is used by NHirSA and i ts contractors in the narrow condition:; in which i t is applicable, and the results a r e always carefu l ly evaluated to determine the i r validity. A s a result of its narrow application a n d - tNumbers in parentheses designate references a t end of paper\ninaccuracies, NHTSA does not sell o r provide the OLDMISS program in a n y f o r m t o the public.\nT h e main a s sumpt io r~ is t ha t a force balance must exist a t t he in t e r f ace between the two colliding vehicles a t t hc point of maximum dcformat ion [ I ] t . If the damage to one of t he vehicles is known, t he maximum fol.ce exerted on tha t vehicle can be computed i'rom the known s t ruc tura l properties associated wi th tha t vehicle. An equal a n d opposite fo rce is assumed t o ac t on the vehicle f o r which the damage dimensions a r e unknown. F rom this, t he c rush dimensions of t h e uninspecte d vehicle can be estimated. Once this is donc, thc AV can be computed frorn the damage algori thm of CRASH3.\nT h e relat ionship between the damage on the two vehicles is derived as:\n(see Figure I). T h e vehicle is assumed to behave l ike a l inear dissipat ive spring, wi th of fse t A and s t i f fness B (Figure 2). T h e detai ls of t he above methodology c a n be f o u n d i n [I]. I t should be noted tha t t h c l incar dissipat ive spr ing model is also described by the use of c rush coeff . d o a n d d ,\n[6] as t he intercept a n d slope of vs crush.\nS O U R C E S O F ERROR Ilrl T H E PRESENT OLDMISS ALGORITHM\nUSE O F CORRECTION FACTORS CF,,AND CFZ - O n e of t he possible sources of e r ror in the OLDMISS formula t ion is the use of Cfl and C,,\n25", + "i t . , the \"correction factors\". These arise due to modelling of fr ict ion when the PDOF is not perpendicular to the vehicle surface. The need fo r such correction factors has been questioned [3] and its value limited in the current version of CRASH3. It should be noted that fo r head-on impacts, this factor tends to a value of I . For 90' impacts, the effect of the correction factor is minimal (as Cf, tends to equal CI? fo r this configuration). OLDMISS was re-run w ~ t h C,=C,=l fo r the various crash test:; and the results presented in a later section of thi:; paper (Sec. 6.0). Some improvement in the estimated crush is observed (Table I ) although the errors are still significant.\nUSE O F VEHI'CLE CATEGORIES FOR STIFFNESS PARAMETERS - OLDMISS uses the stiffness categories used in the present CRASH3 [4]. These categories result in the use of approximations for the values of stiffness parameters A, B, and G's. In all likelihood, the use of model-\nspecific crush coeff. (or do and dl) will improve the accuracy of the :ilgorithm. Such a list of do and d l has been generated fo r f ront and rear impacts f rom the Agency's crash test data-base [ 5 ] . A similar list fo r side impacts has also been developed [6] and is inclucled in Appendix A.\nThe effect of using model-specific values of do and d l will be illustl-ated in Section 6 when the proposed algorithm (detailed in Sec. 5) will be compared to the present OLDMISS.\nPRESENCE O F INDUCED CRUSH - One of the major sources of error in the OLDMISS formulation is the modelling of direct and induced crushes. OLDMISS eritimates crushes by applying equal and opposite rorces on the two vehicles. Clearly, this method will work fo r crushes caused by direct contact between the vehicles. However, fo r certain impact modes, specially in side impacts, substantial amounts of induced crush may be present. The current (and proposed) damage measurement protocol calls for the measurement of total crush, i.e., both direct and induced crush. Also, total crush is used in generating the vehicle crush parameters do and d l (Ref [S]). Therefore, the OLDMISS algorithm. which is valid for the direct crush only, results in large errors in generat ion of the missin!: crush profile, and subsequently, in the estima te of the energy absorbed by the missing vehicle.\nIt will be shown in the next section that the crush energy fo r the missing vehicle can be estimated directly, b y p a s i n g the need togenerate the unknown crush prof ,le.\n26", + "PROPOSED METHOD FOR ESTIMATING THE ENERGY ABSORBED\nThis section details a n alternative formulation for estimating the energy absorbed by the missing vehicle. This method underscores the fact that the damage algorithm of CRASH requires the value of the energy absorbed by the two vehicles. This can be generated directly fo r the missing vehicle (as shown below), bypassing the need to estimate the crush profile, and then integrate across that profile. The proposed method therefore avoids the errors due to inaccurate correction factors, possible interpolation of the profile of the known vehicle, and errors due to lack of data about the extent of direct damage. Instead, a simple expression directly relating the energies absorbed by the known vehicle, and that by the missing vehicle is derived below:\nWe model the vehicle to behave like a linear dissipative spring with a n offset (Figure 2) i.e., the force per unit width of the crush is F=/l +(Ex) (2)\nThis is the same model used in CRASH3 program. A l s ~ , let us assume that vehicle I is available and vehicle 2 is the missing vehicle. At maximum cru:;h, assuming force balance, we get A,+(B,x,) =A,+(B,lr,)\nwhrre A,, B, are the stiffness cocff of veh.1 A,, 9, are the st iffness coeff of vch.2 x,, x , a re the displacements (crushes) of vehicles 1 and 2 resp. Rearranging, we get:\nAlso, from Figure 2, we get the energy absorbed per unit width of the direct crush for the vehicle 1 and missing vehicle 2 an\nSubstituting eq.(3) and (4) in Eq.(5), and simplifying, we get\nThe energy absorbed by the direct crush for the known vehicle I is obtai~ned by integrating the quanti ty Ed, across the width of the direct crush as\n'Dl = l c D @ (7)\nFrom Eq.(6) and (7) we gt:t the energy absorbed by the direct crush on the missing vehicle as\nAssuming uniform stirfness along the side of the vehicle, we get\nIt can be shown (Ref [6]) that the quant i ty A and B in Eq.(2) can be related to the intercept (do) and\nslope (d,) of vs. crush as: A ='i,,.d, (9)\nSubstituting in (3.7), we g,et\nTherefore, if the energy absorbed by the direct crush of the available vehicle is known, the corresponding energy absorbetl by the direct crush in the missing vehicle can be calculated by the above equation.\nHowever, the damage algorithm of CRASH3 requires the energy absorbed by the total crush (direct and induced). The relationship between the energy absorbed by the direct crush (ED) and the energy absorbed by the total crush (E) is explored in the next section.\nDIRECT AND TOTAL CRUSH ENERGIES\nThis section examines the correlation between energy absorbed by the direct crush (ED) and that absorbed by the total (direct and induced) crush (E).\nFOR FRONT AND F!EAR - OLDMISS has been shown (in Ref [2]) to be reasonably accurate for collisions involving damage to only the front or the rear of vehicles. This has also been demonstrated in the validation of the proposed algorithm discussed later in this paper. Therefore, it is assumed that fo r the f ront and rear crushes, the total crush energy equals the direct crush energy, 1.e.. E=E, (12)\nWhile this may not be true in some offset and narrower impacts, it does provide a rirst order estimatc of the energy ;~bsorbed. The above" + ] + }, + { + "image_filename": "designv11_65_0001686_018-Figure4-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001686_018-Figure4-1.png", + "caption": "Figure 4. Diagram for aberration evaluated by means of Lorentz\u2019s turns in the PoincarC model. Turns at 0 are simply labelled by their speeds.", + "texts": [ + " Let K, K\u2019 be two inertial frames, and V the velocity of K\u2019 as measured in K, and let a particle move in K with velocity U. What is the velocity U \u2019 of the particle relative to K\u2019? We introduce another inertial frame K\u201c moving with velocity V relative to K, but with the same orientation of spatial axes. It is clear that U \u2019 is determined by the diagram, (u \u2019R\u2019 ) I v, 1) K\u201d -K\u2019 - K L@, 1 1 2 and hence U \u2019 is the \u2018velocity-like part\u2019 of the product (- v, N u , 1). The diagram in the PoincarC model is shown in figure 4. For simplicity, the turn associated with the pure inertial transformation of velocity V is simply labelled as V in the diagram. The direction of V can be read directly, whereas the modulus must be read using the formula (7a) . Ordinarily, the discussion of this phenomenon makes use of the angles 8, 8\u2019 of the motion of the particle with V as measured in K and K\u2019. These angles are also indicated in our diagram. Turns for the Lorentz group 3419 Let an electron describe a plane orbit around a proton" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001696_s0263-8223(01)00168-4-Figure3-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001696_s0263-8223(01)00168-4-Figure3-1.png", + "caption": "Fig. 3. (a) Spring subject to in-plane line load F . (b) Reduced loading system of a semi-spring subject to in-plane line load F .", + "texts": [ + " By simplifying the total strain energy expression, we obtain U \u00bc F 2R3 L D11D66 D2 16 jDj g2 sin a \u00fe p 2 a g 1 2 sin2 a \u00fe cos a \u00fe sin3 a 6 \u00fe p 8 a 4 \u00fe sin 2a 8 ; \u00f04\u00de where g \u00bc sin2 a \u00fe 2 cos a 2 sin a \u00fe p 2a : The displacement in y-direction is given by dy \u00bc oU oF \u00bc FR3\u00f0D11D66 D2 16\u00de LjDj g1; \u00f05\u00de where g1 \u00bc g2 2 sin a\u00f0 \u00fe p 2a\u00de g sin2 a \u00fe 2 cos a \u00fe sin3 a 3 \u00fe p 2a 4 \u00fe sin 2a 4 : By definition, the in-plane compressive spring stiffness is Ky \u00bc F dy \u00bc LjDj R3\u00f0D11D66 D2 16\u00deg1 : \u00f06\u00de 2.1.2. In-plane bending-shear stiffness (Kxy) The composite spring is subject to unidirectional line load \u2018\u2018F \u2019\u2019 in the x-direction at point B as shown in Fig. 3(a). In order to derive the energy expression, a semispring as shown in Fig. 3(b) is considered. The reduced load F =2 and the indeterminate shear force Q are taken into account, while the indeterminate bending moment at B can be shown to vanish by the principle of the least work [14]. The strain energy expression for the complete spring derived from the semi-spring model is U \u00bc L Z n D11D66 D2 16 jDj M2 dn\u00fe Z n A11A66 A2 16 Aj j N 2 dn BE \u00fe Z n D11D66 D2 16 Dj j M2 dn EF \u00fe Z n D11D66 D2 16 jDj M2 dn FC ; \u00f07\u00de where jAj \u00bc A11\u00f0A22A66 A2 26\u00de A12\u00f0A21A66 A16A26\u00de \u00fe A16\u00f0A21A26 A16A22\u00de; where Aij are elements of the extensional stiffness matrix \u00bdA ; M is the bending moment per unit width and N is the in-plane force per unit width" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001311_6.1997-2631-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001311_6.1997-2631-Figure1-1.png", + "caption": "Figure 1 Cross section of the 36\" diameter aspirating face seal", + "texts": [ + "003\") of the rotor surface regardless of the seal diameter. As such, they have a potentially significant performance advantage over conventional labyrinth seals.' In addition, aspirating seals are inherently not prone to wear, owing to their non-contacting nature, so their perfomance is not expected to degrade over time. A 36\" diameter aspirating seal, for application to the GE90 low pressure aft outer seal location, was designed and fabricated by the Stein Seal Company2'3. A cross-section of the design is shown in Figure 1, and major seal components are listed in Table 1. A test plan has been established to evaluate seal performance under a variety of conditions that the seal would be subjected to in the GE90 aircraft engine application. The tests are being executed on a full scale rotary test rig that was originally configured to test 50\" diameter brush and honeycomb seals, but has been modified to Copyright to 1997 by the AIAA 1 American Institute of Aeronautics and Astronautics D ow nl oa de d by M O N A SH U N IV E R SI T Y o n N ov em be r 10 , 2 01 5 | h ttp :// ar c" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002100_pime_proc_1986_200_139_02-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002100_pime_proc_1986_200_139_02-Figure1-1.png", + "caption": "Fig. 1 Geometry of contact angles and deflections", + "texts": [ + " The traction model is purely empirical and has been evolved over a number of years by the author and colleagues to fit experimental data from full-scale bearing tests (7-10, 13, 15). The traction characteristic is typical of a 5 cSt synthetic gas turbine lubricant. In this case the analysis has been approached in the following sections : 1. Contact angle 2. Cage speed 3. Spin power 4. Cagedrag 5. Cage slip 6. Traction equation 2.1 Contact angle For a thrust loaded bearing under quasi-static conditions the loads on each ball are identical. From Fig. 1, Z cot a. = cot ai + $pn - d,w: - (3' FA F* p . = - ' z sin ai F A Po = - z sin a, ( 3 ) (4) where 6* is obtained from standard tables [for example Harris (6)] - (D/2)cf; - 1 + 6,) cos a, I (D/2Ui - 1 ) + 6i ai = cos-' Equations ( 1 ) to (6) are solved by assuming an initial contact angle and working through the equations in the order that they are presented above. The value of ai obtained from equation (6) is then used to start a new iteration. The process is repeated until convergence is achieved" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000040_bfb0020238-Figure4-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000040_bfb0020238-Figure4-1.png", + "caption": "Fig. 4. Sensorimotor configuration of a 2D agent: (a) direction of reeeptor inputs I for' motion detectors s, (b) weighted connections for sensorimotor coupling~ and (~:) direction of motor forces.", + "texts": [ + "01, introduces new genetic material, thus providing further exploration of the search space and preventing the algorithm from premature convergence in a local optimum. Using an elitist approach, the fittest individual of each generation remains unchanged. 4 R e s u l t s In this section we present three flight trajectories to demonstrate the evolved navigation behavior. First we show an agent that is restricted to 2D motion in the horizontal plane. Its sensorimotor configuration is depicted in Fig. 4 and comprises two horizontally oriented motion detectors and two motors. Stable flight and obstacle avoidance behavior (Fig. 5) emerged after 14 generations. A similar agent is used for 3D navigation. It has two horizontal and two vertical motion detectors (Fig. 6). In order to compensate for the simulated gravity in the virtual world, the two motor forces are directed 24 degrees upwards, allowing altitude control in addition to rotation about the vertical axis. The other two rotational body axes are fixed in order to prevent the agent fl'om losing its horizontal orientation" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003067_iros.1998.724635-Figure3-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003067_iros.1998.724635-Figure3-1.png", + "caption": "Fig. 3: Definition of the projection plane", + "texts": [ + " . - ,n ) and DO are shown in [13]. In eq.(ll), dimDB = (Cj\u201d=, 3cj +3r) x 6m and dimDF = (Cj\u201d=, 3 c j +3r) x Cyx, sj , where sj shows the number of joints of finger j Note that, while we dealt with a 3D model, for a 2D model, the skew-symmetric matrix equivalent to the vector product ax is redefined as ( a x ) = [-av a,], and dimDB = (cj\u201d=, 2cj + 2r) x 3m. 4 Kinematics for Enveloping In this section, we consider whether two objects can be lifted up by a simple pushing motion(Fig.1). As shown in Fig.3, the common tangential plane of two objects are defined as n. The plane which is normal to II and tangent to the gravity vector is defined as I?. We consider the motion of the objects projected on I?. The rolling condition is assumed to be satisfied at each contact point. The kinematic relationship between the objects projected on r is shown in Fig.4 where the suffix y denotes a vector on the two dimensional plane r. For simplicity, we assume that an object contacts with one finger at one point or that contact points between an object and fingers are overlapped when they are projected on r" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003842_iros.1992.594476-Figure5-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003842_iros.1992.594476-Figure5-1.png", + "caption": "Fig. 5. Shift Points", + "texts": [ + " The shape of the trajectory with a comer point, a shift point and a reversal point has 3 bend points in the Timeapace shown in Fig.6. Fig.3. Characteristic Points comer point The corner points occur where the robot changes direction and makes a bend in the trajectory. The reversal points are where robot goes back the way it came, and the shift points are where the robot changes the velocity magnitude. These points are identified in time-space, which is 4 dimensional space including the time axis. (Fig.4 & Fig.5) position :versa1 Point axes point toward the eigenvectors of the dispersion matrix based on the teaching points, and the center is the average of them. (Fig.7) The fitting line is the major axis of the ellipsoid body which is the first eignevectorm, of the dispersion matrix A of these teaching points. the first eigenvalue is the length of the major axis. A=E(SXk) here x=x-X X is the center of the teaching points. Similarly the second eigenvectom, points toward, the second axis that is the most scattered direction" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000700_978-3-642-97646-9_13-Figure12.5-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000700_978-3-642-97646-9_13-Figure12.5-1.png", + "caption": "Fig. 12.5. Amplitude of stator currents required for constant rotor flux", + "texts": [ + " The pertinent stator currents are obtained from the equivalent circuit, resulting in I = [1 j WI (1 - a) Ls S (1 + aR )2] I = (1 . T) I -s + RR -mR + J W2 R -mR\u00b7 (12.6) Hence in order to maintain a given rotor flux, the magnitude (RMS- value) of the stator currents must follow a function which depends on the rotor frequency W2 and contains only rotor parameters, (12.7) This is achieved with the help of another function generator which produces a reference according to Eq. (12.7) for the amplitude of the stator currents. The function is plotted in Fig. 12.5 for two values of the rotor time constant; it is symmetrical to W2 = 0, i.e. for operation below and above synchronous speed. Of course, the fact that the function depends on the rotor time constant constitutes also a source of inaccuracy since the rotor resistance may vary considerably with temperature and the inductance with the degree of saturation. Even though this might be alleviated by making the function generator dependent on tempera ture, the realisation would be difficult in practice because a measurement of rotor temperature is not easy with the motor in operation" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002350_robot.1998.680983-Figure3-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002350_robot.1998.680983-Figure3-1.png", + "caption": "Figure 3: A casting manipulator with two links", + "texts": [ + "use abbreviated characters as follows: The paper is organized as follows: Section 2 introduces casting manipulation and presents the dynamics of the casting manipulator model with 2 d.0.f. We propose a method of generating and controlling swing motion by considering the tension of the flexible string. In Section 3, we show the results of a numerical simulation of the proposed method. In Section 4, we describe the results of the experiment and discuss the effectiveness of the proposed method. Conclusions are given in Section 5. 2 Swing Motion Control 2.1 Model of two-link Casting Manipulator The analytical model of the casting manipulator is shown in Figure 3. The model is a two-link planar manipulator with an actuator at joint 1 but no actuator at joint 2. Link 1 is a rigid link, and a string is used as link 2. We consider the negative direction of the Y axis in Figure 3 as the direction of gravity force. In Figure 3, Li is the length of link 1, d , is the distance between joint 1 and the center of gravity of link 1, mi is the mass of link i, Zi is the moment of inertia about the center of gravity of link i, 0 is the angle of joint i, and g is the acceleration of gravity. In order to generate a steady swing without the string becoming slack, the constraint that the tension of the string is positive is given by the following inequality: - n 1 2 ~ S , > ink (1) When inequality (1) is satisfied, the dynamic equations of the system are AII 61 +Ai2 &+Ant 6z2+Aiiz 61 6z+gl=Z AZI 61 +AZZ &+AZII 6i2+gz= 0 (2) (3) where, A I I = ~ I d l z + m z ( L ~ 2 + L z 2 + 2 L ~ Lz Cz)+Ii+Iz Aiz=Azi=mz ( L z * + L I Lz C2)+Z2 A22 =mz Lz' + 1 2 Aiu=-Azii=-ms LI Lz Sz A112=2Aizz gl=mi g di Si,+mz g ( L i Si+LzSiz) g z = m g LZ SU" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000309_cp:19940323-Figure2-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000309_cp:19940323-Figure2-1.png", + "caption": "FIGURE 2 - Cross-sectional coordinate system", + "texts": [], + "surrounding_texts": [ + "A study of helicopter blade vibration control is presented in this work. The blade is modelled by the finite element method and it is considered as a rotating beam undergoing the coupling motions of flapping, lead-lagging, axial stretching and torsion. The blade model also considers a pretwist angle, offset between mass and elastic axes and isotropic material. The finite element matrices are obtained by energy methods and a linearization procedure is applied to the resulting expressions. The linearized aerodynamic loading is calculated for hover and the state-space approach is used to design the control system. The eigenstructure assignment by output feedback is used in the blade reduced model resulting from the application of the expansion method by partial fractions. The simulations for open and closed-loop systems are presented, having exhibited good response qualities, what shows that output feedback is a good alternative for helicopter vibration attenuation.\nINTRODUCTION\nThe high level of vibration undergone by helicopter, is a bad characteristic of this kind of aircraft. The main source of these vibrations is the rotor formed by flexible blades excited by oscillatory aerodynamic and inertial forces. These vibrations are transfered to the helicopter structure and usually create a hostile environment for another devices, crew, or even the passengers. A great quantity of work have been developed to look for a solution to these problems, and a considerable progress has already been reached.\nAmong the ways to reduce vibration problems in helicopters it could be detached the passive and active control. The first one tries to control the helicopter vibratory response through an accurate structural design, e.g., using optimization techniques or installing devices in the rotor or fuselage in order to absorb, or isolate, the vibratory sources (Reichert (1); Loewy (2)). The second one uses the techniques from autoniatic control systems and in Kretz and Larche (3), (2 ) and Friedmann (4). it can be observed a good view of the development of these techniques applied to helicopters.\nFrom automatic control systems applied to helicopter vibration control, the modern control theory has been\nvery used these days. Two are the ways followed by the studies about application of modern control theory on helicopter vibration reduction. The first one considers the rotor-fuselage coupling model as showed in Straub (5). Takahashi and Friedmann (6,7). The second way, showed in Jonhson (8), Robinson and Friedmann (9) and Nguyen and Chopra (lO,ll), considers only the rotor and the control is applied by exciting the blade with higher harmonics of the blade rotational speed, called, Higher Harmonic Control - HI-IC.\nHowever, another kind of control approach has been studied for applications in structures. This is the eigenstructure assignment. Using the state-space representation, the control is obtained by assigning the eigenstructure of a desired closed-loop system and assessing gain matrices. In aeronautics this approach is very applied for stability augmentation and autopilots as showed by Stevens and Lewis (12). For helicopter vibration reduction this approach can be seen in Straub and Warmbrodt (13) who used state feedback, but no other work using the eigenstructure assignment was found\nSince. the eigenstructure assignment technique makes possible to have a better insight of feedback gains assessment and it is more apropnate for multivariable systems, so in this work, the eigenstructure assignment by output feedback for vibration control of a helicopter blade is studied. The blade was considered as a cantilever rotating beam undergoing the bendingtorsion coupling motions. The finite element method was used to compose the blade model, which had its order reduced by the method of expansion of partial fractions. The gain matrix was calculated by the blade reduced model using the output feedback, applied to the complete blade model. The open and closed-loop systems were simulated and the results were analysed.\nMATHEMATICAL MODELING\nThe blade studied here is modeled as a rotating cantilever beam with length R, undergoing the coupling motions of flapping, lead-lagging, axial stretching and torsion and was based on Houlbolt & Brooks (14), Hodges & Dowell (15), and Marques (16). A pretwist angle 8, is adopted in the model, considered null in the blade root and varying linearly through the span. It is also supposed that elastic and mass axes are\nCONTROL\u201994.21-24March 1994,Conference Publication No.389,QIEE 1994.", + "1291\nnoncoincidents\nThe main coordinate systems of the blade model are shown in Figures 1 and 2. The first one shows the main coordmate system x,y and z, that is fixed in the blade root with its origin in the intersection of blade root cross-section and elastic axis. When the blade is not deformed the x axis is exactly coincident with the elastic axis. Figure 1 also shows the deformed blade and elastic displacement U, v and w, in the x, y and z directions, respectively. Fibare 2 shows an arbitrary blade cross-section and its local coordinate system q and 6. 'The torsional defletion 4, due to the blade deformation can also be seen.\nStrain and kinetic energy\nThe strain energy, supposing a rotating beam undergoing axial stress, shear in the lead-lagging plane and in the flapping plane, is given by:\n+ E I ~ (-v\"sinfj, + w \" c o s ~ , ) ~ + (1)\n+GJg\" +Fcvr2 +F,w\"} dx\nwhere EA, Ely, El, and GJ are the axial. lead-lagging, flapping and torsional stiffness, respectively. The tcrm F, is the centrifugal effect and is a function of the miss (m) and the blade rotational speed (n):\nF, = d m x d x\nTo obtain the kinetic energy expression, the approach presented by Magari et a1 (17) IS also used here.\nThe velocity of an arbitrary point i.n the blade crosssections is given by:\nd7 - - ? dt - = ~ x r + r (4)\nwhere:\n- - - c;=ni; ; 7=x, i + y , j + z , k ( 5 )\nThe coordinates (x,,y,,z,) of an arbitrary point in the deformed blade cross-section are the same as shown by references (14) and (16).\nThe kinetic energy is obtained by substituting the equation (4) in the equation ('3) and calculating the double integrals for the blade cross-section areas. Since this expression is too long it will not be presented in this paper.\nAerodynamic loading\nThe steady aerodynamic approach was adopted to yield the expressions of lift (L), drag (D) and aerodynamic moment (M) in the hover condition. Some simplifications were adopted. The first one is neglecting the induced velocity, which yield:; a free air flow velocity parallel to the y axis. The ma l l displacement consideration results in the assumption that the blade cross-section remains parallel to the yz plane. There is no coincidence between mass and elastic axes, but the aerodynamic center is taken at the same point of the elastic axis and cross-section inters1:ction. The profile NACA 0015 was assumed, therefore the aerodynamic and the pressure center of the blade cross-section are the same.\nA blade element of dx length was taken and the corresponding load element was calculated. Considering that the blade elastic displacenients in the free air flow and supposing an operational region of the blade angle of atack, a inatricial expression representing the aerodynamic loading, results as follow!;:", + "1292\n+[v + f l (x + U)]' + w2}dx\n(6)\nwhere a and b are the proportionality factors between lift and drag coefficients due to angle of atack, respectively (as C,= a x a and C,= b x a ); e is the &set between elastic and mass axis; par is the air mass dcnsiN c is the blade cross-section chord; 8, is the command pitch angle; 8, is the nominal value of pitch angle in the operational region (IO' in this work).\nFinite element method\nThe finite element discretization is done in terms of beam elements, with twD nodes at each end having six degrees of freedom: displacements in the x, y and z directions, rotation in the xy, xz planes and in the cmsssection plane. The nodal displacements (generalized coordinates) form the q vector and are related with blade displacements through the following equations:\nU = Hi ( x ) u ~ +H2 (x) U, v = H3 (x) VI + H, (x) V; +H, (x) ~2 +HE (x) V;\nw = H ~ ( x ) w ~ + H ~ ( x ) w \\ + H ~ ( x ) w ~ + H ~ ( X ) W ;\nO=H,~)oO, +Hz~)o92 (7)\nwhere H,(x) through H,(x) are the shape functions given by third degree Hermite polynomials, whch are the same as in (17) and Sivaneri & Chopra (18).\nNow, the matrices Me, G, and Ke, of each finite element can be obtained. Each coefficient mii. gii and 4 for i,j = 1.2 ,... n, is obtained by substituting equation d; in the expressions of the strain and kinetic energy. However, these d i c i e n t s are not linear in q. The linearization accurs adopting the hipothesis of small motions about the equilibrium point (ref. Meirovitch (19)), what yields the expressions of the coefficients as fotlavs:\nmil = -- h~hj I WO\nThe same procedure is done in order to obtain the\nloading vector Q. By substituting equation (7) in equation (6), nonlinear loading expressions &-e obtained and linearized next. This loading vector Q is composed by two parts. A first one depends only of system inputs and a second one depends only of system generalized coordinates.\nThe system matrices are formed (ref. (19)) by superposing each Me, 0, and K,, respectively, and considering the system constraints.\nThe damping effect was put into the model by using the Rayleigh approach described in Clough and Penzien (20). The damping matrix C, is given by C, = a. M + a, K, where a, and a, are arbitrary proportionality factors. In this work a, = 3.4843 and a, = 0.0006, what yields a damping factor of 5 = 0.05.\nThe mathematical model obtained results in the following matricial equation of motion:\nM i + (0 + C, )II + K q = Q(@p ,q,II) (9)\nState-space representation\nThinking about application of control techniques, it is convenient to transform the equation (9) into statespace representation. Then, taking the state vector x(t) = [ qT (dq/dt)T IT, and premultiplying the equation (9) by M-l, it follows that (ref. (19)):\nX(t) = Ax(t ) + B Q (10)\nwhere A is the state matrix and B is the input matrix. The loading vector Q, when represented in state-space form can be written as Q, x(t) + Q2 u(t), where u(t) is the control input vector.\nTaking each part of the loading vector and applying in equation (10) results:\nwhere A, = A + B Q, , B, = B Q, , y(t) is the output vector and C is the output matrix.\nMODEL REDUCTION\nThe high order of original blade model makes necessary the use of a reduction procedure. The technique adopted here is the same as that described by (12). It uses the original system represented in the for of a transfer function matrix written as a partial fraction expansion:" + ] + }, + { + "image_filename": "designv11_65_0002810_1.1711820-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002810_1.1711820-Figure1-1.png", + "caption": "Fig. 1 A modification of Bennett\u2019s representation of the twelve-bar linkage, an example of Sub-type 4", + "texts": [ + "1711820# Among many topics examined in a densely written work @1# appeared Bennett\u2019s treatment of a multiloop planar linkage comprising six anti-parallelograms. Although the network had been included in Dixon\u2019s @2# investigation of \u2018\u2018deformable frames,\u2019\u2019 Bennett recognized the value of a device already applied in a related context by Bricard @3#, who also seems to have introduced the expression \u2018\u2018contre-parallelogramme\u2019\u2019\u2014Bennett continued to use his own term, the more general \u2018\u2018crossed isogram.\u2019\u2019 His illustrative diagram is reproduced here, in modified form, as Fig. 1. While Dixon\u2019s approach had involved fixing one loop, say BA8CD8, so that its companion isogram B8AC8D could move with one degree of mobility, Bennett interpreted the two degrees of mobility overall as a property of the whole network which, according to the Gru\u0308bler/Kutzbach criterion, should have mobility 1. Bennett pointed out that these two degrees of freedom were embedded in the radii of the concentric circles which guide the network\u2019s motion. Such a linkage could be set up on predetermined circles with links of three given or arbitrary lengths, or by laying out three defining links with a common pivot in arbitrary directions, all of these within some limits", + " A separate development has been an analysis @8# of six-bar loops incorporated in the skew type of Bennett\u2019s network, it being found there that all 16 generally possess two degrees of mobility each. Our intention here is to advance the findings of Refs. @1,6,7# in considering the suite of closure modes which can be assumed by the network while displaying the six anti-parallelogrammatic loops. Analytical relationships in linkages are commonly stated by means of displacement-closure equations expressed in terms of joint-angles, such as those shown at pivot A8 in Fig. 1. Because any four-bar in the linkage is determinate once one of its jointangles is known, link-lengths a.b.c having been specified, then a single equation relating ua , ub , uc is sufficient to define the network\u2019s possible configurations for two independent angles. The linkage\u2019s two degrees of mobility are alternatively expressible through the radii r and r8 of the variable circles. Wunderlich @6# employs a polar co-ordinate system for locating nodes on the two peripheries, angles a, b, g defined relative to a chosen node\u2019s position, and we adapt the method for replacing joint-angles, as follows", + " , tan g8 2 5A~r1r82c !~r1c2r8! ~r1r81c !~r81c2r ! . Hence, for example, tan b81g8 2 5 A$~r22@r82b#2!~@r81c#22r2!%1A$~r22@r82c#2!~@r81b#22r2!% A$~@r81b#22r2!~@r81c#22r2!%2A$~r22@r82b#2!~r22@r82c#2!% . It is possible, then, to effectively formulate closure equations in terms of the interplay between r and r8. It should be noted that, unlike the joint-angles, angles a8, b8, g8 are not signed. The foregoing expressions contain hints of special configurations for the network. For the case depicted in Fig. 1 we can now provide the transformation, which is given by 04 by ASME Transactions of the ASME 17 Terms of Use: http://www.asme.org/about-asme/terms-of-use Downloaded F ua5p2~b81g8!, ub5p2~g82a8!, uc5p2~a81b8!, but the dispositions vary from one case to another. As indicated in Ref. @7#, the two degrees of mobility mask distinctions between different closures of the network and so we adopt the ruse introduced there of imposing symmetry on the solutions, where necessary, in order to clarify changeover configurations", + " In the special symmetrical configuration alluded to above which provides a device for separating the sub-types, and in which the c-links are directed radially, we have r85 1 2 S ab c 1c D , r5 1 2 S ab c 2c D . For the solution to be of Sub-type 1, then, we conclude that r82>r21ab . The next transitional configuration allows passage to Sub-type 3, rendered by Fig. 6. The imposition of symmetry for this case requires that the a-links be radial, and it is recorded in Ref. @7# that the displacement-closure equations will then be ua5ub1uc2p and ~Kca1Kab!~ tb 2tc 21KcaKab!1~KcaKab11 !~Kabtb 21Kcatc 2!50. These equations also apply to Sub-type 4, of which an instance is limned in Fig. 1, when symmetry is demanded. The two cases are differentiated by the a*-loop (AC8DB8), for which we can write 466 \u00d5 Vol. 126, MAY 2004 rom: http://mechanicaldesign.asmedigitalcollection.asme.org/ on 08/24/20 ta*56 tbtc1KcaKab Kabtb2Kcatc , the 1 sign associated with Sub-type 3. Transition between the cases is marked by an in-line configuration of loop a*, as exemplified in Fig. 7. We find the algebraic condition for this set of poses to be r822r25bc . In the special symmetrical position, for which the a-links are orientated radially, we have r85 1 2 S a1 bc a D , r5 1 2 S a2 bc a D " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003396_robot.2004.1308808-Figure7-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003396_robot.2004.1308808-Figure7-1.png", + "caption": "Fig. 7. CF-compliant mation, where h e CF has WO edge-face pc\u2018s. (a) side view and (b) front view", + "texts": [], + "surrounding_texts": [ + "Given two feasible neighboring contact states and with CFj a less constrained neighbor of CF,, we search a compliant relaxation motion from C, to Cj. This compliant relaxation motion consists of three separate motions. First, a CF,-compliant motion from C, to an intermediate CF,-compliant configuration C: that contains all ECs that are required by CFj. We call this intermediate configuration, which allows an instantaneous relaxation motion to CFj, the relaration confrguration. Then, an instantaneous relaxation motion from C: in CF, to C; in CFj, and finally a CFjcompliant motion from C; to Cj. A. CFi-compliant motion The first CFS-compliant motion starts from a given CFicompliant configuration Ci, towards a CFi-compliant relaxation configuration C:. Using the vinual controller, we generate this CF,-compliant motion by keeping the ECs that are required by CF,, making the new ECs that are required by the less constrained neighbor CFj , and avoiding collisions at the same time. The new ECs that are required by CFj act like an attractor towards the relaxation configuration C:. Figure 6 shows two possible compliant motions within a face-face CF, from a start configuration C, towards one of the relaxation configurations C:. In this example, the relaxation configuration C: can be any configuration, within the face-face CF, that allows an instantaneous relaxation motion towards the desired edge-face CF. The configuration C, has four vertexface ECs and each valid C: has two edge-edge ECs, as required by the edge-face CF, and in addition to that CL has one, two or three vertex-face ECs, as required by the face-face CE Figure I shows an example of a feasible CF-compliant motion, generated by the virtual controller on the Kuka 361. Fig. 8. Compliant relaxation motion. with {nl, . . . , nr} the camact normals B. Instantaneous relamtion motion The instantaneous relaxation motion is a compliant motion from a given CF,-compliant configuration Cl to a less constrained CFj-compliant neighboring configuration C;. This motion is generated by the virtual controller by keeping all ECs in CFi that are also present in CF, and breaking all other ECs. Figure 8 shows an example of a relaxation motion generated by the virtual controller. C. CFj-compliant motion The CFj-compliant motion stans from the CFj-compliant configuration C;, towards a given CFj-compliant configuration C j . Note that for this CFj-compliant motion, the start and end configurations are given, unlike for the CF,-compliant motion, where more than one possible end configuration C: exists. Using the virtual controller we generate a CFjcompliant motion from C; to Cj by keeping the ECs that are required by CFj and positioning a torsion spring at each joint to move it towards the position that is required by C j . Again, other collisions are avoided during the motion. VI. FEASIBILITY CHECKING In this paragraph we describe the method to verify if the motions generated by the virtual controller define a feasible compliant relaxation motion. By using the approach based on a set of local specifications for each configuration of the held A, we in fact build a potential field for the held A, with a global minimum at the goal configuration. Unfortunately, most potential field consuuctions suffer from the well documented local minima problem [SI, [9]. A \u201ctrap situation\u201d in a local minimum occurs when, due to a balance between all local specifications, the held A does not move. When the held A gets trapped in a local minimum, we use the following approach. First we shut down the goal attractor. This disturbs the balance between the local Specifications and makes the held A move away form the local minimum. To prevent the held A from getting trapped again after the goal attractor is re-activated, we then apply a random force on the held A during a random time interval, to help steering it to a different configuration. We repeat this process until (1) the held A escapes from the local minimum and moves again towards the global minimum, or (2) no way out of the local minimum can be found after a certain number of attempts. In the former case, we consider that the generated compliant relaxation motion is so far feasible for the held A, if the held A reaches a goal configuration, we say the compliant path is feasible. In the latter case, we consider the motion infeasible for the held A. Instead of using a simple goal attractor for the held A, we plan to develop a full-fledge compliant motion planner for the held A to generate an optimized compliant path. The emphasis in this paper, however, is not on finding any optimized compliant path, hut on finding a feasible compliant path to verify that the compliant relaxation motion between two neighboring contact states can he achieved. VII. CONCLUSIONS This paper addresses how to revise a valid contact state graph between two polyhedral objects A and B by adding the constraint that A can only be moved by a manipulator. An effective approach is introduced to check the reachability of contact states and the connections between neighboring states for A held by a manipulator via a virtual compliant controller. A compliant motion path generated based on the revised contact state graph can actually be executed by a manipulator. Our main short term research goal is to complete and extend the implementation of the approach for generating the revised contact state graph. Next we will study compliant motion planning strategies for the kinematic chain of the object A held by a manipulator, based on the contact state graph. We will, consider different optimization criteria in planning and interface the output plan with a low-level controller to execute the planned motion on a real robot manipulator. ACKNOWLEDGMENT All authors gratefully acknowledge the financial support by the U S . National Science Foundation under grant IIS-0328782 and K.U.Leuven's Concerted Research Action GOA/99/04, REFERENCES [ I ] B. Donald. On motion planning with six degrees of freedom: solving the intusection problerus in configuration space. In Pmceedings of the 1985 IEEE Inremotional cm?ferencc on Robotics orid Automotion, pager 536-541. March 1985. 121 E. Gilben. D. Johnson, and S:Keerthi. A fast pmedure for computing the distance between complex objecls in lhreedimendonal space. IEEE b u m 1 of Robotics and Automotion, 4(2):193-203, April 1988. [,I X. li and 1. Xiaa. Planning motions compliant to camplu contact stales. The internotional b u m 1 of Robotic$ Rereorch. 20(6):44&465, June 2001. . [4] X. Ji and 1. Xiao. Random sampling of contact configurations in two-pc contact formations. In K. Lynch B.R. Donald and D. Rus, editors. New Diremiom in Algorithmic and Cowuroriorull Robotics, Boston. 2001. I51 Kluwer. B. H. Kmg. A generalhi potential field approach to obstacle avoidance contmL In The n a j w pors and beyond, Bethlehem, PA, USA. August 1984. International Svmwsium of Robotin Research. 161 T. Lefebvre. H. Bruinikhx, and 1. De Schutter. Polyhedral confact formation modeling and identification for autonomous compliant motion. IEEE Trmoctionr on Robotics and Automiion. 19(1):2&41, February 2003. [7] B. 1. McCmgher and H. Asada. The discrete event modeling and uajectory planning of rabotic assembly tacks. ASME Jnumol ofornamic Syslsremr. Measurement. and Conrml. 117(3), 1995. [8] F. Pan and 1. Schimmeb. Effidenl conta~t stale graph genedon for assembly applicarions. In IEEE Int. Con$ Robo~ics and Aulomotion, pages 2592-2598. Taipei, Taiwan, September 2003. [9] E. Rimon and D. E. Kodischek. Exacl mbol navigation using artificial potential functions. IEEE Tromoctions on Robotics and Automation. [lo] 1. Schimmels and M. PesNdn. Adnlittance matrix design for force guided assembly. IEEE Tromactions on Robotics and Automotion, 8(2):213-227, August 1992. Fine motion planning lhrohgh constraint network analysis. In IEEE lnternorionol Symposium on Assembly and Tosk Planning, pages 16&170, Pittsburgh, PA, August 1995. 1121 1. Xiao. Automatic determinaion of lopological contacts in the presence of sensing uncertainties. In Pmceedinss of the 1993 IEEE Int. Con$ Robotics and A~tomtion. pages 65-70, Atlank?, GA. USA, May 1993. 1131 1. Xim and X. Ji. A divide-and-merge approach to automatic generation of cantact s u e s and contact motion planning. In Pmceedings of the 2wO IEEE Int. Con! Robotics and Automotion. San Francisco, CA, 8(5):501-518, October 1992. [ I l l R. H. Srurges and S. Laowattana. USA, April 2000. [I41 1. Xiao and X. Ji. On aulamalic generation of high-level contact state space. Imernotionoi Journal of Robotics Research, 20(7):584406. July 2001. 1151 L. B a n g and 1. Xiao. Derivation of Contact States from Geomemc Models of Objects. In Pmceedings of the IEEE Int. Con! Ammbly a d Task Plonnmg. pager 375-380. Piusburgh. PA, August 1995." + ] + }, + { + "image_filename": "designv11_65_0001136_s0020-7403(98)00106-4-Figure2-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001136_s0020-7403(98)00106-4-Figure2-1.png", + "caption": "Fig. 2. The ship's motion.", + "texts": [ + " A comparison of the results from these mathematical models showed a high degree of agreement and gives the authors con\"dence to propose an amendment to the Code of Practice [8] to cover cases where &&live'' or swinging loads are to be shipped. The present work looks at the e!ects on the lashing loads of introducing the bending of the trailer chassis (i.e. spine) into the torsionally #exible model [4] and re-modelling the distributed mass as six discrete masses. The authors can report from the conclusions of this paper that in this case, the e!ects on the lashing loads is shown to be marginal, and therefore no modi\"cation to the Code would be necessary. Fig. 2 shows the ship's six degrees of freedom and the terminology used to describe them. Of these six degrees of freedom, the movement which gives rise to the largest lashing forces is roll, and this is followed (at some distance) by pitch. As a consequence of this, the most prominent feature of the ship's model is roll. The pitching motion (as it is small) has been modelled as a heave and surge, equal in magnitude to the vertical and horizontal motions of the deck where the trailer is located. Thus, any point on the ship could be reasonably modelled by (a) A vertical linear motion, Zin (i" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001215_pime_proc_1996_210_489_02-Figure9-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001215_pime_proc_1996_210_489_02-Figure9-1.png", + "caption": "Fig. 9 Basic design of the reciprocating Amsler rig", + "texts": [ + " Again, the maximum values were slightly offset from the position of maximum valve lift, which is consistent with previous wear modelling studies (3) and with measurements of the distribution of wear depth around the surfaces of cams from similar systems to the one treated here [for example the results of Purmer and van den Berg (14)]. 5 THE RECIPROCATING AMSLER MACHINE The reciprocating Amsler machine has been fully described by Coy and Dyson (5). Its operating principle is illustrated schematically in Fig. 9. The test block, which simulates the follower, is loaded against a 28 mm diameter test disc with a spring-loaded rolling contact bearing which is free to follow the motion of the block. The disc is driven at a constant speed of rotation, while the block is actuated by the same drive, via a crank linkage system. This ensures that corresponding points on the disc and block surfaces make contact repeatedly at every rotation of the machine, thus reproducing an important feature of cam and pivoted follower contacts" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000529_41.778253-Figure5-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000529_41.778253-Figure5-1.png", + "caption": "Fig. 5. Sinusoidal motor phase excitation oscilloscope and phasor representations.", + "texts": [ + " In the laboratory system from which the results presented in this paper were obtained, the shared switching leg (S1 and S2) was driven by a complementary switching sinusoidal pulsewidth modulation (PWM) scheme in synchronism with the two phase currents, but phase offset, so as to maximize the available voltage across each phase at any given point in time. If the phase-A and B currents are given by and , respectively, then the S1/S2 junction voltage is modulated by the function , i.e., the S1/S2 junction voltage is 180 out of phase with the current flowing into it. This is shown in Fig. 5, where experimental current and voltage waveforms are shown with an equivalent phasor diagram. The six power switches in the experimental drive were driven by a control system which provided sinusoidal modulation of the current in each phase and maintained the correct sinusoidal voltage waveform on the S1/S2 switch junction. A fixed-frequency chopping controller approach using current feedback was used. Each phase was energized at the start of a 50- s chopping cycle, and deenergized when the current feedback indicated that a reference level had been reached" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003629_aim.2003.1225071-Figure8-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003629_aim.2003.1225071-Figure8-1.png", + "caption": "Figure 8. State transition diagram", + "texts": [ + " When an angular velocity determined by X-position of the joystick is different from the reference velocity the reflective force is generated to head the position of the joystick for the position of the reference velocity. The magnitude of the reflective force is set as the maximum value of the force for transparency. SO is zem. .1. I Free Moving Srnre - The user can freely maneuver the Obrroele Defection 5afe -The user feels the reflective force for obmole d e e d o n when the forward velocity ofthe robot The five control states and transition conditions are organized using a state transition diagram as shown in Figure 8. The diagram plays important role in the reflective force navigation control since it determines the reflective force considering user\u2019s input and robot states. The operator can see the display setup of the simulator and maneuver the virtual robot by a force feedback joystick. In the simulation the permissible value of the maximum forward velocity is set as 3.60(lanm) and the sampling period r, is defined as lOO(ms). In Figure 10, the reference and actual forward velocities are described by the solid and dotted l i e s respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001106_elan.1140040109-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001106_elan.1140040109-Figure1-1.png", + "caption": "FIGURE 1. Flow-through silicone rubber-based calcium sensor: (a) rigid polyethylene T-tube, (b) Ag/AgCI reference electrode; (c) CaCI, 5 mmol/L reference solution; (d) impregnated silicone rubber tube.", + "texts": [ + "1 M constant ionic strength. Preparation of a Flou-Zbrougb Calcium Electrode. The silicone tubes were immersed into an impregnating solution for 15 min. The impregnating solution was prepared by dissolving 2 mg of ionophore in 1 mL of a single solvent and, in some cases, adding 100 p1 of o-NPOE and 0.5 mg of KTpCIPB. The solvent was CHCI,, THF, or xylene. Following impregnation, the silicone tubes were removed and the solvent was allowed to evaporate. The tubing was then fitted into a rigid polyethylene T-piece (Figure 1) to obtain a flow-through electrode; a reference solution, 50 mM CaCl,, and an WAgCl internal reference 1040-03971 92183.50 + .25 41 electrode ensured the electrical contact. The solution at the entrance of the flowing cell was grounded, and the tube was suitably shielded. Reference electrode was a flow-through saturated calomel electrode. The flow-through electrodes were evaluated in a flowsystem (Figure 2). Standard and wash solutions were pumped at 0.4 mL min.-', using the peristaltic pump. RESULTS AND DISCUSSION Following the interesting hemocompatibility of potassium sensors [ 21 where the ionophore valinomycin could be embedded directly into a silicone structure without any additive, we tried to obtain similar behavior with a calcium ionophore" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003924_0022-4898(71)90023-1-Figure7-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003924_0022-4898(71)90023-1-Figure7-1.png", + "caption": "FIG. 7.", + "texts": [], + "surrounding_texts": [ + "The variation of model parameters gave the following results. The force varied with the 2.3 power of the depth which is less than the power derived dimensionally, however, it is felt that the influence of the walls and bot tom served to lower the exponent slightly. A linear dependence of force on projected model area was found, and it was shown that the force did not vary with cone angle, cone surface, or backing cylinder changes. Lastly, it was shown that the force did not vary with the model speed within the range of model speeds selected. It is apparent that the sand failed in a wedge, the shape of which is determined by slip-lines, that is disrupted by the moving mode. This size, weight and shape of this idealized wedge is determined. The relation of force to the two pertinent model characteristics, depth and projected model area, can be stated in terms of the displaced wedge. As the model depth is increased so is the volume of the wedge and thus the cubic power of depth is introduced. Also a larger model produces a larger, but not deeper, wedge and the force increases linearly as a result. FORCES ON BODIES DRAGGED THROUGH SAND 37 The design of ear th penet ra t ing equ ipment will depend upon the force necessary to d is rupt the wedge. This force will be de te rmined by the body forces in the wedge (densi ty related) and the angle o f in ternal f r ic t ion of the mater ia l . I f cohesive ma te r i a l is considered, then the force will also depend upon the unconfined compressive s t rength o f the mater ia l displaced. Acknowledgement--This research was supported by an equipment grant from the Science Research Council." + ] + }, + { + "image_filename": "designv11_65_0003429_s0007-8506(07)60688-x-Figure8-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003429_s0007-8506(07)60688-x-Figure8-1.png", + "caption": "Figure 8 : Tooling system for cold sizing of gears", + "texts": [ + " However, the deviations depicted in Figure 4 are averaged ones. Checking the absolute values of the deviation of the finished part, obtained using the corrected die, yields a different picture. Figure 5 shows the absolute deviation of the finished part in 5 section planes A to E perpendicular to To facilitate an experimental comparison of the two strategies to compensate elastic die deflection examined numerically, a tooling system for the cold sizing of straighttoothed gears is designed and built up. The tooling system (Figure 8, right) permits the cold sizing of gears employing both strategies. The interchangeable form-giving tool elements, e. g. the necessary dies, are designed and built theoretical tooth profile (conventional die after tooth profile in accordance with the results of the numerical correction) investigations. Figure 6 shows the gear's mean deviation from the theoretical tooth profile for the three relevant stages of the cold sizing process involving a conventional die. The deviation plots exhibit substantial discrepancies", + " Afterwards, the optimal compression distance of the elastomer is determined, which is essential for the generation of the correct pressure for the compensation of the elastic die deflection. Finally, a correction of the die cavity's profile takes place, in order to compensate the minute residual deflection of the die due to the upsetting of the die's teeth between the workpiece and the elastomer. For the application of the necessary compressive force on the elastomer ring, the mechanism represented in Figure 8, left, is integrated into the tooling system. Most important part of the construction is the tubular compression punch, which, in the first place, provides the transmission of forces between the press ram and the elastomer ring and, thus, allows for applying the necessary compressive force. For the sizing process subject to this study the average profile deviations of the sized gears can be reduced substantially with the help of the concept of active 6 SUMMARY deflection compensation. The investigations subject to this paper have proved the tooling concept of active deflection compensation to be a practical way to compensate elastic die deflection and its negative effects in cold sizing operations" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000332_jsvi.1996.0786-Figure3-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000332_jsvi.1996.0786-Figure3-1.png", + "caption": "Figure 3. The accelerometer locations.", + "texts": [], + "surrounding_texts": [ + "The bearing parameters are obtained from the experimentally obtained random response in terms of the linear stiffness parameters v2 n1 and v2 n2 , the non-linear stiffness parameters l1 and l2, and the disc\u2013bearing mass ratios m1 and m2. These parameters are obtained for both the vertical and horizontal directions, with the problem formulation, in the horizontal direction, remaining identical to that in the vertical direction. The joint probability density function p(x1, x2) for a set of displacements (x1i , x2j ) (x1(i+1), x2(j+1)), (x1(i+1) q x1i and x2(j+1) q x2j ), from equation (32) are as follows: p(x1i , x2j )= c2 exp$\u2212zm1m2 sx\u03071sx\u03072 61 2 0v2 11 +v2 n1 \u2212 v4 13 v2 331x2 1i + 1 2 0v2 22 +v2 n2 \u2212 v4 23 v2 331x2 2j +0v2 12 \u2212 v2 13v 2 23 v2 33 1x1ix2j +(v2 n1 l1)g(x1i )+ (v2 n2 l2)g(x2j )7%, (33) p(x1(i+1), x2(j+1))= c2 exp$\u2212zm1m2 sx\u03071sx\u03072 61 2 0v2 11 +v2 n1 \u2212 v4 13 v2 331x2 1(i+1) + 1 2 0v2 22 +v2 n2 \u2212 v4 23 v2 331x2 2(j+1) +0v2 12 \u2212 v2 13v 2 23 v2 33 1x1(i+1)x2(j+1) + (v2 n1 l1)g(x1(i+1))+ (v2 n2 l2)g(x2(j+1))7%. Defining Dx1 = x1(i+1) \u2212 x1i , Dx2 = x2(j+1) \u2212 x2j for small Dx1 and Dx2, one can write p(x1(i+1), x2(j+1))= c2 exp$\u2212zm1m2 sx\u03071sx\u03072 61 2 0v2 11 +v2 n1 \u2212 v4 13 v2 331x2 1i + 1 2 0v2 22 +v2 n2 \u2212 v4 23 v2 331x2 2j +0v2 12 \u2212 v2 13v 2 23 v2 33 1x1ix2j +(v2 n1 l1)g(x1i )+ (v2 n2 l2)g(x2j )7% \u00d7exp$\u2212zm1m2 sx\u03071sx\u03072 60v2 11 +v2 n1 \u2212 v4 13 v2 331(x1iDx1) +0v2 22 +v2 n2 \u2212 v4 23 v2 331(x2jDx2)+0v2 12 \u2212 v2 13v 2 23 v2 33 1(x1iDx2 + x2jDx1) + (v2 n1 l1)G(x1i )Dx+(v2 n2 l2)G(x2j )Dx217%. (34) Combining equations (34) and (33) gives p(x1(i+1), x2(j+1))= p(x1i , x2j )exp$\u2212zm1m2 sx\u03071sx\u03072 61 2 0v2 11 +v2 n1 \u2212 v4 13 v2 331(2x1iDx1) + 1 2 0v2 22 +v2 n2 \u2212 v4 23 v2 331(2x2jDx2)+0v2 12 \u2212 v2 13v 2 23 v2 33 1(x1iDx2 + x2jDx1) + (v2 n1 l1)G(x1i )Dx1 + (v2 n2 l2)G(x2j )Dx27%. (35) For N values of each displacement, x1i , x12, . . . , x1N , and x21, x22, . . . , x2N , equation (35) can be expressed as a set of (N\u22121)2 linear simultaneous algebraic equations, as $ sx\u03071sx\u03072 Dx1Dx2 ln 6 p(x1i , x2j ) p(x1(i+1), x2(j+1)7%6 1 zm1m2v 2 n1 7\u2212$G(x1i ) Dx2 %{l1}\u2212$G(x2j ) Dx1 %6l2v 2 n2 v2 n1 7 The laboratory rig for the experimental illustration of the technique is shown in Figures 2 and 3. The rig consists of a disc centrally mounted on a shaft supported in two ball bearings. The shaft is driven through a flexible coupling by a motor and the vibration signals are picked up (after balancing the rotor) in both the vertical and horizontal directions by accelerometers mounted on both of the bearing housings. The signals from the accelerometers are digitized on a PC/AT after magnification. Typical displacement and velocity signals, in the vertical direction, picked up by the accelerometer are given in Figures 4\u20137. The joint probability density function, p(x1, x2), of the displacements is shown in Figure 8. The bearing parameters estimated from equations (36), with the following set of data, EI=1\u00b703\u00d7108 N mm2, m3 =0\u00b7515 kg, L=250\u00b70 mm, are given in Table 1. Bearing stiffness parameters estimated for the horizontal direction are also given in this table, along with those estimated for the vertical direction. The analysis in the horizontal direction remains similar to that in the vertical direction and the horizontal parameters are obtained from the horizontal displacement and the velocity signals; i.e., (y1, y2, y\u03071, y\u03072) from equations similar to equations (36)." + ] + }, + { + "image_filename": "designv11_65_0000835_ip-cta:19971028-Figure3-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000835_ip-cta:19971028-Figure3-1.png", + "caption": "Fig. 3 Characteristics of rudder limiter", + "texts": [ + ") with ( r , v, 6). Thereafter, once the four hydrodynamicrelated and two control-related coefficients of this mathematical model have been identified online (using a least-squares technique), the position of the vessel can be controlled. A similar control strategy was proposed by Lu et al. [3], but in this case the second-order mathematical model involved ship heading (q) and drift angle (p), i.e. the ship model is based upon the second-order Nomoto model [4], and the rate of track-error (4, see Fig. 3 of [3], is dependent upon -Vsin(B - q), where V is the resultant forward speed of the ship. This ship speed dependent term is linearised subject to smallness assumptions concerning (p - v), and then coupled with a simplified version of the original Nomoto model to provide a relationship between the Laplace transforms of track error d and ship heading q. As changes of rudder 6 must modify q, the final expression for track error is intuitively extended by assuming a relationship between 6 and d in a form similar to that derived between q and d", + " rudder I - - - \\ \\ \\ - - - I I \\ \\ \\ 0 \\ \\ - _ I Fig.2 Schematic diagram of SISO controller The inclusion of the rudder limiter in Fig. 2 is based on the understanding that the command rudder signal generated by the neural network may not be physically realisable, because (a) the difference ISkc ~ Sk-ll is too large to be achieved in one time interval, (b) the actual rate of rudder operation is limited. To represent these physical limitations of rudder operation, we introduce a ramp threshold function, see Fig. 3. The upper and lower limits correspond to the maximum deflections of the rudder to port and starboard, respectively, whereas the slope is the maximum rate at which the rudder can be operated (typically 2.5\" per second). More formally, the ramp threshold function is defined as +La, if 6 ( t ) 2 +Lax -4\" if d ( t ) I -4\" S ( t ) = y . t if IS( t ) l 5 +La, (1) 155 { where y, the maximum rate of change of rudder, is pos- itive, and S, is a value commonly referred to as the saturation level. The learning technique is to be based on the backpropagation method [15]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000024_s0167-8922(08)70489-9-Figure13-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000024_s0167-8922(08)70489-9-Figure13-1.png", + "caption": "Figure 13. Contours of the normal stresses in the axial direction on the inner and bonding faces of the bearing aluminium lining", + "texts": [ + " The shear component is in the axial direction on the circumferential plane. It shows a significant level of the strain in the region where bonding face crack is measured. However, it is confined only to the vicinity of the bearing edge. It is believed that this is one of the sources of the crack generation in the bonding layer. For the cracking in the bonding layer further towards the centre of the bearing, the normal stress in the axial direction is potentially the major contributory factor. Figure 13 shows the normal stresses on the bonding face and on the surface of the bearing lining. the area of high stress corresponds to the observed area of cracking in the bonding layer. The magnitude of these stress is also much higher than the yield limit of the bearing lining material. Should the stresses induce plastic deformation of the bearing, then this may be resisted by the more rigid steel of the backing. To fully understand the resulting complex stress state and the crack generation mechanism, it will be necessary to carry out an elastoplastic analysis. Combining the two mechanisms described above, it is possible to explain the observed areas of cracking. The kinks in the measured outline of the bonding layer cracking appears to mark the boundary between two areas. Towards the bearing edge, only bonding layer cracks occurs. Towards the centre of the bearing, both types of cracking occur. By comparison, it is relatively easy to explain the vertical crack growing towards the bonding face. As can be seen from Figure 13, the maximum axial normal stresses are over 100 MPa through the bearing lining. This can cause compression strain in axial direction, but tensile strain in the circumferential direction. When such a strain becomes excessive, it can then contribute to the form of crack faces normal to thc circumferential direction, which is what is observed in the Sapphire bearings. On the bearing surface the maximum stress is over 10 MPa higher than that in the bonding layer. Should crack appears, it is reasonable therefore to expected that it will start from the bearing surface and grow downword into the bearing lining, which is what was observed in the tested bearing" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001196_s0925-4005(97)00224-4-Figure2-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001196_s0925-4005(97)00224-4-Figure2-1.png", + "caption": "Fig. 2. Experimental setup for in vivo experiments.", + "texts": [ + " Reproducibility is checked by repeating 5 min long cycles. For example for a twig carbon probe for oxygen concentration of 10 and 20% the corresponding current values were respectively 10.691.4 and 17.591.25 nA. 3.3. In 6i6o experiments The microelectrodes which yielded good results in vitro are tested on rat liver. The liver is a convenient organ on which it is easy to perform pO2 measurements, since it is large, easily pierced, highly perfused and oxygenated. New probes are therefore fabricated for in vivo purposes only. Fig. 2 shows the experimental setup. For each oxygen probe, we try 15 various locations on the liver. For each location, the current is measured for 5\u20138 min, in order to let it stabilize. Current responses were also studied when the rat was breathing pure oxygen together with behavioural changes. 4.1. In 6itro results For each microelectrode assumed to be effective during the average 24 h long series of experiments, we collect the starting and ending current values and the calculated current drift per hour (see Table 2)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003576_j.triboint.2004.01.004-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003576_j.triboint.2004.01.004-Figure1-1.png", + "caption": "Fig. 1. Coordinate system.", + "texts": [ + " For example, if the frequency characteristics and the corresponding variation of oil film pressure from a pressure transducer installed at the inner surface of the bearing were detected and monitored, it is expected that the information might Nomenclature A surface area of the bearing (m2) ca speed of sound in air (m/s) cs speed of sound in bearing (m/s) e mass eccentricity of rotor (kg) f frequency (Hz) fpx, fpy fluid film reaction forces in x and y directions (N) h film thickness (m) m mass of rotor (kg) N sound pressure level at outer surface of bearing (dB, dB A) Nb averaged sound pressure level of bearing (dB, dB A) p oil pressure (N/m2) pf oil pressure fluctuation (N/m2) pfi pressure of incident wave at outer surface of bearing (N/m2) pfr pressure of reflected wave at outer surface of bearing (N/m2) pft pressure of transmitted wave at outer surface of bearing (N/m2) pm mean pressure of oil (N/m2) pref reference sound pressure (N/m2) R inner radius of bearing (m) Ro outer radius of bearing (m) T period of steady state response (s) t time (s) W static load of journal bearing (N) x,y,z coordinates (m) Za acoustic impedance of air (kg/m2 s) Zs acoustic impedance of bearing (kg/m2 s) l oil dynamic viscosity (kg/m s) h angular coordinate (rad) qa density of air (kg/m3) qs density of bearing (kg/m3) X rotational angular velocity of journal (rad/s) provide diagnostic information on abnormal phenom- ena of the rotor-bearing system. The purpose of the present paper is to investigate the effects of design parameters on the noise of rotor- bearing systems supported by oil lubricated journal bearings. To do this, the effects of radial clearance and width of bearing, lubricant viscosity, and mass eccen- tricity of the rotor are examined for various rotational speeds of the rotor. The coordinate system of an oil lubricated journal bearing is shown in Fig. 1. It is assumed that the rotor and bearing are circular and rigid, and the rotor moves only in parallel mode. The bearing load is applied in the x direction, and the axial groove, which is located at the top of the bearing, is filled with a lubricant at constant pressure. The equations of motion including rotor imbalance for a rotor-bearing system can be written as: m\u20acx \u00bc meX2cosXt\u00fe fpx \u00feW \u00f01\u00de m\u20acy \u00bc meX2sinXt\u00fe fpy \u00f02\u00de where m is the mass of the rotor, x and y are the coordinates of the journal center, e is the mass eccentricity of the rotor, and W is the static load of the journal bearing" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002503_tdcllm.1993.316227-Figure7-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002503_tdcllm.1993.316227-Figure7-1.png", + "caption": "Fig. 7. Illustration of the manipulator system (ground operation type)", + "texts": [ + " Therefore, noise from the telescopic aerial vehicle must be reduced as far as possible. For this purpose, most Japanese telescopic aerial vehicles are equipped with a noise reduction power unit from which hydraulic pressure is obtained. The on-boom operation type system is provided with two noise reduction power units, one for driving the boom and the other for driving the manipulator and hydraulic generator on the working base. This has reduced noise by about 10 dB. 6. SYSTEM -UP II (GROUND OPERATION TYPE) (1) System overview Figure 7 illustrates the ground operation type manipulator system. In this system, the operator in the cabin installed at the assistant driver's seat of the truck on the ground remotely controls the manipulator on the boom end while observing its motions through the 3-D monitor and supervision monitors. In marked contrast to the on-boom operation type, the ground operation type operator, who is away from what is to be worked on, cannot aid in arm end mounted tool replacement. Although this system is of a non-human-intervention type, i t assures increased safety because the operator does not have to go up to a high working place and can stay away from live parts" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001411_thc-2002-10205-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001411_thc-2002-10205-Figure1-1.png", + "caption": "Fig. 1. Five element biped.", + "texts": [ + " The model presented here emulates the adjustments that individuals with pathologies use when they ambulate. The trends in the resulting gait patterns of the model is contrasted with the ones that arise in pathological human gait. The main objective of the present article is to assess the capability of a relatively simple model in capturing dynamics of pathological human gait by comparing the gait patterns of the model with the ones reported in the clinical literature. In this section we define the main features of the five-link bipedal shown in Fig. 1. We consider gaits of the biped that include the single support phase only (i.e. only one of the lower limbs is on the ground surface at any given time). Although, we allow moments at the ankles, we neglect feet structures and assume point contacts between the lower limbs and the ground (see Table 1 for various dimensions, masses, and moments of inertia). The motion of a biped includes two stages. The first stage is the continuous forward motion during which the biped is pivoted on one limb (stance limb) and the other limb (swing limb) is moving in the forward direction", + " Erect body posture: One of the most important aspects of bipedal locomotion is that the biped should maintain an erect posture during locomotion. This requirement can be achieved for the present system when the net rotation of the upper body is kept to be zero at all times. The condition that yields erect body posture can be written as q1 + q2 + q3 = 0 (2) 2. Overall progression speed: We define the \u201coverall progression speed\u201d as the linear velocity of the center of mass of the upper body in the positive \u00d7 direction (see Fig. 1). A steady progression speed can be maintained by letting x\u03073 = Vp (3) where x\u03073 is the velocity of the center of mass in the x direction and Vp is the desired progression speed. 3. Trajectory of the tip of the swing limb during the single support phase: The motion of the tip of the swing limb (in practice the swing foot) during the single support phase is the dominant factor in the trajectory planning process of a bipedal machine. One can generate various locomotion tasks such as stair climbing, walking on a flat or inclined surface by simply specifying the spatial trajectory of the tip of the swing limb" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003306_978-94-017-0371-0_27-Figure2-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003306_978-94-017-0371-0_27-Figure2-1.png", + "caption": "Figure 2. A body-flXed Cartesian coordinate system {x'. y'. z'l and the inertial Cartesian coordinate system {x. y. z} erectedfor a regular truncated octahedron.", + "texts": [ + " la are described by the product oftwo generators: R, the rotation by n/2 with respect to a diagonal joining two opposite vertices, and F, the flip by n with respect to a median defined by connecting the centers of two facing edges. The octahedral group also describes the congruent motions of both a regular truncated octahedron and a regular truncated octahedral tensegrity module. Three diagonals form an orthogonal triad. Therefore, they are selected as the body fixed Cartesian x'-, y'-, and z'-axes, as weIl as the inertial Cartesian X-, y-, and z axes, as shown in Fig. 2. In the figure, both coordinate frames coincide. However, as R and F are applied, the body-fixed frame {x', y', z'} rotates with the octahedron while the inertial frarne {x, y, z} remains stationary. Let the length of the edges of a regular octahedron be denoted by b. The edge length ofthe truncating square is denoted by a. The edges of the regular octahedron intercept the x'-, y'-, and z'- axes at \u00b1bl..fi. Figure 2 also shows the positive direction of the twist angle with respect to square 1'-2'-3'-4' twisted in the \"clockwise direction\" by a. to become square 1-2-3-4. The rotation R about a fixed vector u = [uJ, U2, U3]T by angle cp is described by using the Euler parameters { eo, el, e2, e3 } as folIows: ; e 0 =COS-, 2 . ; e\u00b7=u\u00b7sm- I I 2 i=l, 2, 3, (2a) e2 e 3- e O e l (2b) 2 2 1 e O +e 3 -2\" The rotation matrix R with respect to the z'-axis through vertex 1 is described by the axis [0, 0, I]T and the rotation angle cjl = 1t/2 with respect to the body-fixed frame {x', y', z'}", + " The matrices Rand F with respect to the body-fixed frame become [ 0 -1 0 J [ 0 0 IJ R= 1 0 0, F= 0 -1 0 . o 0 1 1 0 0 (3) Let the identity of the group be denoted by I. The generators, Rand F, satisfy the following relations: RRRR=I, FF=I, FRFRFR=I, R- 1 =RT , F-1 =FT . (4) The above group operations are expressed by matrix multiplication, i.e., the operators are applied from the right to the left. The graph of the octahedral group will be used: (i) to generate .all nodal coordinates from the coordinate of node 1 of the twisted square 1-2-3-4 in Fig. 2 and (ii) to define the connections of cable elements. Let the coordinates be defined at node I in Fig. 2. By following the lines in Fig. 3a and performing the R or F operations, one can move node 1 to any desired node identified in the original configuration. For example, node 1 can be moved to the node 3 position by RR or RTRT as shown in Fig. 3a. The motion to node 15 is achieved by multiplying RRFRR. It is noted that the number of congruent configurations is the same as the number of nodes of both the truncated octahedron in Fig. 1 b and the truncated octahedral tensegrity module in Fig. lc. In order to take advantage of this coincidence between the nodes of the tensegrity module and the congruent configurations of the octahedral group, node numbers are assigned to the nodes of the graph in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003723_s1474-6670(17)31181-3-Figure6-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003723_s1474-6670(17)31181-3-Figure6-1.png", + "caption": "Fig. 6. Example of a clamping device", + "texts": [ + " APPLICATION: CLAMPING DEVICES The positioning and configuration of clamping devices used in various machines or stations of a production line for car bodies is a current question from automotive engineering. Clamping devices have the job to fix different sheet metal parts quickly to one another at an exact position and with a pre-defined clamping force. After the parts are fixed, they are joined together e.g. by welding. The criteria for such clamping units are: Compact size of clamping device flexibility in use and operation closing speed number of cycles high reproducibility accessibility for (spot) welders In figure 6 we can see two sheet metal components to be welded together. In order to keep the parts in the correct position, they are clamped between a fixed abutment and a movable element (closing element) with two clamping jaws. For quality assurance it makes sense to measure the clamping force in order to detect deviations from the ideal geometry. This can be done by additional sensors (e.g . strain gages, pressure sensors) or by analyzing system variables (e.g. electric current, air pressure) from the drive unit" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003437_0954406042690498-Figure7-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003437_0954406042690498-Figure7-1.png", + "caption": "Fig. 7 Coordinate systems for TCA", + "texts": [ + " e \u00f02\u00de II \u00f036\u00de The orientation of the minor axis is determined by a directional angle a\u00f01\u00de, measured counterclockwise from e \u00f01\u00de I to the minor axis and resolved by the two equations cos\u00f02a\u00f01\u00de\u00de \u00bc g2 g1 cos\u00f02s\u00f012\u00de\u00de g21 2g1g2 cos\u00f02s\u00f012\u00de\u00de \u00fe g22 1=2 \u00f037\u00de sin\u00f02a\u00f01\u00de\u00de \u00bc g1 sin\u00f02s\u00f012\u00de\u00de g21 2g1g2 cos\u00f02s\u00f012\u00de\u00de \u00fe g22 1=2 \u00f038\u00de 4 TOOTH CONTACT ANALYSIS AND GOVERNING EQUATIONS FOR CONTROLLING THE MAGNITUDE OF THE PARABOLIC-LIKE KINEMATIC ERROR AND DIMENSION OF THE MAJOR AXIS OF THE CONTACT ELLIPSE In tooth contact analysis (TCA), simulated meshing and contact of two surfaces in point contact, which provides a contact point at every instant, is performed. TCA can determine the contact paths on mating tooth surfaces, the kinematic errors caused by gear misalignment and the bearing contact as the set of instantaneous contact ellipses [1]. As shown in Fig. 7, the coordinate system S1, rigidly connecting to the pinion tooth surface S1, performs pure rotation about the z1 axis with the parameter f0 1 and drives the coordinate system S2, rigidly connecting to the gear tooth surface S2, to perform pure rotation about the z2 axis with the parameter f0 2. To simulate assembly errors, four parameters Dc, De, Dv and Dh are embedded into the setting of coordinate systems. The parameters Dc,De,Dv and Dh represent the centre distance error, axial error, intersecting angle error and crossing angle error respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003214_01977261.1981.11720839-Figure4-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003214_01977261.1981.11720839-Figure4-1.png", + "caption": "Figure 4.", + "texts": [ + " After this several Folsom projectile point preforms were made by Bradley, a grooved log support was constructed, and at tempts were made to remove channel flakes from the preforms using the antler with chest pressure. The results were minimally acceptable. Channel flakes were remov ed; however, they were short and tended to bounce and form irregularly. Figure 3. It was decided that more continuous and greater pressure was needed. A simple leverage device was devised by the authors which allows the loading of a much greater force with consistent and slow application (Fig. 4). Use of the experimental elk antler tool in a lever system proved effective as a preform fluting tool. Several preforms have been fluted to or almost to the tip using a variety of raw materials, including a tough, coarse grained quartzite. The preform is placed vertical ly in a log support (Fig. 4) that has a slot cut from it through which the channel flake falls upon removal. The elk antler tine tip is placed on the prepared platform at an angle slightly less than 90\u00b0. The lever is fit into a notch in a log buried vertically in the ground in front of the log preform support. The angles are adjusted and the plat form contact checked. When everything is deemed cor rect, slow force is applied downward to the end of the lever. When the force of the antler against the platform produces enough pressure to initiate fracture the chan nel flake is removed" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002835_2.4863-Figure4-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002835_2.4863-Figure4-1.png", + "caption": "Fig. 4 Side view of helicopter\u2019s axis system.", + "texts": [ + "B is used to design a hover-to-forward- ight (FF) transition controller for the following model representing the longitudinalchannel dynamics of an Apache helicopter34 constrained to have no verticalmotion; only longitudinal and pitch rotation motions are allowed: X D X trim C X Px . Px \u00a1 Pxtrim/ C X P\u00b5 . P\u00b5 \u00a1 P\u00b5trim/ C X \u00b1e .\u00b1e \u00a1 \u00b1e;trim/ M D Mtrim C M Px . Px \u00a1 Pxtrim/ C M P\u00b5 . P\u00b5 \u00a1 P\u00b5trim/ C M\u00b1e .\u00b1e \u00a1 \u00b1e;trim/ Rx D X=[m \u00a2 cos.\u00b5 /] \u00a1 g \u00a2 tan.\u00b5/; R\u00b5 D M=IY where Rx , R\u00b5 , and \u00b1e represent the forward acceleration (ft/s2 ), pitch angle acceleration (rad/s2), and longitudinal cyclic input (deg), respectively. X represents the aerodynamic force along the x axis and M represents the pitching moment about the y axis. Figure 4 shows the axis system of the helicopterwith respect to the sideview. Table 1 describes the aerodynamic and physical parameters of the longitudinal channel dynamics model. The parameters X trim , X Px , X P\u00b5 , X \u00b1e , Mtrim , M Px , M P\u00b5 , M\u00b1e , Pxtrim , P\u00b5trim , \u00b1e;trim are functions of Px . The physical constants m and IY have values of 4:5528\u00a3 102 and 3:7409\u00a3 104, respectively.The state vector of the helicoptermodel is [x1 x2 x3 x4]T D [ Px Rx \u00b5 P\u00b5 ]T . It is assumed that the output vector of the model is the same as the state vector" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003287_ias.1989.96681-Figure4-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003287_ias.1989.96681-Figure4-1.png", + "caption": "Figure 4 Voltage error vector plane", + "texts": [ + " The amplitude and the position angle of this voltage error vector are given by I = tan - l (Vlqr /Vldr ) (13) Princide of smce voltage vector selection Figure 3 Space voltage vector Selection based on the Dosition angle of voltage error vector In this study, the transistor inverter shown in Figure 1 is employed and therefore this inverter can generate eight kinds of instantaneous space voltage vectors. An example of the relation between the voltage error vector Vr and the space voltage vectors is shown in Figure 2. The position angle of Vr measured from the real axis becomes (8 + V) when the phase angle between the real axis and the d-axis is 0. Selection based on the amditude of voltage error vector In this method, the voltage error vector plane is divided into three regions as shown in Figure 4. The hysteresis bands, H I and H2, are decided by the experiment. The relation between the amplitude of Vr and the hysteresis band can be classified as follows: E In the region [AI, Vr is considerably large and the current deviations A i l d andlor A i l q are also large as is known from equations (7), (8), (9), and (10). In this case, the induction motor is operated in the transient condition, and these deviations must be rapidly reduced. Therefore, the space voltage vector nearest to Vr should be selected" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002783_vib-48424-Figure5-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002783_vib-48424-Figure5-1.png", + "caption": "Fig. 5 Ball load equilibrium with centrifugal force", + "texts": [ + "8 20.68 35 0.396 4.32 9.4 8.51 125.7 0 45.9 20.68 36 0.503 2.29 10.67 8.51 378.4 0 58.5 20.68 37 12.53 6.1 23.88 22.86 2546.9 1.7578 9445.2 20.68 38 12.24 0 0 0 0 1.7219 9232.8 20.68 nloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 02/01/2016 Ter (a) (b) Fig. 4 (a) Bearing reference coordinate, (b) Inner race groove center reference coordinate Let the vector },{}{ zr T vvv = be motion of a ball center. Then, the equation for a ball including centrifugal force Fc is derived from Fig.5. \u2212 +\u2212 = eeii ceeii z r b QQ FQQ v v m \u03b1\u03b1 \u03b1\u03b1 sinsin coscos && && (17) where Qi,e are contact loads, mb is mass of a ball and \u03b1i,e are contact angles. Neglecting the axial displacement of the outer race let wr be the displacement of the outer race groove center q. Displacements of the groove centers p, q and the ball center in the radial and axial directions are shown in Fig.6. The parameters loi, loe represent the distances between the ball center and race groove centers with no force, and the parameters li, le the distances after imbalance forces are applied" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002360_icec.1995.489164-Figure4-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002360_icec.1995.489164-Figure4-1.png", + "caption": "Fig. 4: Single pole system", + "texts": [ + "Oq-,) using much smaller computation compared to ASAGA[j]. 4.2 Cart-pole problems Pole balancing problem is difficult since the system is nonlinear. The problem becomes harder when we do not have any a priori information about the cartpole system. Two cases of this kind of problems are considered in this section. The first is the balancing of a pole on a cart(sing1e pole problem) and the second is the balancing of two poles on a cart(doub1e pole problem). The double pole problem is mlDre difficult than the single pole problem. Fig. 4 and Fig. 5 show the single pole and the double pole systems respectively. The simulation procedure is; similar to that of [ll] except that we use MGA instead of the evolutionary programming method. The equations of motions for the single pole .. (M+m)gsin6 -cos0[u+mli2 sin61 (10) e = - i j t = (11) where A4 is the mass of the cart and in is tlhe mass of the pole. 1 is the half of the pole length and U is the control force. The equations of motions for the double pole system can be derived from dynamic etquilibrium equations" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000556_63.145144-Figure4-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000556_63.145144-Figure4-1.png", + "caption": "Fig. 4. Schematic representation of the motor.", + "texts": [ + " The system is actually described by four blocks, the supply, the converter, and the regulation being gathered in the same block. This block itself is composed of primitive elements-' 'bricks\"-each representing a functional element (supply, converter, timer, regulation loop, sensor). We have developed a set of models for permanent magnet synchronous machines, dc machines, and induction machines. IV. THE EXPERIMENTAL DEVICE The experimental unit consists of a three-phase, eightpole machine with a power rating of 1.7 kW at the speed of 3000 r/min (Mavilor SE 808 AC servomotor). Fig. 4 is a schematic representation of this disc rotor machine with surface-mounted rare earth magnets. The parameters of the machine are given in Appendix I. The motor is supplied by a three-phase amplifier, modified to allow different current control structure experiments (two or three loops) and to study the influence of imperfections, such as offsets or gain errors. The general organization of the control device is presented in Fig. 5 . The neutral point voltage has been computed from the control variables Experimental curves are obtained with a digital storage oscilloscope and printed by a plotter" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000574_1.2831314-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000574_1.2831314-Figure1-1.png", + "caption": "Fig. 1 Schematic of a liead-disl( interface", + "texts": [ + " In this analysis, however, no rolling and no yawing are assumed. The flows around the rails are assumed to be isolated from each other. Therefore, the analysis can be conducted on only one of the two rails. The performances of both rails are similar, being proportional to the ratio of the speeds of the disk at the respective radii where the rails reside. The contact between the rail and the lubricant may vary from full to partial contact depending upon the operating conditions. The following analysis focuses on the fully flooded condition at steady state. Figure 1 shows a schematic of the head-disk interface. The slider rail has a leading taper angle /3, a pitch length \u20ac], a total length \u20ac, and a width b (not shown). Under steady state conditions, the slider rail is stationary while the disk moves at a tangential velocity u\u201e; the slider rail has a pitch V = VOV*VO.TV'* (1) where 770 is the viscosity at ambient pressure, at room tempera ture, and at zero shear rate; 77*, r]*j, and 77* are the shear thinning, thermal thinning, and piezothickening correction fac tors, respectively. By assuming that the fluid is isoviscous, these correction factors become constants and their functional forms are then not required for the dimensionless parametric study conducted here. Although the viscosity is highly dependent of pressure (Van Alsten and Granick, 1988), the piezothickening effect may be neglected (i.e., 77\u0302 \u00ab! 1) because the applied load in this application is considerably low. By assuming no roll (refer to Fig. 1) and setting the origin of the coordinates at the trailing edge, at the centerline, and at the disk surface, the isoviscous Reynolds equation for steady state sliding can be written as d_ dx dp dx ay \\ dy ^ dh = 677M\u201e \u2014 dx (2) where N o m e n c l a t u r e b = width of slider rail, m fx = friction force, N F, = dimensionless friction force, fj'i'quj) h = film thickness, m /?* = dimensionless film thickness, hit hi\u201e = leading-edge film thickness (ap plied film thickness), m /i,* = dimensionless leading-edge film thickness, hi\u201e/(", + "org/about-asme/terms-of-use On the other hand, for a relatively thin film the pressure builds up significantly in the outlet region. The pressure in the outlet region determines the load capability instead. Thus, a larger pitchlength ratio means a larger load support. As the dimensionless minimum film thickness approaches 1 X 10 '', however, the load becomes less and less sensitive to the pitch-length ratio. The lower overall pressure due to the more side flow in the inlet region that results from the longer pitch-length ratio may starts to offset the larger bearing area to support the load. Figure 1(b) shows that the pressure center moves toward the leading edge in a linear manner as the pitch-length ratio increases from 0.92 to 0.98. Figure 7(c) shows that the friction force is also insensitive to the pitch-length ratio; a larger pitch-length ratio exhibits only an almost unperceivably larger friction force. Com bining the synergistic effects of pitch-length ratio on the load capability and the friction force, Fig. 1(d) shows that, first, at a thicker film the larger pitch-length ratio results in a larger friction coefficient. Second, at a relatively thin film the smaller pitch-length ratio gives a larger friction coefficient. Then the friction coefficient converges for different pitch-length ratios as the dimensionless minimum film thickness approaches 1 X 10\"'. 5 Conclusions The Reynolds equation for the head-disk interface lubricated with a liquid film for near-contact recording was solved in closed forms by assuming no roll, no yaw of the slider, and an isoviscous lubricant" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003702_gt2004-53860-Figure5-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003702_gt2004-53860-Figure5-1.png", + "caption": "Fig. 5 The natural frequencies and mode shapes for the central disk/stinger assembly", + "texts": [ + " Additionally the material had to be stiff enough that, at high frequencies and high amplitude of motion, the deformations of the inner ring/disk assembly would remain small. An FEA analysis was conducted to identify the natural frequencies of the inner ring/disk and stinger assembly. In this FEA model, the boundary condition included elements that contain both stiffness and damping to be representative of the bump foil damper during dynamic testing. The central disk/stinger natural frequencies and mode shapes resulting from the FEA are presented in Fig. 5. Since the operation frequency was chosen in the range of 0 to 400 Hz, note that mode shapes one and two (214.4 and 310.8 Hz) are within the range of operation. It was decided that if the resonant frequencies were causing large error in data reduction, an additional stiffener would be fabricated and initial impact tests would be conducted to identify the resonant frequencies. The stinger was a solid rod made of PH13-8Mo with a 25.4 mm in diameter and 23.038 cm in length. Both ends of the stinger had inside threaded holes for connection to the inner ring/disk assembly and shaker" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002142_s0378-5955(00)00196-9-Figure6-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002142_s0378-5955(00)00196-9-Figure6-1.png", + "caption": "Fig. 6. Schematic presentation of two stereocilia used for calculating the distance between them. Further explanations are given in Appendix A.", + "texts": [ + " 7 we introduced a function x : x N a 1; N a 2; N a 3 MMa MR N a 1 M l N B 12 Mel B 12; N B 12 M MMR N a 2 M l N B 12 HEARES 3564 1-11-00 M l N B 23 Mel B 12; N B 12 Mel B 23; N B 23 M MMR N a 3 M l N B 23 Mel B 23; N B 23 M 8 Solutions of Eq. 7 were found as minima of Eq. 8. The minimisation problem for Eq. 8 was then solved using the Downhill Simplex Multidimensional Method (Nelder and Mead, 1965). It is necessary to calculate the distance between adjacent stereocilia in order to estimate electrostatic interaction between them. Let us introduce the Cartesian frame xO1y as shown in Fig. 6. Because the angle a1 is small we assumed that the distance between stereocilia, h1, is the length of segment ED, which is parallel to O1x. The positive direction for all angles in Fig. 6 is chosen clockwise. One can obtain from Fig. 6: xD H L2sina 2; yD L2cosa 2; xE yDtana 1 L2cosa 2tana 1 9 where (xD, yD) and (xE, yE) are projections of points D and E onto axis x and y, respectively. Hence: h1 xD3xE H L2 sina 23cosa 2tana 1 10 Obviously, when the top of the longest stereocilium is displaced by Nx with corresponding angle Na1 then we have: a 1 N a 1 arcsin N x=L1 sina 1 11 Using the same approach one can \u00a2nd the distance between second and third stereocilia as: h2 H L3 sina 33cosa 3tana 2 12 Assad, J.A., Corey, D.P., 1992. An active motor model for adaptation by vertebrate hair cells" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003902_imece2005-81929-Figure6-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003902_imece2005-81929-Figure6-1.png", + "caption": "Figure 6: Catapult Cantilever Design Project. Top: Catapult design showing the \u201cBlank\u201d students had to design from. Bottom: Finished catapult with cuts made at \u201cX\u2019s\u201d above.", + "texts": [ + " Students utilize the material on solid mechanics and beam bending theory presented in lecture to design a beam that will be able to impart the largest amount of kinetic energy on to the projectile while paying attention to the object\u2019s trajectory. However, students must pay careful attention to design restrictions related to dimensioning and material properties. Their device must fit snugly in the launcher base we provide as well as be completely contained within a specified area of the stock material (Fig. 6). Students must also consider mechanical constraints such as the yield strength and failure mode of the material. A selection of inexpensive plastic materials is provided, including polycarbonate, polyvinyl Downloaded From: http://proceedings.asmedigitalcollection.asme.org/pdfaccess.ashx?u chloride, acrylic, high-density polyethylene, and polypropylene, from which the students must choose based on the provided material property data sheets. The second week of the design contest is focused on developing a fully constrained CAD model in SolidWorks" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002529_00037-1-Figure8-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002529_00037-1-Figure8-1.png", + "caption": "Figure 8 Modified ASTM D 695 compression test fixture and specimen (all dimensions in mm).", + "texts": [ + " It actually does not conform to that ASTM standard, and is not an ASTM standard in itself. It was developed by the Boeing Company in conjunction with Hercules, Inc. in 1979 (Berg and Adams, 1988), and then included in Boeing Specification Support Standard BSS 7260, first issued in 1982 (Boeing BSS 7260, 1988). Thus, it is often called the Boeing Modified ASTM D 695 compression test fixture. It was later adopted by the Suppliers of Advanced Composite Materials Association in 1989 as SACMA Recommended Method SRM 1\u00b188 (SACMA SRM 1\u00b188). The fixture is shown in Figure 8. It incorporates I-shaped lateral supports like the ASTM D 695 fixture (ASTM D 695, 1996), but that is where the similarity ends. The specimen, rather than being an untabbed dogboned specimen, is straight-sided and tabbed. Actually an untabbed straight-sided specimen is used to measure modulus, and the tabbed specimen for determining compressive strength. The specified 4.8mm gage length between tabs is too short to accommodate a strain gage. The gage length is very short to prevent gross buckling since a specimen only 1mm thick is specified", + " Roughened rails are clamped onto the specimen, but the bolts do not pass through the specimen, thus eliminating the need for clearance holes, and the associated preparation cost. The fixture is shown in Figure 11. It essentially eliminates the slipping problem. While less research has been performed on the three-rail shear fixture, the problems are very similar. (iii) Double-notched shear The ASTMD 3846 specimen configuration is shown in Figure 12. The Modified ASTM D 695 compression fixture (see Figure 8) can be used to apply the shear loading. The specimen can also be loaded in tension (Hercules, 1990b). Although an ASTM standard, this test specimen is justifiably criticized because of the severe normal as well as shear stress concentrations induced at the bottoms of the notches. This leads to premature local failures which can then immediately propagate across the entire section between the notches. Although this test method is not used very extensively at present, it was relatively popular for a period of time after the Modified ASTM D 695 compression test fixture first became popular in the mid1980s" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001522_robot.1995.526020-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001522_robot.1995.526020-Figure1-1.png", + "caption": "Figure 1. The Toshiba Direct Drive Arm.", + "texts": [ + " It takes the form feeds forward the desired surface normal force in the onedimensional joint space subspace normal to the surface, and (22;) feeds back the integral of the force error in the one-dimensional joint space subspace normal to the surface. The difference between the IDCF and the P D F controller is tha t the P D F controller omits all model-based plant compensation. The IDCF and P D F controller are identical when the IDCF parameter vector B has value zero 1 3. Force Control Experiments The arm used for these experiments (Figure 1 ) is a three degree of freedom direct-drive arm developed at the Toshiba Corporation for advanced robot control research [5, 91. Each joint is equipped with a direct-drive DC brush motor, and high resolution IO6 count laser optical encoder. The control algorithms were executed on a 40Mhz Motorola 96002 DSP at a frequency of 333Hz. T h e IDCFA adaptive parameters were initialized to zero at the beginning of each run. Coulomb and viscous friction model terms were included in the implementation of IDCF and IDCFA, though omitted from the derivation of Section 2 for clarity [IS]", + " - - -w derivative proportional integral force position posit ion force feed feedback feedback feedback forward T h e P D F controller ( a ) uses simple P D feedback in the joint space subspace tangent t o the rigid surface, (ii) ~ \u2018The well known \u201cCraig/Raibert hybrid controller\u201d &st reported in [lo] is similar in many respects to the \u201cPDF\u201d controller described in Section 2.2. The controllers differ in two respects. First, the system dynamics in [lo] are formulated in workspace coordinates, rather than jointspace coordinates. Second, the controller presentedin [ IO] is presented ad-hoc - without a proof of stability. To the best of our knowledge, an experimental comparison between the similar force controllers which differ only in the coordinate system (workspace or jointspace) has not been performed. - 1847 - of the rigid, flat, aluminum plate (as pictured in Figure 1). The reference trajectory circle diameter was 0.2 meters. The reference trajectory speed was a constant 0.0628 meters/second. The robot t,ool-tip thus completes a complete traversal of the circle in 1 0 seconds. The initial robot position error was about 0.05 meters. The initial robot velocity error was 0.0628 meters/second. Figure 2a (left) shows three graphs of actual IDCFA controller performance. The top graph shows the outline of the robot tool-tip actual and reference position trajectories in the plane of the rigid surface" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003698_978-1-4020-2249-4_35-Figure5-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003698_978-1-4020-2249-4_35-Figure5-1.png", + "caption": "Figure 5. 11M-type singularity with e E int (X) and ~f == ~f , V L. In (b) the wrenches \".,L of (a) are drawn with the lengths of their pitches, to show the linear dependence. In (c) the moment components of all \".,L are coplanar .", + "texts": [], + "surrounding_texts": [ + "(If ll\"r5 II 1l\" 0, the new joint is prismatic.) The new joint velocity is the sum of the velocities it replaces.\nWhen ll\"r5 == 1l\" 0, there is no equivalence. This occurs when 0 E ll\"r5 and the distance, rt, from et to Oz equals rr \u00b1 Zr5' where rr is between Oz and er, and Zr5 is the length of the last leg link (Figure3c).\nSuppose ll\"r5==1l\"0' while the other legs are nonsingular. An 10 occurs only if qL E Crl' Then, there is also RPM and RI. The 5 joints of leg L span a 3-system and the platform has 3 dof. Therefore, the PM has 5 dof (like a closed loop with 8 joints spanning a 3 system) and IIM is present. The configuration also belongs to both or neither of types II and RO depending on the other 3 leg chains. (Indeed, replace leg L with a passive nonsingular leg with the same 3-dof. Then this equivalent PM can have only a singularity of class (RO, II).) Therefore, the configuration is either (RI, RPM, 10, IIM) or with all six types.\nOtherwise, if qL f/. Crl' although leg L is singular, the platform is constrained only by Wo and dim T = 4 (no 10). Either qL E Cr - Crl and we have an RPM singularity, or qL E cf and there is RI. In either case, since there is no 10, there must be an IIM-type singularity. (This follows from the interdependence rules of the singularity types, Zlatanov et al., 1995.) Moreover, if qL E cf and there is no RPM, an RO-type singularity must be present as well. Thus we have at least (RPM, IIM) or (RI, RO, lIM) and each can be augmented with (RO, II) if the other three legs cooperate. Figure 4(a) shows an example of a singularity of class (IIM, RI, RO). Thus, ll\"r5 ==1l\"0 allows the six singularity classes with lIM and without 10, impossible for \"usual\" PMs (Zlatanov et al. , 1994). Below, we generally assume that no ll\"r5 == 1l\" o.", + "of the passive joints screws, dim PL < 4, i.e., dim VL > 2. According to Section 2, this is equivalent to qL E cf for some L.\nEach C4 leg has one passive dof, so the redundant passive freedoms of the mechanism are as many as the C4 legs. Figure 4b represents a configuration with two RPM freedoms.\nProvided that a1l7l\"r5\u00a27I\"0, a necessary condition for OE7I\"r5 is e E 13. Therefore, under this hypothesis, all RPM-type singular configurations have an end-effector in 13 and the workspace interior int (X) is free of them. (As we saw, if the geometry of some leg is such that 7I\"k5 == 71\"0, it is possible to have an RPM-type singularity and e E int (X) .)\nmechanisms is given in Zlatanov et al., 1994. The approach can be applied to the instantaneously equivalent PM. A necessary and sufficient condition for lIM is the existence of four linearly dependant wrenches, \",A , ... , \",D , not all zero, such that \",L E WL n M*, V L.\nIf 7I\"r5 \u00a2 7ro and af2 \"# af3 for all L, the mechanism configuration has an end-effector in ax and int (X) is free from 11M-type singularity. Examples of lIM-type configurations are presented in Figures 4-5.", + "For configurations with e E B, either qL E cf or qL E C.r (7r.r5\u00a27rO) for some L . All these are IO-type singularities with respect to extrusion. If qL E C.r an RPM-type singularity occurs. Else, if qL E cf, there is an RI-type singularity. If qL E C.r}: both RI- and RPM-type are present.\nWhen e E aX - B, qL E Cll u Crl for all legs. Assuming 7r.r5 \u00a2 7ro, such a configuration is an (RI, 10) singularity. Iff af2 = ar3' there are leg singularities with e E int (X). They are (at least) in the 10 and RI types, but never in RPM. Figure 6 shows singularities of class (RI, 10).\nThe analysis in the previous sections shows that if no leg is singular, the only possible singularity class is (RO, II), Figure 7. When 7r.r5 \u00a2 7r \u00b0 and af2 -f; ar3' \"no singular legs\" is equivalent to e E int (X).\nBy using the equivalent PM with 4R legs we obtain the 4 x 4 Jacobian:\n1 A kA A -:An 23' 4 \u00b023 T45\n1 B kB B -:11\u00b023 \" 4 \u00b023\nZ= T45 (1) 1 G kG G\n- -::c; \u00b0 23 ' 4 \u00b023 T45 1 D kD D - -::rr n 23 \u00b7 4 \u00b023 T45\nThe rows of Z represent wrenches in M* . These can be transformed to an equivalent 4-element spanning system with three (or four) pure mo ments. The condition for singularity is obtained as a 3 x 3 determinant, and in case of singularity, the RO-motion screw is derived symbolically." + ] + }, + { + "image_filename": "designv11_65_0002873_s0928-4931(02)00064-4-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002873_s0928-4931(02)00064-4-Figure1-1.png", + "caption": "Fig. 1. Picture of the detector for a dry sample containing 1:1 mixture of BSA and laponite: contour plot of the scattered intensity in arbitrary units going from white to black colors for increasing intensities. The platelets were parallel to the neutron beam and either in the horizontal (a) or vertical (b) orientations. The detector was out of axis, so that the zero angle where the primary beam reaches the detector was close to the right-hand-side of the detector. The detector is protected against damage by the primary beam by means of a beam-stop appearing as a white circle. The sample-to-detector distance was SD= 1.00 m, giving q= 0.48 A\u030a 1 at the left-hand side of the detector in (b). Sketches of the configurations are given at the bottom.", + "texts": [ + " Thus, the structural heterogeneities were found in the direction perpendicular to the faces of the platelets and could be observed when the scattering vector was oriented in this direction. The platelets had to be oriented parallel to the neutron beam which entered the material by its edges. Thus, quartz cuvettes were filled with stacks of platelets oriented in either the vertical or horizontal directions perpendicular to the neutron beam [24]. Strong and anisotropic scattering was then observed as in the typical scattering pattern obtained for a 1:1 dry mixture of BSA and clay shown in Fig. 1. Thus, when the platelets were oriented in the vertical direction, the structural inhomogeneities perpendicular to the faces were found in the horizontal direction and the scattering was observed for horizontal scattering vectors (Fig. 1a). Obviously, when the sample was rotated with respect to the neutron beam axis, the orientation of the maximum scattered intensity also turned around by the same angle. Fig. 1b shows the scattering pattern with a vertical orientation for the structural inhomogeneities that were observed when the platelets were horizontal. Since the scattered intensity is strongly anisotropic, the structural orientation is quite perfect. The scattered intensity in the direction parallel to the faces of the platelets is very low as compared to the perpendicular direction; it decreases slowly as a function of the scattering vector and reaches values close to the incoherent background for q > 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003764_rob.20045-Figure3-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003764_rob.20045-Figure3-1.png", + "caption": "Figure 3. Minimum overload trajectories of last link in example 1.", + "texts": [ + " c 530 90 T Nm, c 6 6 T rad/s. The dimensions (lx ,ly ,lz) and the geometric center of one hexahedral obstacle are (0.4, 0.6, 0.4) and (1.2, 0, 0) m in base co- ordinates. The manipulator moves from ( 30\u00b0 , 30\u00b0) to (30\u00b0 ,30\u00b0) in joint space. The velocities and the accelerations at two end points are zero. As mentioned in Section 4.3, we performed the global search for the minimum-overload trajectories by changing the location of the seed point of the obstacle. Our algorithm produced two local minima as shown in Figure 3, where the motion of link 2 is shown. When we located the seed point in the right lower corner of the obstacle, the motion converged always to the local minimum (A) regardless of the total motion time. When it was located in the right upper corner and the total motion time is around 0.65 s, the motion converged to the local minimum (B). When ignoring the obstacle, the resultant motion is very simple [Figure 3(a)]. However, if we lowered c , the motion resembles the local minimum (A). The minimum times are 0.69, 0.99 and 1.16 s in Figures 3(a), 3(b) and 3(c), respectively. Figure 4 shows the actuator torques (solid lines) and the equivalent torques (dotted lines) during a minimum-time motion, where joint 2 is almost saturated during the entire motion, and joint 1 is saturated in the latter half. The model is a 3-link arm shown in Figure 5, where all joints are revolute pairs around their z-axes and it is the configurations of zero-displacements" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000935_elan.1140090215-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000935_elan.1140090215-Figure1-1.png", + "caption": "Fig. 1. Structure of a glucose biosensor using Langmuir-Blodgett films and a schematic diagram of the measurement system.", + "texts": [ + " The adsorption method of enzymes on the LB film has significant effects on the output current of the sensor. Especially, modification of the LB film is a key factor. In this work, various LB films having different electric charges at a head group were used to study the effect of adsorption of enzymes on the output current of the sensor. We also demonstrated a novel adsorption method of enzymes suitable for the fabrication of the LB film biosensors. Furthermore, the effect of the number of deposited layers on the current response was examined. 2. Experimental to form a desired pattern as shown in Figure 1. The substrate was treated with a silylating agent to make the substrate surface uniformly hydrophobic, since it was thought that coexistence of a hydrophilic region (glass) with a hydrophobic region (gold) on the substrate surface would lower the deposition ability of LB films on the substrate. The substrate was immersed in 0 . 0 2 % ~ . stearyltrichlorosilane (Tokyo Kasei Kogyo Co. Ltd., Tokyo, Japan) solution in a 4: 1 v./v. mixture of toluene and carbontetrachloride for 30 min. Then, the substrate was dried at 80\u00b0C for 30min. LB films with adsorbed enzyme were deposited on the substrate with the procedure described in sec. 2.2. The sensing area (surrounded by the bold line in Fig. 1) amounted to 13Omm\u2019. Lead wires were connected to the electrical terminal areas with an electroconductive silver paste (DOTITE D-550, Fujikura Kasei Co. Ltd., Tochigi, Japan). The connection was reinforced with an epoxy adhesive (Araldite Rapid, CIBA-GEIGY, Switzerland) and covered with a silicone adhesive to avoid penetration of water. Response to glucose was measured by the two-electrode electrochemical method with a home-made potentiostat and noise filter. The sensor was immersed into a measuring cell filled with 90 mL of 10 mM HEPES (Dojindo Laboratories, Kumamoto, Japan) buffer solution (pH 7", + " A standard glucose solution of lOmL was added and mixed with a magnetic stirrer. Hydrogen peroxide, which was produced by the enzymatic reaction between glucose and glucose oxidase (GOD, EC 1.1.3.4, G6125, 24900 units/mg, from Aspergillus sp., Sigma), was electrolyzed at a constant voltage (+0.7 V vs. counter electrode). The oxidation current was recorded with an X-t pen recorder (Type 3063, Yokogawa Hokushin Electric, Tokyo, Japan) [lo]. All measurements were performed at 20\u00b0C. 2.1. Fabrication of an LB Film Biosensor for Glucose Figure 1 shows the structure of an LB film biosensor for glucose and a schematic diagram of the measurement system. LB films with adsorbed enzymes were deposited on a glass substrate with Au electrodes by the following procedures. A 50 nm thick chromium film was deposited by evaporation on a mirror-polished glass substrate (50 mm x 50 mm x 2 mm, Coming 7059, Corning Japan Co. Ltd., Tokyo, Japan) to improve adhesion of the gold film to the glass substrate, and a l5Onm gold film was deposited on the chromium film using an ULVAC evaporating system (EVC-SOOA, Tokyo, Japan)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000584_pime_proc_1995_209_411_02-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000584_pime_proc_1995_209_411_02-Figure1-1.png", + "caption": "Fig. 1 Geometrical arrangement of the tri-taper bearing", + "texts": [ + " These thermocouples had been positioned at a circumferential position corresponding to the perceived location of the minimum film thickness (one at the bearing centre-line and one near to each of the bearing edges). The gearbox and hence the problem bearing is a critical component of the whole production process. If the bearing temperature reaches a pre-set value, then the whole plant is shut down. The bearing has operated under similar conditions for many years, albeit under the close scrutiny of the maintenance engineers. The possibility of requiring increased production rates precipitated a more detailed analysis of the bearing. The geometry of the tri-taper bearing is shown in Fig. 1. The arrangement is similar to that for a tri-lobe journal bearing. Here, there are three lubricant supply grooves equi-spaced around the bearing periphery. Each of the three sections consists of equal lengths of the bearing base circle and a section offset from the base circle such that it merges smoothly with the base circle and provides a predetermined depth at or near to the lubricant supply groove. This geometry is akin to three tapered land composite thrust bearings (Fig. 2) arranged around a bearing circle" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002679_ias.1995.530622-Figure7-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002679_ias.1995.530622-Figure7-1.png", + "caption": "Fig. 7: A cross section of a CWT winding", + "texts": [ + " This is due to the fact that the leakage field between the 2509 primary winding and each of the secondary windings is dominated by the spacing between each individual secondary and its corresponding primary tube. At high frequencies, the proximity effects force the bulk of the return current of each of the secondary windings in their corresponding outer primary tube. As a result, very little leakage coupling exists between the individual secondary windings. This phenomenon can be seen by carrying out a field analysis for one of the windings as shown in Fig. 7. The problem of Fig. 7 can be solved using Maxwell's equations in cylindrical coordinates. The flux density diffusion equation in cylindrical coordinates can be written as, Note here that the @ component is the only non-zero component of the flux density. The solution for (5 ) has the form, where I1 and K 1 are the modified Bessel functions of the first and second kind of order one, respectively. The constant k is defined as, (7) where 6 is the skin depth. The constants ci and cz can be found by using the boundary conditions at ro and r2" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003918_j.jmatprotec.2005.02.163-Figure16-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003918_j.jmatprotec.2005.02.163-Figure16-1.png", + "caption": "Fig. 16. Solution proposal.", + "texts": [ + " The mentioned triangular shape a p t t s has to remove the bar before it is able to enter the inlet. If the plug touches the bar, it starts a rotational movement and it will leave the empty space of the inlet. The centre plane is again symbolised by dotted lines. As soon as the plug reaches its final position, the spring presses the bar into of the inlet, the plug returns to the initial position and the plug is fixed by the swivel bar. The reference plug and two swivel bars to fix it are shown in Fig. 16. If the plug reaches the slanted surface, the swivel bars would move around their rotation axis which is symbolised by the drilling. As mentioned earlier, the plug has to be underneath the surface and this could cause a problem for removing the plug. The simplest solution would be to produce two openings, which allows the operator to grab it. Another way to deal with this problem is to attach a replacement part as shown in Fig. 17. By operating the opening and a little further movement, the slanted surface causes the plug to be lifted leading to better accessibility. llows the usage of the same element for all different kinds of lugs; only the position of the rotation axis has to be modified o the certain plug in order to make sure that the right part of he swivel bar is entering the inlet. The second realisation to achieve the normally closed etup is shown in Fig. 16. Similar to the old system, the plug Fig. 14. Reference plug. The proposed solution described in the former chapters solves the space problem by shifting the source to fix or to open in another level. Despite the amount of space necessary for the air tightness test, all available space can be used to generate the force. That means that another technology with lower power density compared to pneumatic could be used. The usage of the readily available compressed air would be the easiest way to produce the force" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002325_3.25718-Figure3-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002325_3.25718-Figure3-1.png", + "caption": "Fig. 3 Surface pressure measurement technique.", + "texts": [ + " Concurrently, the other scanivalve directed the surface pressure being measured by that tap out through the sting to the pressure transducer and associated recording equipment located outside the tunnel. The gap between the face of the pressure tap seal unit and the inner surface of the shell was sealed by means of an O-ring located on the outer face of the seal unit. The cavity created within the seal unit was open to the pressure acting on the outside surface of the shell when the vent hole was aligned with the tap as illustrated in Fig. 3. Once the vent hole in the spinning shell rotated past this aligned position, the seal caused the cavity to retain the pressure. The cavity eventually assumed a constant pressure with time equal to the pressure acting on the surface of the spinning model at that particular circumferential location. Details of a pressure tap seal unit are shown in Fig. 4. The outer surface of each seal unit was contoured to match the radius of the inner shell surface at that location. The 0.5-in.-diam O-rings were composed of lubricant impregnated rubber and were retained in the circular groove of the seal block by high viscosity silicone oil" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002529_00037-1-Figure7-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002529_00037-1-Figure7-1.png", + "caption": "Figure 7 German modified celanese compression test fixture, German Standard DIN 29 971, 1983 (all dimensions in mm).", + "texts": [ + ", 25mm shorter than the IITRI and Celanese fixtures, demonstrating that the 64mm long tabs commonly used with those fixtures are really longer than necessary (Irion and Adams, 1981; Berg and Adams, 1989; Adams and Odom, 1991). Specimens up to 12.7mm wide and 6mm thick can be accommodated. The fixture weighs 4.5 kg, just about the same as the standard Celanese fixture, and is only about 70% as expensive. Thus, it has become a very popular alternative to the standard Celanese and IITRI compression fixtures (WTF, 2000). Another modification is the German Modified Celanese Fixture (DIN Standard 65 380, 1991), shown in Figure 7. This fixture uses flat wedges like the IITRI, but circular holders like the Celanese. Unfortunately, it incorporates an alignment sleeve like the Celanese also. This fixture has experienced somewhat limited use to date, primarily in western Europe. 5.06.6.4.2 End-loading fixtures By far the most popular end-loading compression test fixture at the present time is the socalled Modified ASTM D 695 compression test fixture. It actually does not conform to that ASTM standard, and is not an ASTM standard in itself" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003342_robot.2003.1241866-Figure4-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003342_robot.2003.1241866-Figure4-1.png", + "caption": "Fig. 4 Meuurcment af 6a, and dp, Fig. 5 Measurement of &yo", + "texts": [ + " This treatment does not affect the calibration accuracy. Consequently, = @, = .tio = 0 and the first three components of (15) become e x o = O - J ~ o @ , e s , = O - J y o @ , e z o = 0 - J z 0 @ (18) Note that only the first two equations are independent as the last one has already been considered in ( I I ) . Equations associated with Sa, and SP, Sa, and Sp, can be calculated by measuring the relative distances h,, = h , - h, and h,, =h, - h , from the x - yplane to three points located on the bottom surface of the end-effector as shown in Fig. 4. Assuming that 1) the z , axis is normal to the bottom surface; 2 ) the projection of the points onto the x - y plane constitutes an equilateral triangle with 0 being the center and one side being parallel to the x axis, it can then he shown that da, = ((a2 -a , )& -al ) ) * i ((a2 -aI)x(u, -1 ))*i W O = J5. 3 . (19) 1('2 - a 1 ) x ( a 3 - u 1 ] 1(.2 - ' I )x(a3 where a2 -a I = - b - i r b b i + h 2 , k r a l -a l =-b&+h,,k h,, = h, -h , , h,, = h, -h , and i , j , k are the unit vectors of the axes of the 0 - xyz " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001164_s0303-2647(97)00052-x-Figure2-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001164_s0303-2647(97)00052-x-Figure2-1.png", + "caption": "Fig. 2. Basic construction used in the simulation model of diffusion-limited accretive growth. A new layer j+1 of tangential and longitudinal elements is constructed on top of the preceding layer j. The thickness of the new layer (the length of a longitudinal element) is determined by the local nutrient gradient and the local radius of curvature.", + "texts": [ + " In this paper, we will restrict attention to the \u2018classical\u2019 numerical approach and to diffusion without convection. In another paper we will discuss the application of the particle-based solvers in diffusion and flow limited growth phenomena (Kaandorp et al., 1996). The basic idea is that we finally hope to be able to compare both approaches and be able to identify which method is the most suitable model in parallel simulations (Sloot et al., 1995). 2.1. Simulation of radiate accreti6e growth on a sequential platform The basic construction applied in the 3D geometrical model is shown in Fig. 2. The surface of the object is tessellated with a pattern of triangles which are again organized in a pattern consisting of mainly hexagons and pentagons. The edges of the triangles (the tangential elements) are of nearly equal size s, while the longitudinal elements vary between 0 and s. In the construction the length l of a new longitudinal element, the edge connecting the vertices Vi, j and Vi, j+1 of two successive layers is determined by the growth function: l=s \u00b7k(c)h \u00b7h(r\u0304, max\u2013cur6) (1) where k(c) is an estimation of the local nutrient gradient and h(r\u0304, max\u2013cur6) the amount of contact with the environment, a function based on the average local radius of curvature r\u0304", + "5 A and depends on the number of iteration steps it\u2013step in the iterative geometric construction. The The tip-splitting in the object C is caused by the function h(r\u0304, max\u2013cur6): as soon as the average local radius of curvature r\u0304 exceeds the maximum allowed value, locally the growth velocity decreases and the old branch splits into new branches. One step in the sequential version of the iterative geometrical construction consists of eight successive functional stages. These distinct stages are listed in Table 1 Stage 1 is the actual geometrical construction as shown in Fig. 2. In stage 2, the amount of contact function h(r\u0304, max\u2013cur6) is determined. In stage 3, triangles are respectively inserted or deleted in the new layer j+1. Insertion of triangles occurs if triangles become too large in the construction. In the simulation, the length of the tangential edge is only allowed to vary slightly around the basic unit s of the system. This unit corresponds to the size of the basic building elements (for example the size of spicules in sponges, corallites in stony-corals (Kaandorp, 1994a)) in the actual organisms. In Fig. 2 for example, layer j+1 in one of the triangles a critical limit is exceeded and the triangle is subdivided into four new ones. Also the reverse situation may occur. For example in triangles between branches, the size of the edges may become smaller in subsequent growth steps. Triangles with edges below a certain critical lower limit are removed from the layer j+1. The insertion and value of the amplitude A is expressed in units s. By selecting different values of the start value start\u2013pt it becomes possible to generate a population of forms which are different realizations for the same parameter setting of the two control parameters (max\u2013cur6, h) in the 3D model" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000700_978-3-642-97646-9_13-Figure12.1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000700_978-3-642-97646-9_13-Figure12.1-1.png", + "caption": "Fig. 12.1. Mechanical model of (a) DC and (b) AC motor", + "texts": [ + " This is caused by the fact that the main flux and the armature current distribution of a DC machine are fixed in space and can be directly and inde pendently controlled while with an AC machine these quantities are strongly interacting and move with respect to the stator as well as the rotor; they are determined by the instantaneous values of the stator currents, two of which represent independent control variables. An additional complication stems from the fact that the rotor currents cannot be measured with ordinary cage rotors. Hence the AC motor is a highly interacting nonlinear multi-variable control plant that kept control engineers puzzling for a long time. The differences in control dynamics of a DC and an AC motor are best ex plained by the simplified mechanical models shown in Fig. 12.1. In Fig. 12.1 a, corresponding to a DC motor with a mechanical load, a disk is driven by the tangential force h acting on a pin P which can be moved in a radial slot by a radial force fR acting against a spring. Caused by velocity-dependent friction, the radial motion of the pin is assumed to be relatively slow. Between this mechanical arrangement and the electric model of a DC ma chine the following analogy holds: IT == armature current, R == main flux, fR == field voltage, R h == electrical torque, R w == induced voltage (e.m.f.), R wh == electrical power. Controlling torque, speed and angular position of the disk is straight-forward if hand fR can be separately chosen; the analogy applies to operation below base speed as well as in the field-weakening range with the limitation being either saturation (length of the slot) or maximum induced voltage (circumfer ential velocity of pin). A similar arrangement applies to the AC motor (Fig. 12.1 b); however, the pin is now driven by three connecting rods that apply forces h, 12, is in three fixed directions spaced by 1200 \u2022 In order to produce a smooth circular motion while at the same time keeping the pin at a given radius, a well coordinated set of alternating forces is required for producing constant radial and tangen tial resultant forces. With the analogy that the orientation of the pin (R, g) W. Leonhard, Control of Electrical Drives \u00a9 Springer-Verlag Berlin Heidelberg 1996 corresponds to the position of the fundamental flux wave and the forces to the stator currents, it is easy to see why an AC motor is so much more difficult to control than a DC motor", + " With a digital realisation, the position controller may be a simple proportional (P)-controller, as is discussed in Sec. 15.2. It would also be pos sible to omit the torque controller and generate iSqRef directly with the speed controller. In the direct axis there is a flux controller, the reference signal of which is reduced above base speed in order to achieve field weakening; the flux reference signal also serves to reduce the torque limit at high speed. When recalling the mechanical model in Fig. 12.1b, the strategy of field orientation is as follows: When the speed controller determines that a certain value of torque is required in order to maintain the commanded speed, a refer ence for the necessary tangential and radial forces is computed which is then, by coordinate transformation, converted to reference quantities for the forces 11, 12, h\u00b7 The transformation must be based on the instantaneous position of the pin; hence the need for flux acquisition in a field orientated control scheme. Clearly, the principle of control in field coordinates looks like a very effective way of decoupling the complex multivariable control structure of the induction machine" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003073_cp:20020168-Figure8-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003073_cp:20020168-Figure8-1.png", + "caption": "Figure 8. Locus of stator flux linkage with dc offset.", + "texts": [ + "experimental results demonstrate that the sensorless drive is capable of working from very low speed to high speed and exhibits very good dynamic and steady state performance. V. THE EFFECT OF OFFSET ERROR AND ITS COMPENSATION A. The EjJect of OjJset Error The stator flux is estimated in DTC by integrating the difference between the input voltage and the voltage drop across the stator resistance. When the stator flux is indirectly estimated from the integration of the hackemf, any dc-offset is also integrated and would eventually lead to a large drifts in the stator flux linkage as shown in figure 8. Operational amplifiers are normally used to amplify the signal from the current and voltage sensors. These sensors have temperature dependent dc-offset. Since the current measurement path contains many analog devices, dc offset is an inevitable problem. The offset error causes error in flux estimation and in turn causes the torque of a motor to oscillate at the stator electrical frequency. Figure 9 shows the estimated speed and its harmonic spectrum. As shown in fast 508 Fourier transform (FFT), the estimated speed has many undesirable harmonic ripples, which are, particularly, one, and two, and six times the stator electrical frequency" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002819_s001700170023-Figure2-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002819_s001700170023-Figure2-1.png", + "caption": "Fig. 2. Decomposing compensation laser beam tracking control.", + "texts": [ + " This work adopts a new design so that the laser positioning system is partly absorbed into the 5- axis CNC system. This 5-axis CNC control system can achieve the laser beam positioning functionality by using a decomposing compensation control algorithm. The advantage of this method is that it can delete the control part (motor and its transfer mechanism) of the laser beam focusing system compared to a \u201cstand-alone\u201d laser beam focus system, which will improve the dynamic response characteristics of the laser beam positioning system. Figure 2 shows the decomposing compensation algorithm. The tracking position can be resolved into the movements of five axes. In Fig. 2, u is the direction of the normal vector of the laser beam. The compensation algorithm is shown as follows: ux sin sin u uy sin cos u uz cos u Where and are the absolute coordinate of the A- and C-axes. ux, uy and uz are the compensation discomposing components in X-, Y- and Z directions, respectively. From this equation, it can be seen that the influence of the position tracking error of the laser beam control system will influence, not only the cutting or welding quality, but also the precision of the contour, as the position tracking error can be divided into the movement of the other five axes", + " In order to design a laser beam position sliding mode variable structure controller, we must consider the control mode of the laser beam control system. Figure 3 is the control diagram of the laser beam focus control system: In Fig. 3, w(s) is the variable tracking height of the workpiece, the ISMC is the sliding mode controller with integral, and ku/( s 1) is the approximate servo control model [16]. The servo controller timer is a variable, and for the laser focus servo system, it is usually in the range 20\u201340 ms. When the sampling time is 4 ms, ku 250. On Fig. 2, 0 is the reference value of distance, and is the actual distance e 0 is the distance error. Choose state variables x1 e, x2 e\u00b7 From the control diagram, the state vector model of the system can be described as: x\u00b71 x2 x\u00b72 1 x2 ku u W\u0308 1 W \u00b7 w is the rising and falling height of workpieces, the undulating workpiece can be assumed to have a sine function w kwsin t, its derivative w\u00b7 kw cos t shows the slope of the rising and falling of the workpiece. Suppose that the cutting speed is not more than 5 m min 1, and the slope angle of the workpiece is less than 15\u00b0, so, wkw 5000 60 tan15\u00b0 18" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003931_isie.2006.296115-Figure3-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003931_isie.2006.296115-Figure3-1.png", + "caption": "Fig. 3. Fuzzy sets definitions.", + "texts": [ + " The main idea is: (i) when S is far away from the nominal sliding surface, then parameter p takes a high value, giving a large weight to the error in position; (ii) when S approaches the nominal sliding surface, the gain is adjusted to a smaller value for a smoother approach; (iii) once S is close on the nominal sliding line, then parameter p takes a high value that corresponds to its nominal value, for convergence along the sliding line. The range and the functions for the input and output fuzzy variables are defined in Fig. 3. Center of gravity method is used for defuzzification. VI. FPGA IMPLEMENTATION OF FuzzY CONTROLLER Fuzzy logic algorithms are easily implemented on computers. However, while custom fuzzy circuits require long development time, they offer higher performance. As a compromise or for validation purpose, FPGA-based systems can reduce development time and operate at high speed. Algorithms are described in languages like VHSIC hardware description language (VHDL) and are verified by simulation. Fig. 4 presents the structure of the fuzzy logic controller for adaptation of parameter p as described in section 5" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002350_robot.1998.680983-Figure11-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002350_robot.1998.680983-Figure11-1.png", + "caption": "Figure 11: Structure of Casting Manipulator", + "texts": [], + "surrounding_texts": [ + "4 Experiments\n4.1 Hardware\nA simple casting manipulator was constructed in order to evaluate the swing motion control method proposed so far. Figure 9 shows an over view of the casting manipulator system. The first link made of aluminum plate is driven by the DD motor attached at its base. On the other end of the first link a force-torque sensor is attached in order to estimate joint angle of the second link, a string for fishing. At this moment the string is attached at the force-torque sensor's measurement plate for swing motion control. In order to throw the gripper by releasing the string a fixture is now being developed. The gripper with four fingers is set at the tip of the string. An optical encoder with the resolution of 81000 pulsehev is installed on the DD motor. The specification of the casting manipulator is given in Table 1.\nThe control system is illustrated in Figure 10. The joint angle of the first link is measured by the encoder, and sent to the DOS-V machine (Pentium) through a 24-bit counter. The torque command for the DD motor control is sent from the DOS-V machine through a 12-bit DA converter and motor driver. The joint angle of the second link is estimated from the output of force-torque sensor. This information is given to the DOS-V machine through the 12-bit AD converter. A Tracking Vision, which can\nmeasure plural target positions within one frame, is used to measure the joint angle, 02. These computers, the DOS-V machine and Tracking Vision, are controlled by VxWorks in order to keep a short constant cycle time. The cycle time of the DD motor control is 1 msec. The cycle time of the Tracking Vision is 33 msec. One workstation SPARC Station is used for the man-machine interface.\n4.2 Swing Control\nFor the preparation of swing motion control, estimation of the joint angle 0 2 was investigated. In the case of using the encoder or potentiometer, friction of joint 2 becomes a serious problem. Then we use the force-torque sensor shown as Figure 1 1 . The joint angle 0 2 is estimated from the force information f, and f,, with the following equations:\ne? = atan2 (Fy -hc , Fx -$w) + n t2", + "m is the total mass of the sensor plate and the stopper. With the compensation of the inertial force and the gravitational force this method gives a fairly good estimation.\nThe swing experiments about generation of constant swing (91=50 degree) from initial stationary state were carried out. This experiments corresponded to the simulation described in the previous Section. The time history of the joint angles 81 and 132 are shown in Figure 12. According to the planning method of Section 2.4 the amplitude of 81 is increased in every swing phase. Finally the joint angle 91 reached the command reference, though the joint angle 82 did not converge to zero. We consider that the viscosity resistance of joint 1 and the torsion of sensor codes prevented the motion of the first link. So then the torque for keeping zero did not generated sufficiently during the swing motion. However, the error of the amplitude of joint 1 remained less than 5 degrees. Then this result indicates that the proposed swing motion control method works well.\n5 Conclusions\nIn this paper, first we showed the process of casting manipulation and discussed the swing motion phase for the first steps. Then we proposed a method of generating the desired swing motion without becoming the string slack from the initial state when the links are hung. We investigated the method of choosing feedback gain by an off-line simulation chart. The method was evaluated by numerical experiments and shown to be effective. We also constructed a two-link casting manipulator and proposed the method of measuring the flexible string by the forcetorque sensor. Finally we showed the experimental results that indicated that the proposed swing motion control method to be effective.\nReferences [1]H. Kobayashi, E. Shimemura, and K. Suzuki, \u201cAn Ultra-Multi\nLink Manipulator\u201d, in Proceedings of the IEEWRSJ International Workshop on Intelligent Robots and Systems\n[2]N. Takanashi et al. \u201cSimulated and Experimental Results of Dual Resolution sensor Based Planning for Hyper-Redundant Manipulators\u201d, in Proceedings of the IEEWRSJ International Workshop on Intelligent Robots and Systems (IROS\u201993),\n[3]0. Khatib, K. Yokoi, K. Chang, D. Ruspini, R. Holmberg, A. Casal, \u201cVehicle/Arm Coordination and Multiple Mobile Manipulator Decentralized Cooperation\u201d in Proceedings of the IEEEJRSJ International Workshop on Intelligent Robots and Systems (IROS\u201996), pp.546-553, 1996. [4]H. Arisumi, T. Kotoku, K.Komoriya, \u201cA study of Casting Manipulation (Swing Motion Control and Planning of Throwing Motion)\u201d, in Proceedings of the IEEWRSJ International Workshop on Intelligent Robots and Systems\n[5]F. Saito, T. Fukuda, F. Ami, \u201cSwing and Locomotion Control for Two-Link Brachiation Robot\u201d, in Proceedings of the IEEE International Conference on Robotics and Automation, pp.\n[6]K. Furuta, M. Yamakita, S. Kobayashi, \u201cSwing Up Control of Inverted Pendulum\u201d, in Proceedings of the IEEE International Conference on Industrial Electronics, Control and Instrumentation, pp. 2193.-2198, 1991.\n[7]M.W. Spong, \u201cSwing Up Control of the Acrobot\u201d, in Proceedings of the IEEE International Conference on Robotics and Automation, pp. 2356-2361, 1994. [8]Y. Hashimoto, T. Tsuchiya, I. Sugioka, T. Mastuda, \u201cTransversal Load-Swing suppression Control of Travelling Crane\u201d, The trans. of RSJ* vol. 11, No.7, pp. 1073-1082. 1993 (in Japanese) [9]T. Sakaguchi and F. Miyazaki, \u201cDynamic Manipulation of Ball-in-Cup Game\u201d, in Proceedings of the IEEE International Conference on Robotics and Automation, pp. 2941-2948, 1994. [IOIY. Nakamura and R. Iwamoto, \u201cSpace multibody structure connected with free joints and its shape control\u201d, in Proceedings of the IEEE Conference on Decision and Control,\n(IROS\u2019gl), pp.173-178, 1991.\npp.636-643, 1993.\n(IROS\u201997). pp.168-174, 1997.\n719-724, 1993.\npp. 3126-3131, 1993." + ] + }, + { + "image_filename": "designv11_65_0003163_iros.1996.570845-Figure6-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003163_iros.1996.570845-Figure6-1.png", + "caption": "Figure 6 Ambiguous contacts", + "texts": [ + " (12) The counterparts of the equations (4)-(6) and (9)-( 10) exist for the ev facets and the four classes of edges. These are derived in [3]. The form of these equations are listed in the Appendix for reference. E (Oi\u2019 Ci> = pi - Oill 4.3 C-surface Selection For every observed configuration oi, we obtain a set of feature pairs FPi which satisfy the contact conditions (1)- (3). From the set FPi we need to find the most likely contact configuration and c-surface corresponding to it. This c-surface can easily be obtained when the contact pairs are well defined as shown in Figure 6(a). But in certain configurations such as the one shown in Figure 6(b), due to the proximity of the object features, not all the contact pairs in the set FPi will correspond to real contacts. The true contact and the c-surface in this case is obtained as follows. Given the feature pair set FP, we generate all combinations of contacts which can correspond to a c-surface. We then project the observed configuration onto these c-surfaces and find the corrected configuration in each case. The csurfaces are mathematical constructs, hence there will al- ways be a projected configuration" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000191_bf01107213-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000191_bf01107213-Figure1-1.png", + "caption": "Figure 1 Schematic configuration of the tensile test specimen.", + "texts": [ + " A highresolution optical microscope was used to identify the intrinsic microstructural features and examine the weld quality that included porosity, hot tearing and extent/depth of penetration. X-ray diffraction using a Siemens diffractometer and CuK~ radiation was also used to identify the microstructures. Tensile and hardness tests were performed on the laser-welded specimens. The tension tests were performed on a closed-loop servohydraulic Universal testing machine at a constant crosshead speed. The test specimen configuration is shown in Fig. 1. The specimens were mechanically polished, cleaned and rinsed in acetone prior to testing. Multiple tests were conducted to ensure consistency in results. A Vicker's microhardness tester was used to determine hardness at various locations along the weld, at a load of 500 g. A scanning electron microscope was used to examine the deformed tensile specimens in order to characterize the predominant fracture mode and the fine scale fracture features. 4. Resul ts and discussion 4.1. Laser we ld ing Two mechanisms, namely \"deep penetration\" and \"conduction\" are involved in laser welding depending on the laser parameters (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003012_1350650011543637-Figure3-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003012_1350650011543637-Figure3-1.png", + "caption": "Fig. 3 Journal bearing with two axial oil grooves", + "texts": [ + " However, knowledge of using the real geometry in the analysis was not available and had to be approximated with a rectangular geometry (Fig. 1). The second geometry to be treated is the thrust bearing with sector-shaped pads and a plane inclined gap profile (Figs 2 and 3). Here, the geometry of the pads is the same as in the analysis presented by Floberg [25], i.e. the pads fill 5R of the periphery 2\u00f0R, which is approximately 80 per cent. Then, it follows that the mean length of the pad is L \u02c6 5R=n where n is the number of pads. The geometry of one pad is shown in Fig. 3 where cylindrical coordinates (j, r) are used. The angle which one pad fills is \u00f5 \u02c6 5=n. Finally, the circular journal bearing with two axial oil J01001 # IMechE 2001 Proc Instn Mech Engrs Vol 215 Part J DESIGN FUNCTIONS FOR HYDRODYNAMIC BEARINGS 407 at NANYANG TECH UNIV LIBRARY on June 2, 2015pij.sagepub.comDownloaded from grooves is analysed. A description of the geometry is given in Fig. 3. One drawback with circular journal bearings is that instability may occur. This has been analysed by, for example, Lundholm [28, 29]. One solution to the stability problem is to replace the journal bearing with radial thrust bearings. Another solution is to use the combined hydrostatic and hydrodynamic bearing treated by Zhang [30]. In the present work, dynamic loads are not considered, but the aim is to derive a function for each design quantity covering a larger interval than that presented in reference [20]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001964_s0022-5193(05)80089-x-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001964_s0022-5193(05)80089-x-Figure1-1.png", + "caption": "FIG. 1. A cross-section through the cellular monolayer along one principal direction in the surface shows the local radius of curvature in this direction, Rt, as well as the parameters b t and a~ from which the ratio x = b~/aj is found. Also shown is the variable h, the local cell height. The middle surface is shown as \"dashed\".", + "texts": [ + " It is assumed now that the curvature of the surface at each point is a function of the morphogen level at that point, and that this is determined once the cell shape as a function of morphogen is known. Although we will refer to cell shape in what follows, it is to be understood that the shape may actually refer to the shape assumed by an aggregate of cells, say five or ten, when this is warranted. Then we think of our \"cel l\" shape in a given co-ordinate square as being determined by three positive numbers, only two of which will turn out to be independent. Reference to Fig. 1 will show these relevant parameters, two of which, x and y, are dimensionless, and the third which is the cell height \" h\" , all three evaluated at the point u, v, via their dependence on morphogen \" m \" at that point. The ratio of the linear dimensions of the basal cell width to the apical cell width in one principal direction in the middle surface, say b~/al, is shown in Fig. 1, and will be denoted as \" x \" , while the ratio in the corresponding orthogonal direction bJa2 will be denoted by \"y\" . The two principal directions in the surface are orthogonal directions along which the two (principal) radii of curvature of the surface (R1 and R2) take on their maximum and minimum values. An elementary geometrical argument then leads us to an expression for the principal radii of curvature R~ and R2 as functions of x, y and h, and thus to expressions for both M O D E L O F M O R P H O G E N E T I C P A T T E R N F O R M A T I O N 553 the Gauss and Mean curvatures of the surface at that point. By reference to Fig. 1, it is seen that, from the definition of the angle h, a ~ / ( R ~ + h / 2 ) = b d ( R l - h / 2 ) , where RI is the principal radius of curvature of the middle surface in the \"1\" direction, so that x = (b~/a~)= (g~ + h / 2 ) / R ~ - h /2 ) , (5) and similarly for the \"2\" direction orthogonal to \"1\" in the surface, y = (b2/a2) = (R2 + h / 2 ) / R 2 - h /2) . (6) Solving eqns (5) and (6) for Rl and R2 gives at once that g~ = (h/2)(1 + x ) / ( 1 - x ) , (7) and R2 = (h/2)(1 + y ) / ( 1 - y ) . (8) This gives at once for the Gauss curvature; K = 1~(RIg,_) = (4/h2)[(1 - x ) ( 1 - y ) ] / [ ( 1 +x) (1 +y) ] " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000153_1.2802446-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000153_1.2802446-Figure1-1.png", + "caption": "Fig. 1 (a) Top and (/>) rear (section A-A) views of the ball rolling on the rim. Variables are similar to those of Holmes (1991). A moving orthogonal right-handed coordinate frame A has unit vectors a\u0302 and as in a vertical plane with 32 in the direction from rim contact point to bail center and the third unit vector as = a, x as horizontal. Two angles 6 and <\u0302 describe the ball position on the rim relative to its position 0 = do < 0, c\u0302 = 0 at the initial contact point /.", + "texts": [ + " In the simulations of the model presented below, we show that this term should not be neglected and we compare a full spin model and a correct no-spin model. Finally the barrier between the initial conditions for roll-in and roll-out trajectories is studied and shown to lead to a set of quasi-equilibrium trajectories which, although not obtainable from the horizontal rolling initial conditions, nevertheless yield large values for roll around angles and contact times. The Three Degree of Freedom Model (Spin Model) Figure 1 shows the coordinate system. Many of the symbols are chosen to be identical to those used by Holmes (1991). The intersection of the horizontal plane and the axis of symmetry of the right circular cylindrical hole of radius R,, is labeled O. Since we are concerned with rolling on the rim, we assume that the surface of the smooth spherical ball (with mass m, radius Rb, and centroidal mass moment of inertia / = 2/5mRl) main tains contact with the rim at a variable point B. Although the rim surface often has a small but nonzero radius of curvature in the vertical plane, so that its surface is approximately toroidal, here it is modeled as a circle, the intersection of the planar green surface and the cylindrical hole" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000276_jjap.38.3338-Figure6-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000276_jjap.38.3338-Figure6-1.png", + "caption": "Fig. 6. A hybrid-transducer-type ultrasonic motor with flanges.", + "texts": [ + " \u201cOut of phase drive\u201d is an inappropriate driving waveform, which has the same combination as \u201cin phase drive\u201d but the phase of 3rd mode vibration differs by 180 degrees, as shown in Fig. 5(c). As shown in Figs. 4(a) and 4(b), the value of Wn can be reduced by the \u201cin phase drive\u201d method. Therefore, we can expect low wear of the friction material by using a proposed driving method. The rectangular waveform is highly effective for reducing the Wn value especially when the longitudinal vibration is not sufficiently large. Figure 6 shows the schematic illustration of a hybridtransducer-type ultrasonic motor with flanges.2) The stator has a symmetrical design. The diameter of the stator and the rotor is 30 mm, contact force is 98 N and driving frequency is 21.5 kHz. The maximum rotation speed is about 1.3 rad/s. Friction material, which is a composite of asbestos and rubber, and is called asbestos joint sheet, is adhered on the stator side of the sliding surface. Outer and inner diameters of the tor is made of stainless steel" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002007_0379-6779(88)90597-8-Figure2-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002007_0379-6779(88)90597-8-Figure2-1.png", + "caption": "Fig. 2. Cyclic voltammograms of poly~3-methylselenophene) grafted on Pt (obtained at constant potential using 300 mC cm- ). Sweep rate: 0.1 V s - l in CH3CN + saturated LiC104. (1), 1; (2), 150; (3), 10 000 cycles.", + "texts": [ + " This is performed by recording the charge involved during a double potential step (0 V, 1.15 V, 0 V). The upper limit is chosen at the foot of the cyclic voltammetric curve so that electropolymerization should give a yield of 100%. Assuming that the polymerization involves two electrons per monomer unit, the doping level 5 is given by = 2 Q r d / Q o x - Qrd where Qo. is the charge consumed during the formation of oxidized polymer and Qrd the charge corresponding to its reduction (Fig. 1). With the experimental conditions defined above, we obtain 5 = 0.40 + 0.05. Figure 2 shows the electrochemical behaviour of a poly(3-methylselenophene) film synthesized at 1.15 V until 300 mC cm -2 have been passed. The amount of charge involved during the first cycles is about 20% of the amount used to synthesize the film. Since 5 = 0.40, only 300 \u00d7 (2/2.4), i . e . , 250 mC cm -2, is consumed in the polymer formation and the remaining 50 mC cm -2 is consumed in the oxidation of the film. This value permits the conclusion that the whole film is electroactive. The oxidation and reduction processes of the film are chemically reversible because the amounts of charge involved are equal to within 2%" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000108_la00014a041-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000108_la00014a041-Figure1-1.png", + "caption": "Figure 1. Structure of the spectroelectrochemical cell: (TB) Teflon bushing, (OFW) optical flat window, (ITGC) the internal threaded glass connector, (PG) Pyrex glass.", + "texts": [ + " 0 1994 American Chemical Society SERS from Silver Electrode Surfaces Langmuir, Vol. 10, NO. 2, 1994 587 (i) Reagents and Solution Preparation. The following chemicals were used without further purification: lithium hexafluoroarsenate (LiAsFs), electrochemical grade (Lithco, Lithium Corp. of America); methyl acetate (MA), anhydrous, 99+ % ; and tetramethylammonium bromide ((CHs)rNBr), 98 % (Aldrich Chemical Co.). The salt was slowly added to the solvent in a Braun drybox with an argon atmosphere. (ii) Electrochemistry. The three-electrode spectroelectrochemical cell (Figure 1) for this nonaqueous solution study was designed and constructed in our laboratory. The cell has allglass arms sealed by internal threaded glass connectors (ITGC) (US. Patent 3,499,642). DuPont \u2019Kalrez\u201d O-rings were used to seal the system. Only \u201cKalrez\u201d O-rings were found to be chemically stable toward methyl acetate. The counter electrode was a 0.3-mm platinum wire (99.99% Aldrich Chemical Co.) sealed in glass. The reference electrode was Ag/AgBr/Br. A l-mm silver wire (sealed in glass) was placed in 1 M KBr solution and degassed with Nz for 24 h while being oxidized at a current of 1 mA/cm2 (controlled by a Princeton Applied Research Galvanostat Model 273)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003795_s1554-4516(05)02009-0-Figure12-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003795_s1554-4516(05)02009-0-Figure12-1.png", + "caption": "Fig. 12. Cyclic voltammograms of H2O2 in pH 6.9 phosphate buffer 10 mmol lK1, with 0.1 mol lK1 KCl 25 8C, at a GC electrode modified with MP11\u2013DDAB film: (a) no H2O2 ; (b) 1 mmol lK1 H2O2; scan rate: 50 mV sK1.", + "texts": [ + " CV results of a positive shift of half-peak potential and quasi-reversible characteristic of electrochemistry for MP11 on GC electrodes indicate that conformation change from random coils to a-helix for MP11 hinders the electron transfer between MP11 and the electrode. It may prove that the exposure of the heme of MP11 diminishes and the distance between the electrode and the active site increases. Further more, peak current ip is proportional to the scan rate n (shown in Fig. 11(b)), as predicted by thin-layer electrochemistry theory [58]. The catalytic reduction of hydrogen peroxide was performed on the GC electrode with MP11\u2013DDAB film (shown in Fig. 12) as soon as hydrogen peroxide was added to the buffered solution. Cyclic voltammograms for the electrode with and without H2O2 show that MP11 in DDAB film keeps its catalytic activity. Heme absorption is a very useful conformational probe for the study of heme proteins as well as positions of the Soret absorption band, providing information about the environment of heme [59] on the binding of MP11 to DDAB vesicles. The 400 nm band (shown in Fig. 13, solid line), characteristic of the native random coils MP11 in solution [60], is shifted to 410 nm (shown in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001696_s0263-8223(01)00168-4-Figure7-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001696_s0263-8223(01)00168-4-Figure7-1.png", + "caption": "Fig. 7. Spring subject to bending-torsional surface load in surface-loading condition.", + "texts": [ + " (21), we have U \u00bc L Z p a a D11D66 D2 16 Dj j \" QR sin h F 2 \u00f0R cos a R cos h\u00de L 2 Rdh # : \u00f022\u00de By simplifying the above expression, the strain energy expression becomes U \u00bc F 2R3 L D11D16 D2 66 jDj w2 p 2 a \u00fe sin 2a 2 \u00fe cos2 a p 4 : 2w a 2 \u00fe 1 4 p 2 a sin 2a 2 ; \u00f023\u00de where w \u00bc cos2 a p 2 a \u00fe sin 2a 2 : Thus, the deflection is dxy \u00bc oU oF \u00bc 2FR3 L D11D66 D2 16 jDj w2 p 2 a \u00fe sin 2a 2 \u00fe cos2 a p 4 2w a 2 \u00fe 1 4 p 2 a sin 2a 2 : \u00f024\u00de The spring stiffness is Kxy \u00bc F dxy \u00bc L 2R3 D11D66 D2 16 jDj w2 p 2 a\u00fe sin2a 2 \u00fe cos2 a p 4 2w a 2 \u00fe 1 4 p 2 a sin2a 2 1 : \u00f025\u00de 2.2.3. Bending-torsional stiffness (Kxz) The composite spring is subject to uniformly distributed surface load in the z-direction as shown in Fig. 7. Similar to line-loading condition, the outof-plane bending moment M and the twisting moment T are taken into account in order to formulate the strain energy expression of the complete spring. Hence, U \u00bc 1 Do Z n M2 dn \u00fe 1 b Z n T 2 dn : \u00f026\u00de By substituting the out-of-plane bending moment and the twisting moment into Eq. (26), we have U \u00bc 1 Do Z p a a MB cos h \u00fe FR 2 cos a sin h 2 Rdh \u00fe 1 b Z p a a MB sin h \u00fe FR 2 1\u00f0 cos a sin h\u00de 2 Rdh: \u00f027\u00de Thus, the strain energy expression becomes U \u00bc F 2R3 1 Do u2 p 2 a sin 2a 2 \u00fe 1 4 cos2 a p 2 a \u00fe sin 2a 2 \u00fe 1 b u2 p 2 a \u00fe sin 2a 2 \u00fe 2u cos a \u00fe 1 4 p\u00f0 2a\u00de \u00fe cos2 a 4 p 2 a sin 2a 2 ; \u00f028\u00de where u \u00bc 1 b cos a 1 Do p 2 a sin 2a 2 \u00fe 1 b p 2 a \u00fe sin 2a 2 ; which leads to dxz \u00bc oU oF \u00bc FR3 Do 2u2 p 2 a sin 2a 2 \u00fe 1 2 cos2 a p 2 a \u00fe sin 2a 2 \u00fe FR3 b 2u2 p 2 a \u00fe sin 2a 2 \u00fe 4u cos a \u00fe 1 2 p\u00f0 2a\u00de \u00fe cos2 a 2 p 2 a \u00fe sin 2a 2 \u00f029\u00de and the bending-torsional spring stiffness is Kxz \u00bc F dxz \u00bc 1 R3 1 Do 2u2 p 2 a sin 2a 2 \u00fe 1 2 cos2 a p 2 a \u00fe sin 2a 2 \u00fe 1 b 2u2 p 2 a \u00fe sin 2a 2 \u00fe 4u cos a \u00fe 1 2 \u00f0p 2a\u00de \u00fe cos2 a 2 p 2 a \u00fe sin 2a 2 1 : \u00f030\u00de Eighteen specimens were manufactured by circumferential winding of plain weave E-glass woven cloth impregnated with epoxy resin on mandrels" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001512_bf03184459-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001512_bf03184459-Figure1-1.png", + "caption": "Fig. 1 Schematic diagram of a dynamically loaded journal bearing", + "texts": [], + "surrounding_texts": [ + "The Reynolds equation in short bearing can be written as: (4)p(Z) conditions given above. 6I]CUcos~+ c( \u00a2 - iiJ)J h3 [z2- (B/2)2J F r = - fpcos~. R\u00b7 de : dz (5) F 1J = fpsin~ . R \u2022 d~ . dz (6) The equations of force equilibrium are given by: Wcos\u00a2=Fr (7) Wsin\u00a2=F1J (8) By applying the mobility method to the Eq. (7) and Eq. (8), the following equations for the motion of journal center can be obtained (Boo ker, 1965, Taylor, 1993). S W(C/R)2 \u2022 M (9) I]BD \u2022 c(\u00a2-iiJ) W~Cf:!nR)2. M1J (10) In Eq. (9) and Eq. (10), M. and M 1J are mobility terms, which can be written as: (1- ceos\u00a2) 3/2 M. Jr2(B/D)2\u00b7 [Jrcos\u00a2. (I-ceos\u00a2) -4csin 2\u00a2J (11) (1- cCos\u00a2) 3/2 \u2022 M 1J ~(B/D)2\u00b7 sm\u00a2[4ccos\u00a2+Jr(1 -ccos\u00a2) J (12) From the assumed initial values of e and \u00a2, the The hydrodynamic force components due to the pressure distribution acting along and normal to the line of center (see Fig. I) can be determined from: (3) at z= \u00b1B/2, p=O at z=O 1k.=0, dZ tz (h3 ~~)= 121] \u2022 (~r[ scos~+c \u2022 (\u00a2-iiJ) \u2022 sin \u00bb] (I) mental measurement (Bates and Benwell, 1988, Choi et aI., 1992, 1993) for MOFT in engine bearing have been performed. Engine oil has to be sufficiently supplied to the bearing for preventing the collapse of oil film. However it is not a fundamental means of solving the problems for the economical lubrication. Therefore most of the engine bearings adapt to the oil reservation groove as an alternative solu tion. The circumferential groove has some advan tages for lubrication and cooling. However there are only a few research papers on the effect of circumferential groove on MOFT. Jones (1982) studied the effect of groove in engine bearing by considering oil film history, and Choi et a1. (1992, 1993) measured the minimum oil film thickness in engine main bearing with circumfer ential half groove by using the total capacitance method, and then compared with theoretical results. The aim of this paper is to investigate the effect of circumferential groove on MOFT. The oil film pressure is solved by using the short bearing theory, and the mobility method is used to obtain the trajectory of the journal center. From these, the effect of groove on the oil film thickness in engine bearing is presented in the form of MOFT curve. h(~) =C(I +ceos~) (2) The boundary conditions to solve the Eq. (1) are The fluid film pressure can be obtained by twice integrating Eq. (I) using the boundary Table 1 Operating condition of test bearing Con-Rod Main BID 0.39 0.33 Clearance 24,um II, 20, 23, 25 ,urn rpm 3500, 5500 Load Condition Ful1load Viscosity 6.26cP 9.4IcP (a) 3500 rpm (b) 5500 rpm (b) Grooved bearing(a) Ungrooved bearing Fig. 2 Schematic diagram of axial pressure distribution new values of eccentricity and attitude angle is obtained by the time integration of Eq. (9) and Eq. (10), and the time integration is repeated until the convergence for journal center trajectory is achieved. When the journal center is located at the groove area, the oil film pressure is obtained at each of the two bearing sides except the groove. Figure 2 shows typical axial film pressure distri bution in ungrooved and grooved bearings. In order to balance any given external load, the grooved bearing has higher maximum pressure than the ungrooved bearing." + ] + }, + { + "image_filename": "designv11_65_0001507_iros.1993.583189-Figure2-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001507_iros.1993.583189-Figure2-1.png", + "caption": "Figure 2: definition of Zg", + "texts": [], + "surrounding_texts": [ + "Let At be the sampling period. If At is small enough, bXd(t) can be approximated by: bXd(t) = bXd(t - At) + Atbkd(t - At). Furthermore, bXd(r - At) = bAg.gXd(t - At) where bAg is the rotation matrix from Cb to Cg, and %d(t - At) represents the desired velocity at time (t - At), expressed in the gripper frame. This is precisely the value given by the artificial constraints. Therefore, gXd(t - At) is equal to (-vY) for subtask 1 and (-VIZ) for subtask 2. Let us denote At.Xd(t - At) as AXd(t). Therefore, we have: bXd(t) = bXd(t - At) + bAXd(t) The selection matrix S is then given by: If we take for the initial value bXd(o) = bX(0), we can compute the desired trajectory bXd(t). It is input in the position servoloop as presented in figure 4. Let us recall that the directions are selected in the constraint frame Cg. After this selection, the input g<(t) is possibly expressed again in the base frame before entering the \"position control law\" block of figure 1. This depends on the way of this position control law is programmed. 2.3. Whv is this calculation incorrect 2 The desired trajectory computed above is in contradiction with the task description proposed by Mason. To prove it, let us consider for example the f i s t subtask of our application. If we consider the gripper position at time t, the desired motion which is effectively taken into consideration after selection by S is defined along the direction Y of the gripper frame. But the handle has to rotate about its axis. Therefore, the desired position will never be reached. We will show that, because of this, the increment which is used in the control scheme (after S ) is in general different from vAt: consequently, the actual velocity reference value is not v. Let us consider the first displacements of the grasping point R (figure 5). This point is located at the tip of the handle for the sake of clarity in the figure. Also for clarity reasons, the increments are exagerated in figure 5, although this is not realistic. Di is the desired position at time ti (X,(t,)) Ri is the real position at the same time @(ti)) Yi is the Y axis of the gripper frame at the same time Si is the orthogonal projection of Di on Yi-l ti+l = ti + At At the initial time to, Do = R,. The arc represents the trajectory of the grasping point. Let us consider the error vectors E, and <. Generally speaking, the error at time $+1 is: E,(ti+l) = RiDi+l After selection by S, only the component of &,(ti+l) along Y is taken into account: <(ti+l) = RiSi+l Therefore the desired position which effectively acts in the control is Si+l. We always have: DiDi+l= -vAt Yi Si+l is the orthogonal projection of Di+l on Yi. If Ri was the orthogonal projection of Di on the same axis Yi, we would then obtain: = DiDi+l = -vAt Yi But this assumption is wrong, except by chance or for i=l. As a matter of fact, if we suppose that R, is the optimal position, that is to say the position which minimizes the error RiSi, then Ri is the orthogonal projection of Si on Yi (this is the case in figure 5). Since the grasping point trajectory is on the arc, the direction of Y changes as a function of time. In particular, the axes Yi-l et Yi define different directions. Therefore Si, which is the orthogonal projection of Di on Yi-l, is not aligned with Ri and Di. Consequently, R, is not the orthogonal projection of Di on Yi and therlefore: <(ti+,) f -vAt Yi. This is what can be seen in figure 5. This result is of course valid when Ri is not the optimal position. Generally speaking, the value which is applied to the robot is not vAt. 2.4. Correct calculation of b l ~ d ~ Since the mistake comes from the fact that Ri is not the orthogonal projection of Di on Yi, it is sufficient to replace Di by Ri for calculating the new desired po,sition Di+l : Then: E*(t. ) = = -vAt \\Ii. x 1+1 More formally, the relationship RiDi+l = -vAt Yi yields: bXd(t) = bX(t - At) + bLixd( t). The new desired position is calculated froim the last position actually reached, not from the previous desired position. This method can be generalized to other tasks, because it really uses Mason's description, which is not the case in the computation (widely used) of section 2.3. As a matter of fact, when the gripper is in bX(t - At) at time (t - At), to require it to be in bX(t - At) + bAxd(t) at time t is to require a displacement bAXd(t) which is precisely the velocity input in the base frame (multiplied by At, which is a constant). Figure 6 shows the new method of computing the position error after selection, {(t) : Since the same quantity bX(t - At) is successively added and subtracted, this term has no influence in the computation. Then the f i s t two frame transformations (bAg and gAb) also cancel each other. Finally: gE,(t) = gAXd(t) and there is no more position servoloop. 2.5. CO nseaue nces Since we eliminate the position loop, we also eliminate the problem of expressing a desired trajectory in a base frame. Both position and force reference values can be expressed in the constraint frame. Figure 7 shows the new computation of &:(t) and E;(t). As already mentioned, the frame transformation bAg is optional: it depends on the control laws which are then used. As can be seen in figure 7, S acts on a velocity (times At) not on a position error. We thus obtain a selection of velocities and forces, as described in Mason's approach. Our task description led us to simplify the control scheme, which will be of interest to save computation time in a real implementation. However, the lack of position loop makes the robot sensitive to disturbances. In fact, during the task, the disturbances which act on the directions selected by S cannot be cancelled; in addition, at the end of a subtask, the robot must stay in its final configuration, which requires getting information on its position. It is therefore necessary to include another position servoloop in the control diagram, which is compatible with the results we have obtained and with the hybrid control concept. This is presented in the next section. 3. The new hvbrid control diamam We are looking for a simple solution for the position servoloop in order not to produce too many additional computations. Since there is no loop in Cartesian space, this new position loop must be located after the sum of the position and force contributions, which is compatible with the hybrid principle. The simplest solution would be a PID controller acting on the joint variables. Therefore, the command vector C must be transformed into a jointspace command q* before entering the PID controller. For the position part, we already have a desired velocity x, constant on a sampling period At. So, it is sufficient to define C,(t) and Ckt) as velocities. C,(t) is obtained from &:(t) via a proportional gain We also choose a proportional gain for the force part, which transforms the force error Er(t) into a displacement: AXkt) = K&t). The inverse of K, represents the desired stiffness matrix of the robot. Therefore the choice for K, depends on the type of environment: a bad choice for K, would produce an unstable behavior of the robot [6]. Then Cdt) = 7 is defined as the desired velocity in the force selected directions. Finally, the total desired velocity in task space is defined from the desired displacement in the S selected directions and from the force error in the (I - S) selected directions: E: (t) 1 AXfO) X*(t) = C(t) = CJt) + Cfct) We then use the pseudo-inverse J + of the robot Jacobian matrix J to obtain a jointspace velocity vector, that we integrate to obtain q*. Figure 8 presents the block diagram that we have determined step by step. bAg is optional, but J is generally computed with respect to the base frame and it is therefore necessary to express < and E; in this frame. This is what we have done in our real experiments. and we set gAXd(t) = 0. Then, the force loop no longer plays any role and X* = 0, q* = 0. The reference q* is maintained to its last value and the robot does not move anymore. It is more reasonable to entirely position control the robot for the following reason: when the subtask is ended, the constraints may have changed. This is the case at the end of the first subtask in our door-opening application: the motion along Z is then possible (rotation of the door about the hinge). So, if we apply a disturbance force in this force-controlled direction, the robot will move in order to assure F = Fd (= 0 in our case), and we cannot control this motion. The PID controller has other advantages. Since it is a jointspace loop, there is no need for frame transformation, and the computation of U from q and q* is much faster than the computation of q* from the Cartesian data (see figure 8). In our case (see section 4), the respective computation times are 3 and 30 ms. The jointspace loop runs much faster than the Cartesian space loop. Therefore, the desired value q* is sent via several steps (10 in our case), which is good for the smoothness of the motion. In addition, the PID controller acts on the force loop global behavior; its smaller period improves the time response of the force loop and the disturbances acting in the force-selected directions are cancelled more quickly. Another advantage of the joint PID is that the scheme presented in figure 8 can be easily implemented on industrial controllers which already use joints PID boards in order to stabilize the process. m e r i m e n t a l resu I t s Our experimental setup is composed of a Puma 560 robot, an AICO controller, and a door with its handle. Force measurements are given by an AICO wrist force sensor. The original LSI 11 controller (made by STAUBLI) has been replaced by the AICO controller, because its communication and computational capabilities were not sufficient for our application. The AICO controller is based on a VME bus, to which we have added several boards, including five 68020 CPU boards. This helps to compute force and position loops at the same time [7]." + ] + }, + { + "image_filename": "designv11_65_0001041_1.2831575-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001041_1.2831575-Figure1-1.png", + "caption": "Fig. 1 Observation of fringe pattern formed by object plane and obser vation plane", + "texts": [ + " Two coherent light beams intersecting each other combine to form many equally spaced interference planes parallel to the bisector of the intersect angle. Cutting the interfer ence planes with the observation plane generates a fringe pattern comprising many intersection lines resulting from the interfer ence and observation planes. Hereafter, the surface equation of the object plane is derived using information constituting the fringe pattern. Let us consider two planes, an object plane and a observation plane as depicted in Fig. 1. Let x and y coordinates be located on the observation plane with O' being the origin where a reference beam, e\u0302 (directional vector), intersects; the z axis erected perpendicular to the observation plane intersects the object plane at O, o (positional vector); and a light source be located at 0 , q (positional vector). An incident beam, e, 564 / VOL 118, JULY 1996 Transactions of the ASME Copyright \u00a9 1996 by ASME Downloaded From: http://tribology.asmedigitalcollection.asme.org/ on 02/24/2018 Terms of Use: http://www" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002254_isic.1995.525099-Figure2-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002254_isic.1995.525099-Figure2-1.png", + "caption": "Fig. 2 . Planar wheel driven mobile vehicle", + "texts": [], + "surrounding_texts": [ + "Potential Fields fior Nonholonomic Vehicles K. J.Kyriakopoulos, P. Kakarnbouras and N. J.Krikelis\nControl Systems and Intelligent Automation Laboratory\nDepartment of Mechanical Engineering\nNational Technical University of Athens, At:hens 10682\nGreece\ne-mail: kkyria@central.ntua.gr\nA b a h c t - The problem of motion planing of a wheeled nonholonic vehicle is treated by decomposing the problem to the subproblems : (i) flnd a collision free path and (ii) approximate this path with a nonholonomic coliision free path. Thin is treated, in real time, by solving the flrst subprobllem using a potential flelds strategy and the second with nonhollonomic tracking. Thus, collision avoidance of a nonhollonomic wheeled vehicle in a feedback formulation is achieved.\nmerit of our method is tha t it is robust to environment changes and execution errors because it is a feedback\nscheme.\nIn section 2 we present a mathematical problem statement along with the required analysis. Simulation results to demonstrate the value of our method are presented in section 3. Finally, in section 4 we present some issues for further resarch.\nI. INTRODUCTION 11. PROBLEM STATEMENT A N D ANALYSIS\nThe increasing interest on autonomous vehil:les has motivated during the last few years the interest of robotic researchers on nonholonomic motion planning. A number of approaches have appeared [CFLSl], [dWS91], [MS90], [RL92], [SW92] in the literature but primarily without considering the presence of obstacles. Recently a novel approach for nonholonornic motion planning in the presence of obstacles has appeared\n[JPMR94]. The basic idea presented in this paper is to decompose the problem to the subproblems : (i) find a (possibly holonomic) collision free path and (ii) find a nonholonomic collision free path around that .\nIn this paper we present a strategy for nmholonomic path planning in the presence of obstacles based on a similar idea but with different tools. The find path problem is solved using artificial potential fields and the nonholonomic approximation of that path is realized with our nonholonomic tracking scheme that has earlier appeared in the literature [PK93a,]. The\nConsider a wheeled mobile robot in a geometric environment similar to this depicted in figure (1). The purpose is to find a path connecting the initial configuration with t:he final one by avoiding the obstacles and obeying the nonholonomic constraints.\nA . Mathematical Problem Statement\nIt has been shown in [PK95] tha t the equations describing the motmion of the fratme attached to point R of the rear wheel axis (2) are given by\nX = v . c o s 0\n= v ' s i n 0\nB = LJ\nwhere (2, y) defines it,s position tion. Idet, 11s represeiit by W c by ai, i = 1 , 2 , . . . the obstacles a t a configuration q = [z y e] . T and 0 its orienta!R2 the workspace, and d(q) the robot\n0-7803-2722-5195 $4.00 0 1995 IEEE 46 1", + "We seek a path z(s), y(s) s E [0, SI], parameter- a) connecting the starting and destination points,\nb) satisfying the nonholonomic constraints, written ized by the its length s :", + "now as\nd x - = cos0 ds (4) (5) dY - = sin 8 ds\nsince v = $, and c) satisfying the collision avoidance constraints with the environment, i.e\nB. Potential Fields\nLet us consider a point E on the y-axis of the car that is very close t o R. Assume the existence of a\nglobally converging potential function V ( x , y) that can be constructed to lead the system to the destination (origin) using any one from the numerous tecliniques presented in the literature [RK89], [PR88]. Assuming a holonomic motion model for E\nand considering a Lyapunov function candidate\nglobal asymptotic convergence of E to the origin is achieved.\nI t should be noticed tha t the resulting strategy (13)\nis a feedback one and the resulting path holonomic.\nC. Nonholonomic Tracking\nThe motion point E is given by\nwhere c is the srnall distance of E from R, on the y-\naxis.\nA feedback control strategy for tracking of nonholo-\nnomic vehicles was proposed by Pappas and Kyriakopoulos [PK93b]. I t was shown to be asymptotically converging and can be used for E t o track a desired trajectory. I t has the form\nv\nw\n= -k(xe cos 0 + ye sin 0) + i d COS 0 + y d sin 0\n= 7 [-k(ye cos 0 - x, sin 0) - i d sin 6 + y d cos 81 1\n(16)\nwhere ze( t ) = zE( t ) - \" E d ( t ) , ye(t) = Y E ( t ) - YEd(t) are the position errors of point E with respect t o a reference trajectory z~~ ( t ) , y ~ ~ ( t ) . This scheme has\nbeen reported to be very robust during experimenta-\ntion [AK95a], [AK95b].\nD. Control Architecture\nIf we combine the two control laws described in the previous subsections we derive the control structure\nthat is depicted on figure 3. Stability of the overall scheme can be derived by appropriately satisfying the conditions of a theorem by Khalil, Kokotovic and\nO'Reilly [PK]. Kokotovic.\n111. S I M U L A T I O N STUDIES\nIn order to demonstrate the value or method we present a simulation example. The considered environ-\nment and the equipotential curves are demonstrated in fig. 4. Notice, that the construction of the potential function was not done by using the methodology by Koditschek [RK89] but simply by adjusting the gains to produce a patlh with no local minima.\nThe trajectory resulting from such a potential function without considering the nonholonomic constraints is demonstrated on figure 5 (a) , while the trajectory\nfrom our method is shown on figure 5 (b). Obviously\nthis trajectory is \"smoother\" due to the nonholonomic nature of tracking." + ] + }, + { + "image_filename": "designv11_65_0002421_890146-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002421_890146-Figure1-1.png", + "caption": "Figure 1. Schematic showing essential features of EMA-LS9 simulator.", + "texts": [ + " Provision is made for lubricant spray at the wear interface and the entire unit may be heated to as high as 550\u00b0C in order to simulate advanced diesel top ring reversal temperatures. Data from the EMA-LS9 simulator include friction force as a function of crank angle and average friction coefficient as a function of time. As is often done after such tests, liner samples may be analyzed for wear volume, surface finish, and other meaningful wear parameters. Simulator results are available for both right and left wear interfaces. Figure 1 shows a schematic of the simulator. By pressurizing the air cylinder, ring loads are applied which can simulate the high pressures experienced in highly turbocharged engines under peak pressure conditions. Maximum reciprocating speed is limited to 700 rpm by inertia forces. In effect the simulator attempts to duplicate the most severe ring and liner condition; namely low speed, high load at TDC firing. Most of our testing to date has been at speeds below 500 rpm. Higher speed operation, if desired, would require addition of a reciprocating balance mechanism", + " As discussed previously, unlubricated EMA-LS9 wear rig evaluations were performed on the same wear couple combinations using samples of rings and liners used for the 300 hr engine test. Profilometer traces from these tests (figures 7 & 11) show the performance of the chromium oxide and TiC/Cr3C2 couples to agree with the engine wear results on a relative basis. Interface Oil Quantity and Quality Considerations One of the significant features of the EMA-LS9 is the means by which the friction force is obtained. The EMA-LS9 employs strain gauged pivots (see figure 1 for location) which permit measurement of right and left friction forces simultaneously. Viewing these results, it was apparent that side to side differences initially existed. Unlike an engine, the oil supply at the wear interface of the simulator can be systematically varied in both quantity, and quality. Friction differences side to side are virtually eliminated when oil quantity and quality are also equal side to side. Figures 16 and 17 show how average friction coefficient history varied when oil supply quality was more equalized side to side" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003764_rob.20045-Figure5-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003764_rob.20045-Figure5-1.png", + "caption": "Figure 5. Spatial 3-link manipulator.", + "texts": [ + " When ignoring the obstacle, the resultant motion is very simple [Figure 3(a)]. However, if we lowered c , the motion resembles the local minimum (A). The minimum times are 0.69, 0.99 and 1.16 s in Figures 3(a), 3(b) and 3(c), respectively. Figure 4 shows the actuator torques (solid lines) and the equivalent torques (dotted lines) during a minimum-time motion, where joint 2 is almost saturated during the entire motion, and joint 1 is saturated in the latter half. The model is a 3-link arm shown in Figure 5, where all joints are revolute pairs around their z-axes and it is the configurations of zero-displacements. Base coordinates are the same as the first link coordinates. The specifications are listed in Table 2, where gravity is acting in the z0 direction and c is about twice the static actuator torques necessary to endure gravity in fully stretched configuration. The dimensions of two hexahedral obstacles are both (0.4, 0.4, 0.5) m and the centers are (0.76 0.47 0.25) and (0.76 0.47 0.25) m in base coordinates" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000412_s0263574700017768-Figure2-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000412_s0263574700017768-Figure2-1.png", + "caption": "Fig. 2. Geometric illustration for the manipulator arm.", + "texts": [ + " And the desired position and orientation of the hand are described by the homogeneous matrix n, o, ay Yd *z %d 0 0 0 1 (1) The two operators x and \u2022 denote a vector and a scalar product, respectively, and J, = sin 0, and ci = cos 0, are used as the conventional notation throughout the derivation. 2.1 Closed form arm solutions based on geometric identity method In this section, we restrict our attention to a geometric and an intuitive method for solving the redundant kinematic equations, because analytical approaches may serve some purposes, viz. real-time motion control or better understanding of arm mechanisms throughout the process of finding the solutions. Figure 2 is useful in the present consideration of mainly treating the spatial geometry of the linkage as a plane geometry problem. First, in view of the special arm structures in which the last three joint axes intersect at the wrist, we see that their joint angles have an influence only on the orientation of the terminal link, and thereby the position of the wrist is readily expressed as: - xh - =Xd- = Zd- (2) Likewise, using the direction cosines D4, \u00a34 and F4 of a fore-arm EW, we can describe the position of the elbow as follows: xe = xw ~ (3) In fact, these positions are given straightforward via the Kinematics problems 235 T= n\\(\\ - c o s n,n2(l - c o s (l - cos co-ordinate transformation matrices described in where a transformation operator T is given by: Appendix 1", + " Among four joint angles, the geometrical identity 2lxl2c4-R}-l\\-l\\ associated with the triangle SEW, derives an explicit solution (4) + cos 0 + n3sin - n2 sin cf> l - cos (f>) - n3 sin n\\(l - cos (f>) + cos 0 ~ cos 4>) + rii sin 4> M i n 3 ( l \u2014 c o s ) + n2 s in n2n3(l - cos 4>) - rii sin 0 n\\{\\ \u2014 cos ) + cos (f> (9) where Empirically, we know that the human elbow possesses one single degree of freedom of rotating about a straight line connecting the shoulder S with the wrist W, with the location of the wrist fixed and that this redundant rotating mechanism serves for the flexibility or dexterity of arm configurations. For this reason, a special variable, called the elbow revolute angle, is adopted here as one of the practical methods to resolve the present redundancy.14 Now, let the current position of the elbow be on a plane TT, (oblique shaded portion) of containing points 5, W and z -axis. From a geometric observation illustrated in Figure 2, we see its location E0(x\u00b0e, y\u00b0e, z\u00b0) is given by: 0 cos 0 /, cos {LE0SQ)yw (5) where z\u00b0 = /, sin (S-E0SQ), LE0SQ = a + j8 (for d4 > 0), LEQSQ = a - /3 (for 04 < 0). Note that another possible configuration of the elbow \u00a3 0 indicated by the dashed lines will correspond to the negative sign of 04 directed clockwise. And expressions for these angles a and j3 are found from tan a = cos /3 = \u2022 21, R (6) Here, thinking of an elbow vector {-, as describing rotational motion of the elbow, we have = \u00a3 cos from the initial point \u00a3\u00b0 = (x\u00b0, y\u00b0, zj", + " And the expansion of individual matrices yields sss6 = P4ax + Q4ay + R4az = G,, c5*6 = ~{U4ax + V4ay + W4az) = G2, (14) c6 = D4ax + E4ay + F4az = G3, from which we can calculate and tan 06 = G2 G3S5 Finally, we obhun/~~) tan 07 = r * F4n. 4n. \u00a340, + F4o, (15) (16) (17) As stated above, the analytical solutions for the redundant manipulator were derived with relative ease by means of an observation on the elbow motion of the human arm. In passing, this revolute angle can be evaluated in the following way when each joint angle of the manipulator is given. That is to say, Figure 2 shows that the oblique shaded plane TT,, which may be considered as a reference plane for motion of the elbow 236 Kinematics problems in the Cartesian space, moves to a new plane K2 of containing the triangle SWE along a circumference with the change of the elbow angle. In this context, each plane equation is expressed by z ze x y z 1 x y xw yw 0 1 xt ye X V Z. \\ X V Z. o o o i o o o = 0, or axx + b{y = 0 (for 7T,) and a2x + b2y + c2z = 0 (for n2). (18) Since the revolute angle is equivalent to an intersection angle of both planes, which also may be regarded as the angle between vectors normal to planes KX and n2, the resulting intersection angle is calculated as: COS under some desired position and orientation of the hand" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000759_004051759106101210-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000759_004051759106101210-Figure1-1.png", + "caption": "FIGURE 1. Warp slide on an elliptical cross section of the filling yam.", + "texts": [ + " We also assume that the filling crimp in the horizontal direction is negligible, since the reed holds it at the instant of beat-up and prevents horizontal crimping. Our third assumption is that the filling yam cross section in the cloth is an ellipse with a major axis ( 2a ) and a minor axis ( ~b ) . This flattening of the filling yarns is the result of the pressure generated by the warp tension. The aspect ratio (A = a / b ) is a function of yarn structure [6]. Yam Tension for Elliptical Cross Section When the warp moves relative to the pick with a velocity V, as shown in Figure 1, the warp tension Tro is less than tension Tt. The relationship between tight and slack warp tensions is a function of parameters At, n, r, k, ro, and ~, where u is the frictional coefficient and n is the frictional index between the warp and the pick, respectively, r is the radius of the filling yam, k - 62 , b is the minor radius of the ellipse, a is the a major radius of the ellipse, io is the angle between the radius of the contact point s and the right major radius, and ~ is the angle between the radius of contact point t and the tight major radius" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003086_acc.2001.946148-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003086_acc.2001.946148-Figure1-1.png", + "caption": "Figure 1: Interception geometry.", + "texts": [ + " Guidance Model and Associated Uncertainties The derivation of the suggested robust guidance law requires the presentation of the intercept kinematics and associated assumptions. To this end, the framework of the well-known PNG is adopted. The general formulation of a three-dimensional PNG is rather complicated; however, by assuming that the lateral and longitudinal maneuver planes are decoupled by means of roll-control, one can deal with the equivalent two-dimensional problem in quite a realistic manner. Thus, it is assumed hereafter that the geometry is two-dimensional. This assumption renders the planar interception missile-target geometry depicted in Fig. 1. This figure describes a missile employing PNG to intercept a maneuvering target. Based on Fig. 1, a linearized model of the guidance dynamics can be developed. Such a model is widely used in the analysis of PNG''-l4. A block diagram describing the linear model is given in Fig. 2. In this linear time-varying system, missile acceleration aM is subtracted form target acceleration aT to form a relative acceleration jj . A double integration yields the relative vertical position y (see Fig. l), which at the end of the engagement, I = t , , is the miss distance y ( t , ) . By assuming that the closing velocity V, is constant, the relative range is given by: R(t) = vc " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002726_iros.1994.407422-Figure4-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002726_iros.1994.407422-Figure4-1.png", + "caption": "Figure 4: The physical model of the terrain.", + "texts": [ + " According to this concept, the terrain is represented by a set of interconnected particles @(Pi) having the following properties [lo]: (1) each particle is seen as a point mass m which obeys Newtonian dynamics and which is surrounded by a spherical non-penetration \u201celastic\u201d area; (2) the set of particles corresponds to the mass, the inertia, and the spatial occupancy characteristics of the modeled object; (3) the particles are interconnected using interaction components (refered to as the \u201cconnectors\u201d). Each connector corresponds to a type of interaction, and is modeled using appropriate physical laws -interaction laws like viscous/elastic behaviours, elastic collisions, or dry friction are modeled by combining \u201cspring\u201d and \u201cdamper\u201d components appropriately- (as shown in figure 4). The discretization of the terrain in terms of such elementary physical components requires that several criteria such as the terrain surface shape, the average distribution of the contact points between the wheels and the ground, and the complexity of Q(7) (i.e. the number of particles should tion of the system, the mass distribution is computed by converting the geometric model G(7) into a set S I of spheres Sj whose profile approximates the surface of 7 - a particle @(Pi) is then automatically associated with the center of Si- (see figure 4). This is done in such a way that each point of the terrain surface (i.e. list of the points given by G(7)) should be located on the surface of at least a sphere of S I . The efficiency of such a representation is related to the be taken into consideration. In the current imp 1\u2018 ementa- fact that it allows us to maintain the geometric features of the motion planning problem (i.e. checking the geometric constraints as the no-collision), and easily build the physical model @(7), and to formulate the interactions with the robot. Indeed by describing 7 in terms of a set of spheres, one needs to handle a smaller amount of information to represent its geometric shape (as shown in figure 4). Furthermore, the combination of such simple primitives with an appropriate hierarchical description of the wheels allows us to compute easily the distribution of the contact points using a fast distance computation algorithm involving the structured sets of the considered spheres [9] -these s heres represent the pairs (particle, non-penetration areay which model the wheels of A and the terrain 7. The most difficult problem to solve when dealing with a vehicle moving on a terrain, is t o define realistic vehicle/terrain interaction laws" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000666_bf00322165-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000666_bf00322165-Figure1-1.png", + "caption": "Fig. 1. Principle of IgG detection using anti IgG-AP conjugate. I = addition of anti rabbit IgG-Ab-AP; H = washing off free conjugate, III = addition of substrate (here p-nitrophenylphosphate)", + "texts": [ + " This antibody was bought from Sigma and two charges have been used: 27F-8842; 0.75 mg conjugate per ml; enzyme activity 445 U/ml at 310 K. 18F-8895; 0.6mg conjugate per ml; enzyme activity 345 U/ml at 310 K. According to Sigma, both have been cleaned immunospecifically, which means that all goat serum proteins have been removed that do not bind specifically to rabbit IgG. According to Sigma, the alkaline phosphatase (AP) was prepared from bovine intestinal mucosa. The detection of the rabbit IgG using anti IgG-AP is shown in Fig. 1. The detailed procedure is as follows: 1. For incubation, dip the antibody coated wires (3 wires per ml) in I ml of a 1 : 1000 diluted solution of the basic anti IgG-AP (total 750 ng) in the considered buffer, at room temperature. 2. For washing, dip the wires into 15 ml of the considered buffer for 5 min, repeat three times. 3. Wash with distilled water. 4. Put the wires into the substrate solution (0.5 mg pnitrophenylphosphate = pNPP per ml of 1 mol/1 diethanolamine buffer pH 9.8), 1.5 ml for each wire; prepare simultaneously a reference solution of 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003864_50002-2-Figure2.4-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003864_50002-2-Figure2.4-1.png", + "caption": "Fig. 2.4. Effective rolling radius and longitudinal slip velocity.", + "texts": [ + " A small horizontal shift of the curves is sufficiently accurate to change from one definition to the other. The drawback of the last definition is that when testing on very low friction (icy) surfaces, the rolling resistance may be too large to let the wheel rotate without the application of a driving torque. Consequently, the state of free rolling cannot be realised under these conditions. Nevertheless, we will adopt the last definition where r s = r e and consequently, point S is located at a distance r e from the wheel centre. Figure 2.4 depicts this configuration. According to this definition we will have the situation that when a wheel rolls freely (that is\" at M a = 0) at constant spee~ over a flat even road surface, the longitudinal slip x is equal to zero. This notwithstanding the fact that at free rolling some fore and aft deformations will occur because of the presence of hysteresis in the tyre that generates a rolling resistance moment M y . Through this a rolling resistance force Fr = M y / r arises which necessarily is accompanied by tangential deformations" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002851_j.ijmecsci.2003.10.009-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002851_j.ijmecsci.2003.10.009-Figure1-1.png", + "caption": "Fig. 1. Compression between rotating plates\u2014notation.", + "texts": [ + " The result obtained in the present study shows what kind of interpolation functions is appropriate in the elements near the friction surface. A nonsteady planar 0ow of a material obeying the double-shearing model (see, for example, Ref. [2]) between two inclined plates (total instantaneous angle 2 ) rotating about the same motionless line with an angular velocity ! is considered relative to a 1xed system of plane polar coordinate r with its origin O on the line of intersection of the plate walls as shown in Fig. 1. The plate walls are rough and it is supposed that the frictional stress follows the law adopted in Refs. [3,4]. It is assumed that the radial velocity is directed towards the apex O. Because of symmetry, it is suCcient to study the 0ow in the region 06 6 . In the polar coordinates, the system of stress equations which governs the plane-strain deformation consists of the equilibrium equations @ rr @r + 1 r @ r @ + rr \u2212 r = 0; @ r @r + 1 r @ @ + 2 r r = 0 (1) and the yield condition of Coulomb\u2013Mohr ( rr + ) sin\u2019+ \u221a ( rr \u2212 )2 + 4 2 r = 2k cos\u2019; (2) where rr , and r are the components of the stress tensor and the constants k and \u2019 denote the cohesion and angle of internal friction of the material", + " [3]: @u @r + u r + 1 r @u @ = 0: (3) sin 2 ( @u @r \u2212 u r \u2212 1 r @v @ ) \u2212 cos 2 ( 1 r @u @ + @v @r \u2212 v r ) +sin\u2019 ( @v @r + v r \u2212 1 r @u @ \u2212 2 d dt ) = 0; (4) where u and v are the components of velocity in the r and directions, respectively, and is the angle between the poison vector r and the algebraically greatest principal stress. Since d =dt =\u2212!, the convected derivative can be written in the form d dt =\u2212! @ @ + u @ @r + v r @ @ : (5) The boundary conditions are as follows. Due to symmetry = 0 (6) and v= 0 (7) at =0. Since the material is assumed to 0ow towards the apex, the friction stresses f are directed as shown in Fig. 1 and, therefore, the frictional law adopted in Refs. [3,4] is given by = w = 4 + \u2019 2 ; (8) at = . The velocity condition at = is v=\u2212!r: (9) The main hypothesis is that is independent of r. Assume a stress distribution in the form rr k = cot\u2019\u2212 rm ( ; ) sin\u2019 (1\u2212 cos 2 sin\u2019); k = cot\u2019\u2212 rm ( ; ) sin\u2019 (1 + cos 2 sin\u2019); r k = rm ( ; ) sin(2 ); (10) where ( ; ) is a positive valued function of and and m is a function of . It may be veri1ed that the distribution Eq. (10) satis1es the yield condition (2)", + " (22) in the region )\u00bf \u00bf 0, but involve q2 whose value is unknown (q3 should be excluded by means of Eq. (41)). Using an iterative procedure, the value of q2 can be found by applying the boundary condition (28). The variation of q2 with for di2erent \u2019 is shown in Fig. 3. The radial velocity at the friction surface is determined from Eqs. (17), (26) and (40) as u= !r ( G0 r2m cos2 \u2019 \u2212 mq2 cot\u2019 4 ) (43) at = w. It is seen from Figs. 2 and 3 that the second term in the parenthesis is negative. On the other hand, our choice of the direction of the friction stress (Fig. 1 and Eq. (8)) requires that the velocity component u is negative at the friction surface. Hence, in Eq. (43) the 1rst term in the parenthesis must be negative and its magnitude must be greater than the magnitude of the second term. Even if these conditions are satis1ed, it is clear that the solution is not valid if r is large enough, say r \u00bf+, because u at = w is certainly positive as r \u2192 \u221e, as follows from Eq. (43). However, the value of + may be made as large as necessary by an appropriate choice of G0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002644_2003-01-3743-Figure7-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002644_2003-01-3743-Figure7-1.png", + "caption": "Figure 7: left: Device used to measure the cumulative frictional torque: (a) force transducer (b) oil feed (c) motor (d) spindle (e) test bearing (f) adapters (g) hydrostatic friction balance (h) pressure oil (i) stroke guide (j) load cell (k) load and lifting cylinder", + "texts": [ + " At this stage, the bearings are oiled with anticorrosive agents only in order to maintain a high friction coefficient. This allows precise adjustment of the preload force. The frictional torque displayed during assembly is compared with a reference measurement that demonstrates the correlation between preload force and the bearing frictional torque. The reference measurement is indicated on the bearing drawing for the customer's information. To obtain these reference data, statistically reliable measurements were made for frictional torque in single bearings and in pinion gear bearing supports. Figure 7 shows the test stands used to measure frictional torque. To measure the cumulative frictional torques, a pinion head bearing was braced against a pinion flange bearing, driven by a precision rotary table, and an axial load was applied (Figure 7, left). Figure 6: Measured natural frequencies standardized to match customer specifications. The individual frictional torques were measured on a hydrostatic friction balance (Figure 7, right). The outer ring of the bearing is driven, the axial load is applied from below via a hydraulic piston, and the frictional torque is measured on the hydrostatic balance. The individual frictional torques can be added up to obtain cumulative frictional torques and then correspond to the cumulative frictional torques that were measured directly (Figure 8). measurement by INA 0.90 0.95 1.00 1.05 1.10 1.15 1.20 (m ea su re d fr eq ue nc y) /( de m an de d fr eq u en cy ) axle drive size A, pinion shaft, 6 kN axle drive size A, driving shaft, 5 kN axle drive size B, pinion shaft, 6 kN axle drive size B, driving shaft, 5 kN The friction balance also allows the bearing to be oiled during the test and enables testing using high speeds and axial loads (Figure 7, right). Comparison testing was conducted on an angular-contact ball bearing type and two tapered roller bearing types. The bearings were lubricating using 20 ml/min of SAE75W90 transmission oil, an axial load of 2.22 kN was applied, and the bearings were tested at 2500 rpm. The oil inlet temperature was kept constant at 20\u00b0C. During the test, the temperature was measured on the stationary inner ring, and the bearing frictional torque was measured on the hydrostatic balance. Even though the angular-contact ball bearing as a larger outside diameter, the temperature was 10-15 K lower and the frictional torque was 50 % lower than was the case with the tapered roller bearings (Figure 9)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001158_s0020-7403(97)00082-9-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001158_s0020-7403(97)00082-9-Figure1-1.png", + "caption": "Fig. 1. (a) Roller contact, (b) Equivalent roller bearing line contact.", + "texts": [ + " and - 2 NOTATION inlet distance from the center of contact, m specific heat of lubricant, J/kg K coefficient of friction friction force per unit tength of Roller, N/m film thickness, m minimum film thickness, m dimensionless film thickness, h /R thermal conductivity of lubricant, W/m K lineal mass flux, kg/m s Legendre polynomial, ith order [Po = 1; PL = ~; P2 = (3~ 2 - 1)/2; P3 = (5~ 3 - 30/2] pressure, Pa equivalent roller radius, m length of roller, m slip or slide-to-roll ra:io (u-2 - u 2 ) / u , temperature, K x-wise lubricant velocity, m/s average rolling velocity (u-2 + u2)/2, m/s lateral coordinate in ~:he direction of surface motion, m load capacity per unit length of the roller, N/m dimensionless load capacity, Whmin/24LRu,r/o coordinate perpendicular to gap mid surface, m thermal diffusivity of lubricant, m2/s fluidity functions (see Appendix A) fluidity functions (see Appendix A) dimensionless coordinate transverse to film, 2z/h lubricant viscosity; Pa s lubricant viscosity at inlet, Pa s temperature viscosity of lubricant, K - t fluidity functions (see Appendix A) fluidity (reciprocal viscosity), I/r/, l /Pa s lubricant density, kg/m 3 dissipation function refer to Fig. 1 * Graduate student. Presently: Executive (Design), Mukand Dravo Wellman Ltd., P.O. Kalwe, Thane 400 605, India. 603 604 M.K. Ghosh and K. Gupta I N T R O D U C T I O N Estimation of minimum film thickness for a given operating condition of load, speed in rolling element bearings of jet engines is of great significance in aerospace technology. Contacts in rolling element bearings such as roller and ball bearings are characterized as line or point contact, respectively. More often, the contact operates in the regime of elastohydrodynamic lubrication", + " The momentum equation for non-inertial laminar lubricating films and the corresponding energy equation are, respectively, as follows: ( v~ = ~ ~ ~ : j (1) ~T 0 ( ~ T ) pcpa ~x - Oz K ~ z + c~, (2) where q~ = ~/(t/~ti/0z) 2 is tlhe viscous dissipation function. Along with Eqns (1) and (2), the following mass continuity equation for an incompressible fluid must be satisfied: V. fi = 0. (3) Hydrodynamic lubrication of line contacts 605 The :numerical solution to the flow field is sought by sampling the velocities, pressures and temperatures over chosen grid points and appropriate physical laws mentioned above are satisfied through an algorithm to which these values are interlinked. For a given geometry of line contact shown in Fig. 1 which represents an equivalent roller bearing, the temperature variation across the film is represented by a Legencre polynomial of order N, PN(0 and the sampling points then are N Le, batto points. It can be shown that N such internally selected points permit exact numerical integration of a polynomial of order 2N + 1 over range - 1 < ( < 1. Thus, f a r ( 0 d ( = 2WkTk --1 and for N = 2 the Lobatto location ~k and weight factor Wk are as follows and includes end-point values (Table 1). Therefore, if end-point temperatures are known, then it requires only two interior Lobatto point temperatures to be determined" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003918_j.jmatprotec.2005.02.163-Figure15-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003918_j.jmatprotec.2005.02.163-Figure15-1.png", + "caption": "Fig. 15. Swivel bar.", + "texts": [ + " The existing solution needs energy during the whole test to fix the plug, despite the fact that compressed air is cheap. From the energy point of view, it would be better to fix the plug with a spring and insert energy to the system only for releasing the plugs. But the difference between the normally open and normally closed is not very considerably. A simple bevelling of the fixing element should provide the function to remove the element by inserting the plug to the inlet. In the initial state the element is located inside of the inlet powered by a spring. The first realisation of the swivel bar is shown in Fig. 15. The rotating axis is placed in the centre of the plug. The dotted lines show the centre plane. The mentioned triangular shape a p t t s has to remove the bar before it is able to enter the inlet. If the plug touches the bar, it starts a rotational movement and it will leave the empty space of the inlet. The centre plane is again symbolised by dotted lines. As soon as the plug reaches its final position, the spring presses the bar into of the inlet, the plug returns to the initial position and the plug is fixed by the swivel bar" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002210_s0022-5096(99)00047-2-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002210_s0022-5096(99)00047-2-Figure1-1.png", + "caption": "Fig. 1. The geometry used in our experiment. Right top is a side view and the bottom is an end view.", + "texts": [ + " In the second case, on the other hand, a single developable cone is nucleated and the energy of the cone is measured. We will show that the study of mechanical properties of a single singularity, namely its energy, is necessary to understand the crease selection in a buckled cylindrical panel. The simplest curved surface one can crumple, is a panel (thickness h, width l, and length L ) which is curved along the arc of a circle of radius R and opening angle b. We investigate the crumpling of such a surface which has been exposed to axial compression as shown in Fig. 1. Other surface shapes can be considered, such as hyperbolic or parabolic shapes, but these shapes fall into the same class as the cylinder and we expect no qualitative di erence with respect to the nucleation of creases. The cylindrical panels used in this study are 0.1 mm thick sheets of DHPcopper. The thickness of the panels is uniform within 1%. Such panels are manufactured by cold rolling a thicker plate. This process produces residual stresses within the bulk of the plate. This stress leads to an anisotropy in the material properties with respect to the two in-plane directions", + " The jump in 4 The de\u00afection due to vertical displacement of the border is given by the expression w R2q=Eh 1\u00ff n2 e\u00ffmx cos mx, where n is the Poisson coe cient, h is the thickness of the plate, q is the bending moment and m 4 3 1\u00ff n p 2 =R2h2}, the wavelength of perturbation is l 2p=m0 Rh p , which is negligible with respect to the crease length and the panel length. equilibrium position as the crease appears involves a release of elastic energy, and thus explains the rapidity of the buckling process justifying the sub-critical character of the transition. In Cha\u00f5\u00c8 eb and Melo (1997), open cylinders with a opening angle b, as shown in Fig. 1, were discussed in contrast to the full cylinder case. The dependence of the crease properties on l, R and the additional parameter b was discussed (Cha\u00f5\u00c8 eb and Melo, 1997). In the following section we will discuss the selection of the crease length X as R is varied for an almost closed cylinder, i.e. b22p. The panels we compress can be divided in two categories: small opening angle panels with b < 2 rad. and large opening panels with b > 2 rad. For small opening angles, a single crease nucleates and occupies nearly the entire panel width" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002560_icsmc.1995.538488-Figure2-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002560_icsmc.1995.538488-Figure2-1.png", + "caption": "Figure 2. Coniigurahon of the f o r e m boundary singularity for PUMA manipulator", + "texts": [ + " Thus, the singularities can be identified by checking the determinants of the two 3 x 3 matrices O J l l and O J 2 2 as follows. 2.2.1 Forearm Singulari ties The forearm singularities can be identified by checking the determinant of the matrix O J l l in (12) as h t ( O J 1 1 ) = -aa(drCa - a3S3)(d&3 + a2C2 + a3C23). (15) There are two conditions, denoted as 76 and vi , for forearm singularities. One is the boundary singularity [16] with (16) A 7 6 = d4C3 - a3S3 = 0 and the other is the interior singularity [16] with (17) Referring to Fig. 2, it can be seen that the boundary singularity happens when the Wrist point is located on the x2 axis. In this configuration, the tasks in x2 direction and z2 A 7 ; = d4S.23 + a2C2 + a s C 2 s = 0. 441 1 direction are dependent. This key point can also be identified by viewing the linear velocity of the Wrist in coordinate 2aS Then, 2J11 will have the following structure where x's represent some values other than zero. As for the interior singularity, referring to Fig. 3, it h a p pens when the Wrist locates at the yl -a1 plane" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002529_00037-1-Figure11-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002529_00037-1-Figure11-1.png", + "caption": "Figure 11 Wyoming modified two-rail shear test fixture.", + "texts": [ + " An equally significant limitation in the past has been slipping of the rails, which can negate the test. However, this problem has been significantly reduced with the introduction of tungsten carbide particle surfaces as described earlier (WTF, 2000). A very promising modification of the tworail shear fixture has recently been introduced (Hussain and Adams, 1998, 1999). Roughened rails are clamped onto the specimen, but the bolts do not pass through the specimen, thus eliminating the need for clearance holes, and the associated preparation cost. The fixture is shown in Figure 11. It essentially eliminates the slipping problem. While less research has been performed on the three-rail shear fixture, the problems are very similar. (iii) Double-notched shear The ASTMD 3846 specimen configuration is shown in Figure 12. The Modified ASTM D 695 compression fixture (see Figure 8) can be used to apply the shear loading. The specimen can also be loaded in tension (Hercules, 1990b). Although an ASTM standard, this test specimen is justifiably criticized because of the severe normal as well as shear stress concentrations induced at the bottoms of the notches" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002436_bf01213538-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002436_bf01213538-Figure1-1.png", + "caption": "Fig. 1.", + "texts": [ + " The Euler-Poisson equations are used for the presentation of that motion. Three first integrals of these equations are well known. To solve this problem completely one further first integral has to be found. The main purpose of the paper is to obtain a necessary and sufficient condition for some function F(og~, (o2, w3, Y~, Y~, Y3) to be a new first integral of Euler-Poisson equations. Given a rigid body B of mass M, weight W = M g (g - - acceleration of gravity), center of mass C and a point P fixed in tha t body (Fig. 1). Le t A1, A2, As be the principal moments of inertia of B about P and let the body fixed frame ( P x l x 2 x 3 ) - - basis ( i , i2, i8) - - coincide with the principal axes of inertia through P. Le t the point P be a t tached to the origin 0 of a spatially fixed cartesian coordinate system (Oy~y~y3) - - basis ( j , j ~ , j 3 ) - - where j3 is directed upwards such tha t there holds for the weight vector w = --Mgj~. (1) Let the projections of the unit vector ja on the axes of PXlX~X3 be yi, y2 and Y3- Thus ja = Ylil ~- 72i2 ~- y~i3" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002644_2003-01-3743-Figure3-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002644_2003-01-3743-Figure3-1.png", + "caption": "Figure 3: Axle insert with hypoid teeth pinion and ring gear and normal straight-cut bevel gear differential, Source: ZF.", + "texts": [ + " Axle differentials are designed with spiral bevel gears in order to achieve a high load carrying capacity combined with quiet operation. The bevel gears require rigid bearing supports to obtain the proper tooth contact and uniform load distribution across the width of the teeth [2]. The axle differential's limited envelope often requires overhang pinion supports although in some variants the pinion has locatingnon-locating bearings supports. In these cases, two tapered roller bearings are braced against each other, and an additional cylindrical roller bearing behind the pinion acts as a non-locating bearing (Figure 3). The cylindrical roller supports increasingly high loads as transmission life progresses because preload loss occurs in the tapered roller bearings. The non-locating bearing is thus more of a semi-locating bearing that is installed due to the inadequacy of the tapered roller bearings. Axle differentials should be as compact as possible to save weight and ensure that the vehicle\u2019s ground clearance is not unnecessarily restricted. The size of the ring gear, which in turn depends on the torque to be transmitted, determines the size of the transmission (see Figure 3). Due to the high torques that must be transmitted and the limited design envelope, axle differential bearing loads are very high. Due to their contact angle, tapered roller bearings can support both axial and radial forces. Their function, however, is based on the fact that they are axially arranged. This is why it is important to brace at least two bearings against each other. The different contact angles on the outer and inner ring create a force that presses the tapered rollers against the inner ring rib" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000585_rob.4620121004-Figure5-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000585_rob.4620121004-Figure5-1.png", + "caption": "Figure 5. The auxiliary structure.", + "texts": [ + " Such a circle has center B2 and radius L2 given by: Mutually orthogonal unitary vectors a2 and b2 are now introduced, both fixed to the platform and forming with (By - Bi) a right-handed orthogonal system. Accordingly, the position of point A2 with respect to W,, can be expressed as: (A2 - B 4 , = L2(a2p cost$ + blp sin&) (6) where 82 is the counterclockwise rotation angle about (&\u2019 - Bi) that superimposes a2 on vector Clearly, performing the DPA of the (5-5)B fully parallel manipulator-once the leg lengths are given-is equivalent to finding the assembly configurations of the auxiliary structure represented in Fig. 5, where two revolute pairs are introduced in lieu of the couples of legs converging at points B1 and A2 of the (5-5)B fully parallel manipulator represented in Figure 4. Links AIBl and A2B2 of the auxiliary structure, respectively connected to platform and base by spherical pairs, are orthogonal to the axis of their extremity revolute pair. (A2 - B 2 ) . 3. THE FIRST SUBSET OF CONSTRAINT EQUATIONS With reference to the auxiliary structure of Figure 5, the guidelines for laying down the constraint equations can be summarized as follows. The position of the platform with respect to the base is first expressed as a function of as reduced a number of parameters as possible. This allows a minimum number of constraint equations to be imposed, which, in turn, contributes to keeping the complexity of the mathematical treatment to a minimum. The first parameter needed for describing the position of the platform is angle el, which governs the position of point B1 with respect to reference Innocenti: Analytical-Form Direct Kinematics 665 frame WI, (see Eq. (3) and Fig. 5). The orientation of the platform is then parametrized by referring to a couple of vectors, namely, (B2 - B , ) and (A2 - B2) . These vectors, which are generally non-collinear, have components in W,, and WI, depending on a reduced number of parameters. Actually, vector (B2 - B1) is constant in W,,, while its three components in Wi,: are a priori unknown. Vector (A2 - B2) depends only on angle tI2 when referred to W,, whereas, basing on identity it depends on parameters O,, M I , u2 , and u3 when referred to WI, (components of (A2 - Al)t, are constant, while (Bl - A])/, and ( B 2 - Bl)/, are given by Eqs" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003976_j.jmmm.2005.10.076-Figure2-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003976_j.jmmm.2005.10.076-Figure2-1.png", + "caption": "Fig. 2. Configuration of a four-particle cluster under an applied field; here j1 \u00bc j4 and j2 \u00bc j3.", + "texts": [ + " Reflecting this, the magnetization curve of the cluster at h \u00bc h undergoes a sharp bend, after which magnetization turns asymptotically to saturation, see Fig. 1. To explain this magnetic behavior, consider the configurational change of a square four-particle cluster under the action of a field. In the initial state (H \u00bc 0) all the moments are oriented under the angles j1 \u00bc \u00bc j4 \u00bc p=4 to the sides of the square. At a weak field, the orientational pattern rendering the energy minimum is symmetrical inside the pairs grouped with respect to the field direction, keeping j1 \u00bc j4 and j2 \u00bc j3, see Fig. 2. However, to get parallel with the field, the moments of the \u2018\u2018favorable\u2019\u2019 pair have to cover the angle about p/4 while the moments of the \u2018\u2018unfavorable\u2019\u2019 pair have to sweep 3p/4. From the set of equilibrium equations minimizing the energy of a four-particle cluster, a relationship j1(j2) was obtained which has a sharp break at jn 2 0:91 p corresponding to the field strength h 0:62. At hXh the configuration of the system simplifies, and is characterized by just a single angle due to the equality j2 \u00bc j1 \u00fe p, see Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001240_6144.774758-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001240_6144.774758-Figure1-1.png", + "caption": "Fig. 1. SmartPen system [1] (courtesy of IEEE).", + "texts": [ + " For example, in a S&A device used in a torpedo, a hydrostat senses pressure and only unlocks the slider if the device senses it is at an appropriate depth. The S&A device also includes independent optical verification of its position. Section III of this paper discusses the S&A device in detail. A Computer Writing Tool\u2014A \u201csmart\u201d pen that allows simultaneous writing on paper and on a computer is under development [1]. The pen uses a standard ink cartridge mounted in a shaft that is connected to a force sensor (Fig. 1) and is monitored by a tilt sensor. A rechargeable battery pack is included to power the device, and at the top of the pen is a printed wiring board containing multiple chips comprising a radio transmitter that communicates with a computer. The position of the pen on the writing surface is computed from information obtained from the force and tilt sensors, and serve as input for displaying a bitmap of the written text on a computer display. Solid drug delivery device\u2014Solid drugs are normally delivered by implanting a highly concentrated dose of a drug subcutaneously so that it is slowly dissolved by the body thereby releasing small amounts of the drug" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003638_s021812740401103x-Figure2-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003638_s021812740401103x-Figure2-1.png", + "caption": "Fig. 2. Admissible angular velocities \u03c9 for given (h, l) in the Euler case (A1, A2, A3) = (2, 1.5, 1). The values (h, l) are, from left to right, (0.33, 1), (0.42, 1), (0.6, 1). In the two pictures on right, one quarter of the ellipsoid has been cut away to give a better impression of \u2202Vh,l.", + "texts": [ + " Assume first that (h, l) is taken from the yellow region in the bifurcation diagram. The smallest possible value of g2 is obtained when the space-fixed vector l points in the directions of \u03b3 or \u2212\u03b3: then g2 = l2. The invariant lines are two S1-curves encircling the e1-axis in \u03c9-space. All higher values of g2 up to g2 = 2A1h can be realized by adjusting the angle between l and \u03b3 such that \u3008l, \u03b3\u3009 remains equal to l. The union of the corresponding lines in \u03c9-space is the set of two disks shown in the left part of Fig. 2. Together they are Vh,l = \u2202Vh,l. For (h, l) taken from the green region of the bifurcation diagram, the smallest possible g2 defines two S1-curves encircling the e3-direction. Again, all higher values up to g2 = 2A1h can be realized by bending l sufficiently away from the direction of gravity. The resulting Vh,l = \u2202Vh,l is an annulus, see the middle part of Fig. 2. When energy is increased into the blue region, all points on the energy ellipsoid in \u03c9-space are admissible for an appropriate value of g2 > 0, see the right part of Fig. 2. The general invariance of l2 is a unique feature of the Euler case. In almost all other cases of rigid body motion in a gravitational field, the torques associated with r 6= 0 make the angular momentum vacillate between local minima and maxima of l2 (for exceptions see the last sentence of Sec. 3). The admissible points in \u03c9-space then form a threedimensional set Vh,l. 1.3. Lagrange\u2019s case The Lagrange case is considerably richer in the structure of its energy surfaces and their projections to \u03c9-space" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003675_j.triboint.2003.11.004-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003675_j.triboint.2003.11.004-Figure1-1.png", + "caption": "Fig. 1. Schematic illustration of cams.", + "texts": [ + " Better understanding of elastohydrodynamic and materials technology leading to significantly increased rolling bearing fatigue life and similar advances in the areas of cams and gears was reported by Jost [7]. It is well known that tribological processes occur when there is relative motion between two contacting elements, such as on cams and other tribomechanical systems. The consequence of the development of tribological processes in the operation of the cams is the wearing out and failure of the cams in contact. This tribomechanical system, shown in Fig. 1, consists of three elements, these being : element 1 (lever transmission), the element 2 (cam) and the element 3 (lubricant) in which the contact between the two former elements is realized. The cams have poor conformity between surfaces, very small contact areas and very high unit loading (h\u2014film thickness, B\u2014width of the zone). The tribological behaviour of the cams is influenced by a number of factors that are interrelated in complicated ways. These factors can conditionally be divided into the external and internal ones" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001951_s0389-4304(01)00102-3-Figure6-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001951_s0389-4304(01)00102-3-Figure6-1.png", + "caption": "Fig. 6. Cross-sectional view of linkages and trunnions.", + "texts": [ + " Role of linkages The upper and lower ends of the trunnions supporting the power rollers can be positioned freely by means of upper and lower linkages according to the tilt of the trunnions. Needle bearings are provided between the trunnion and the hole into which a linkage is inserted. These needle bearings serve to reduce friction when the trunnions are tilted, and they also make contact with the convex surface of linkages, thus allowing easy vertical displacement of the trunnions accompanying the tilting of the linkages (Fig. 6). Traction force T in the driving force transmission unit is expressed as T \u00bc mN, representing the product of thrust force N and traction coefficient m. Since m is a characteristic the value of which is determined by the composition of the traction oil, it is necessary to apply correspondingly larger thrust force N in order to transmit greater torque. Increasing thrust force N in the half-toroidal CVT results in a greater thrust load that can cause the power rollers to jump out of the toroidal cavity between the opposing input/output discs" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000835_ip-cta:19971028-Figure17-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000835_ip-cta:19971028-Figure17-1.png", + "caption": "Fig. 17 Route planning at the point of intersection", + "texts": [ + "2 rig-z8g manoeuvre under SlSO and SIMO NN controllers To expect the ship to have minimal track error, when a distinct change in heading is planned without communicating this information effectively to the controller, is unwise. Effective communication between route planning and controller, see Fig. 10, will clearly depend on how much warning the controller has of the directidfi change and how the transition from one defined ship heading to another distinct heading is to be achieved. One possible approach is by selection of the anticipation distance [ll], e, and the allowable error, e, see Fig. 17. The arc PIP2 is circular and of radius R, a n - tred on 0. The point T corresponds to the turning point in the original ideal theoretical zig-zag. The straight line P I T (TI\u2018,) represents a distance, d, between T and the end of the anticipation distance. As R2 + dL = (R + e)2 and d = R tanq, then solving the resultmt quadratic equation in R2, subject to R is real and positive, requires e sec9 - 1 As Q, is specified by Y and X , in the defined zigzag manoeuvre and e is specified by the master, then R is known. Hence, a modified route plan with semibte transition from one fixed heading to the next fixed heading is deducible by the controller, whether SISO or SIMO. Typical values of R and other parameters are provided in Table 3 as a function of e, for tp = 18.57\u2019. This cp is consistent with X, = 2X = 3000m, and Y = 500m in Fig. 17. For the SIMO controller, we require qkd to be defined on the modified route. The obvious, but inappropriate, value of qkd would be the tangent to the cir- (34) R = IEE Proc -Control Theory Appl , Vol 144, No 2, March 1997 cular transition arc introduced. It is inappropriate, because, once the ship has taken up the implicit turning circle, it is unreasonable to assume that, from the point P2 onwards, the ship can immediately take up and sustain the constant heading qAW. Just as a transition arc was introduced to provide a more sensible route plan, so qkd, on the transition arc, must be modified so that the ship is capable of sustaining I)& after point Pz" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002329_s0378-4754(00)00225-1-Figure8-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002329_s0378-4754(00)00225-1-Figure8-1.png", + "caption": "Fig. 8. Ergonomic trajectory of the spray-gun: (a) spatial view of the trajectory and the trace of nozzle axis at the plane that contains the treated surface; (b) top view.", + "texts": [ + " It is expected that the secondary objective Js will be reduced and the quality will be kept above the prescribed level: Jq \u2265 J min q This step is described in the simulation example. First, one may observe that a human worker does not perform a rectangular trajectory like the one shown in Fig. 6a, but rather a curved one shown in Fig. 6b. This reduces the required accelerations and hence the pick torques. Next, one observes that a human worker tends to make the spray-gun\u2019s trajectory more \u201ccompact\u201d and thus reduce the amplitudes of joint motions. He (she) achieves this by deforming the planar trajectory from Fig. 6b in the way explained in Fig. 8. This is called the \u201cergonomic trajectory\u201d. The shape of such a trajectory can be parameterized by using l and dy as shown in Fig. 8. Different trajectories will produce different coverage profiles characterized by different quality levels (Jq) and will consume different amounts of energy (Js). Some examples of coverage profiles are shown in Fig. 7. Besides the \u201cideal\u201d planar motion (case (a) in Fig. 7), two ergonomic trajectories are studied (cases (b) and (c)). The trajectory parameters l and dy should be tuned so as to achieve the constrained minimization of the secondary objective. Fig. 9 shows how the quality-function Jq and the cost-function Js depend on the parameters l and dy " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002407_robot.1986.1087402-Figure2.1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002407_robot.1986.1087402-Figure2.1-1.png", + "caption": "Fig. 2.1 shows some examples of 3D-robot models.", + "texts": [], + "surrounding_texts": [ + "The development of a simulation system for robots is described. The system includes interactive construction of 3D-graphic models, kinematic and dynamic analysis, simulation and animation of the robot on a high resolution graphics terminal.\n1 . Introduction\nSimulation is a valuable tool in designing and testing technical systems. Most commercially available CAE systems lack the special capabilities needed to simulate robots. This led to the descision t develop a special simulation system for obots called ROBSIM (ROBot SIMulation system). The development is done in 3 major steps\na. implementation of basic robot simulation capabilities as :\n- interactive construction of robots ( CAD 1 - creation of databases for the mechanical and graphical data - creation of transformation matrices - creation of a mathematical model,\n- creation of the equations of motion - programming the robot task in a high\n- creation of Cartesian trajectories - manual programming via teach in - test of different approaches of\n- simulation and 3D display of the\nmodel reduction\nlevel language\ncontrol strategies\nmovement\nb. integration of a world model\nc. use of AI-methods to handle complex problems (intelligent robot systems)\nThis paper describes parts of step a, the development of basic robot simulation capabilities. ROBSIM is divided into several modules. The modules ROBCONS, ROBPRO, ROBDYN and ROBGRA are described in this paper.\nROBSIM is written in FORTRAN77 and runs on a 68000 based microcomputer system under UNIX. A Raster Technologies model ONE/380 is used as a graphics display. ROBSIM is commercially available. Installations are made on different computer systems e.g. machines of Digital Equipment, Data General etc. . 2. ROBCONS - Module For the Interactive\nConstruction of Robots\nMost robots are build out of translational and/or rotational joints. The kinematic structure of a robot is determined by the sequence of joints and links, in the sequel called segments. This sequence and the geometric dimensions determine the different robot types and workspaces. Analysis of the structure of robots provides the fact that graphic models of robots can be build using a few basic segments . Construction of a 3D graphic model of a robot with ROBCONS is done interactivly using a graphics workstation. It consists out of a high resolution\n1859 CH2282-2/86/0000/1859~01.00 O 1986 IEEE", + "graphics display combined with an alphanumeric terminal. The user selects his desired segments from a 3D-segment database. Construction starts with selection of the robot base. Then segments are selected, specified and added to the robot. After selection of a segment ROBCONS checks if the segment is allowed at this stage of construction. If the check fails a list of possible choices is displayed. For a segment 3 numbers or less have to be specified e.g. heigth, radius. After specification of the geometry the actual segment is created. All segments are stored as structures with geometrical relationships with respect to a local coordinate system. This internal data is recalculated for the real segment. ROBCONS places the segment automatically using the information about the two segments which have to be combined. The graphic model of the robot is stored in a robot-database which describes the robot graphically and kinematically.\n3. ROBPRO - Module for Task Programminq\nROBPRO is the module to program a robot task. Programming can be done either by teach-in or by creating a robot program. In teach-in modus the robot is driven using a virtual control panel. It is possible todrive single axes or t move the hand in Cartesian coordinates. The programmer can drive the robot to a desired position and store the corresponding frame. Creating a robot program is done in two steps. First all positions have to be specified as frames. Then the sequence of robot commands is given. ROBPRO creates a program file and writes all commands combined with the necessary information into this file. It is possible to specify frames with respect to a chosen coordinate system. In order to drive complex trajectories the specification of via-points is possible. These points are included in the MOVE command.\nAccording to the interpolation formula\nF(r) = D(r) F 1 (3.1 1\nwith D(0) = I, D(1) =F2F1 r = 0,. ..,I\n-1\nwith F as start frame and F as\ndestination frame intermediate frames F(r) are calculated. Fig. 3.1 shows a straight line MOVE, keeping the orientation of the hand constant. Fig. 3.2 shows a MOVE from the same start frame to the same destination frame but using two viapoints. The use of ROBPRO shall be demonstrated by an example, a car wheel assembly.\nFig. 3.3 shows the workcell as a topview. The robot has to move from its initial position (PI) to the wheel (P2), grasp it, make a quick move in front of the car (P3) and make a precision move to position P4 to fix the wheel at the car. First all positions have to be given. They can be determined for example by using the virtual control panel. Then the command sequence has to be specified. The robot program is created and the necessary trajectories calculated.\n1 2", + "Fig. 3.4 shows the simulation of the trajectories. Lnaa-' 9 &+e P 4\nP 3\n4 . ROBDYN - Module For Dynamics Modellinq -\nFig. 3.2 MOVE using viapoints\nROBDYN is the module for modelling the robot dynamics. The robot is assumed to consist out of rigid 1inks.Common approaches to robot dynamics are Lagrange/2/ and Newton-Euler /2/ equations. ROBDYN includes the Lagrangian approach implemented recursivly, using the algorithm of Hollerbach / 1 / and Kane's approach / 3 / . For a six degrees of freedom robot a comparison of computing costs of different methods is given i table 1 . Kane's method seems to be the most efficient." + ] + }, + { + "image_filename": "designv11_65_0001008_0734-743x(92)90486-d-Figure4-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001008_0734-743x(92)90486-d-Figure4-1.png", + "caption": "FIG. 4. The arrangements of Robins\" ballisuc pendulum", + "texts": [ + " This monograph of Robins consists of two chapters only, with a relatively lengthy Preface of some 53 small pages, giving a short Recital of the Rise and Changes in Military Architecture. (This is much concerned with the invention of the Angled Bastion for forts, a still 'live' topic and one which has been the subject of authoritative address [ 15].) The first chapter is on the force of gunpowder, principally the ballistic pendulum and the second on the resisting power of the air; it was brought before the international public, through Fig. 4. The former subject compelled world attention from foreign ordnance departments and, enlarged, became a necessary piece of equipment for research and development centres in France, the U.S.A., and elsewhere. The author endeavoured to discuss the two chapters at length [ 16,17] and where Robins proceeds by a method which uses \"'Proposition\" and \"Scholium\"; this method reveals Robins' attachment, (following Newton), to 'older' geometrical methods as opposed to 'new' algebraic ones. Outstanding points in his writing are that he observes, \"all burning bodies destroy great quantities of air ", + " Put simply, firing musket balls of mass m and speed v normally at a suspended heavy wooden block (metal backed) of mass M and recording the swing of its centre of gravity through vertical height h leads to an initial combined speed of ball and block Vof (2gh) 1/2, where g is the gravitational constant. Then, through the principle of the conservation of linear momentum, we find, v = ( M + re)Vim; this sufficiently outlines the operational principle of the machine though many refinements can be and were included or added. The weight of bullets in Robins' first investigations were 1 / 12 lb and were shot at various velocities up to about 1700 ft/s, and were measured soon after discharge using this device. The size of the early ballistic pendulum is conveyed in Fig. 4. Several rules were laid down for the safe use of this pendulum, see p. 214 of [16]; they included, not scattering gunpowder and concerning the desirability of not breaking-up the pendulum block because pieces of metal backing which could spall would cause injury. Indeed at this point Robins declares, after long experience with his machine that, \"There are Dangers to the braving of which ... in Philosophical Researches no Honour is annexed\"! In I-6] the writer first reproduced a copy of a hand-written account of a French ballistic pendulum and the investigation in which it was used" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003842_iros.1992.594476-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003842_iros.1992.594476-Figure1-1.png", + "caption": "Fig. 1. Sampled trajectry", + "texts": [], + "surrounding_texts": [ + "Proceedings of the 1992 IEEE/RSJ International Conference on Intelligent Robots and Systems Natural Robot Programming System Raleigh, NC July 7-10,1992\nTaro Harima\nMitsubishi Electric Co.\nAbstruct-When applying robots in small scale produc-\ntion, the bottleneck is programming. There are three conventional ways to generate robot programs: by walk through, lead through and off-line programming. Walk through programming has limited application. Lead through programming can be time consuming and inefficient, while current off-line programming systems need large capital investment.\nThis paper introduces a robot programming method\nthat provides a natural interface between man and robot at a low cast. The system translates the human motion into a robot language, using the human hand as a teaching device, so that users never use the keyboard nor CRT.\nHarry West\nCenter for Information Driven Mechanical Systems\nMassachusetts Institute of Technology\n~.INTROSUCTION\nRobot task programming methods are classified in Table 1, Walk though, Lead though and off-line programming [l].\nThe walk through programming method is particularly\nsuitable for robots which have to perform complex continuous path motion such as spray painting and arc welding. A robot is programmed by an operator physically leading the robot through the desired path. the controller records position information from each of the axes. Once the programming is completed the robot can continually repeat the path it has been taught. This method is easy to use, but it has a disadvantage which is that it is only posible for back driveable robots.\nThe most common programming method on the manufacturing floor is the lead through programming method. in this", + "method the operator \"teaches\" the robot task by use of a teach-\ning box. however the teaching box does not nahirally correspond to robot motion. In addition an operator has to learn the complicated robot language to write task programs. Teaching is too complex to use easily.\nEvidently off-line programming is a part of the legacy\nfrom numerically controlled machines, NC. But unfortunately NC programming is not completely transferred to robotics. Robots don't exhibit precision in the same class as NC machines, because of their serial articulation, speed of operation and large workspace.\nThis paper discusses the result of a NRP (Natural Robot Programming) project in which the authors have developed ideas and programming methods for robot task programming. A good interface between man and robot presupposes a high level service from the control and supporting software [2]. [3],\n141. The teaching and programming method described in the paper can be used for on-line programming on real manufacturing floors at a low cost. The key elements are a teaching device and a task program generating method both of which the operator can easily use on-line during jobs.\n11. TEACHING\nThe human arm is used for a teaching. We measure the human arm position by a glove with ultra-sonic transducers and three receivers. The glove gives a natural interface between the operator and robot.\nBecause the computer controls the robot from the human arm position, the operator can execute his task while the robot follows his arm motion. Simultaneously the computer records the robot motion as the teaching data. When the programmer finishes his task, he has finished teaching his task to the robot.\nIn the NFtP system, the operator doesn't have to be familiar\nwith a robot task programming language. The system can automatically acquire and generate the task program.\n111. TASK PROGRAM GENERATION\nBecause the teaching data sampled from the glove consists of many teaching points (Fig. l), we should remove the middle points where robot accelerates and decelerates frequently, and reduce the robot travel time. (Fig.2)\nThe characteristic points which are necessary and suffi-\ncient for the task are as follow: (Fig.3)\n1 Corner points 2 Reversal points 3 Shift point\n75 1", + "d Comer Point Reversal Point Shift Point.\nThe shape of the trajectory with a comer point, a shift\npoint and a reversal point has 3 bend points in the Timeapace shown in Fig.6.\nFig.3. Characteristic Points comer point\nThe corner points occur where the robot changes direction\nand makes a bend in the trajectory. The reversal points are where robot goes back the way it came, and the shift points are where the robot changes the velocity magnitude. These points are identified in time-space, which is 4 dimensional space including the time axis. (Fig.4 & Fig.5)\nposition\n:versa1\nPoint\naxes point toward the eigenvectors of the dispersion matrix based on the teaching points, and the center is the average of them. (Fig.7)\nThe fitting line is the major axis of the ellipsoid body which is the first eignevectorm, of the dispersion matrix A of these teaching points. the first eigenvalue is the length of the major axis.\nA=E(SXk) here x=x-X\nX is the center of the teaching points.\nSimilarly the second eigenvectom, points toward, the second axis that is the most scattered direction. the second" + ] + }, + { + "image_filename": "designv11_65_0003128_robot.2002.1014257-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003128_robot.2002.1014257-Figure1-1.png", + "caption": "Fig. 1. A two-wheeled mobile robot .", + "texts": [ + "(40) XO 1 0 In the new coordinate and by the definitions of q, and q1 , the mobile robot system can be transformed into a system of the form (29) where the matrix-valued functions AA(qo,ql) and M(q0,q1) can be given in the foIIowing . Define the new coordinates by [4] cos6 sin6 0 A: 1. (41) -(1 +p3sin(r) \"1. A&[ +osr-(l-cosr) -(l+jj)sinr -\"+pcosr-l+cosr 0 Thus, Theorem 2 car1 be applied to guarantee a globally exponential practical robust stability for small values of Aq, , Ap,, A r . This is achieved in present literature yet. Note that a two-wheeled mobile robot (see Fig. 1) can be modeled as a system of the form (39). Indeed, let W, and w, denote the angular velocities of the left wheel and the right wheel, respectively. Let R and L denote the unknown radius of the rear wheels and the distance between thein, respectively. Define . Then, it can be W w, +w, and v=-2 2 modeled as a system of the form (39) with p = R and [4]. For simulations, suppose 2R q=L p = R = 0.11 (m) and L = 0.5 (m) . Then, q = 0.44. Let (p,,, pa) = (0.050.25) and (qmn, 4) = (0.2,0.5) . Then, p o = 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003483_imece2004-59492-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003483_imece2004-59492-Figure1-1.png", + "caption": "Fig. 1: Reference systems", + "texts": [ + " This approach allows one to accurately evaluate the variable global stiffness, which depends on the gear position, materials and teeth geometry. The variable stiffness gives rise to a parametric excitation, which is periodic; therefore it gives rise to a Mathieu type instability. A Fourier series of the variable stiffness is obtained from the numerical data evaluated with the FEM model. The dynamic model is analyzed by means of perturbation techniques [14]. MESHING EQUATIONS Let us consider the profile generation by means of an envelope of a rack profile 1, see fig. 1. Three references systems are introduced: -Sf (Ofxfyx) fixed reference; -S1 (O1x1y1) moving respect to Sf; -S2 (O2x2y2) moving respect to Sf; where \u03c31 is fixed in S1 and \u03c32 is fixed in S2. \u03b8 is local variable that identify a point on the curve \u03c31 and \u03c6 is local variable that identify a point on the curve \u03c32. The parameter \u03c6 depends on \u03c6; P1 and P2 are the same point P represented in the reference systems S1 and S2. nloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 02/07/2016 Te Let us consider a regular plane curve 1 describing the rack tooth profile", + " [13], the following sufficient conditions for the existence and regularity of the curve must be respected: 1) The family of curves \u03c32 must have C2 regularity for both \u03b8 and \u03c6 parameters: ( ) ( )baRECPO ,,,, 2 22 \u2208\u2286\u2208\u2208 \u03c6\u03b8\u03c6\u03b8 (6) 2) \u03c31 must be regular: ( ) 0011 \u2260 \u2202 \u2202 \u03b8 \u03b8 PO (7) 3) If a generic point (\u03b80,\u03c60) verifies eq. (1), then it cannot be a singular point and must verify the following equation: ( ) ( ) ( ) 0,vt t PO 00011 \u2260\u03c6\u03b8+\u03b8 \u2202 \u2202\u03b8 \u03b8\u2202 \u2202 \u03c4 (8) where ( )t\u03b8\u03b8 = provides the contact point on \u03c31 while \u03b8 is varying. APPLICATION OF THE ANALYTICAL ENVELOPING METHOD Sliding velocity From fig. 1 one obtains that: [ ]TRRjRiROO 001121 \u2212=\u2212= \u03c6\u03c6 (9) and [ ] ( )t dt dT \u03c6\u03c9\u03c9\u03c9 == ,000 (10) the sliding velocity given by eq. (4) becomes: ( ) ( ) ( ) 0POPO 00 kji 0RR 00 kji iRPv 1y11x1 111111 1 \u03c9+ +\u03c6\u2212 \u03c9+\u03c9=\u03c4 (11) Roto-translation matrix The rotation matrix is given by: ownloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 02/07/2016 Te ( ) ( ) ( ) ( ) ( ) \u2212 = 1000 0100 00cossin 00sincos \u03b7\u03b7 \u03b7\u03b7 \u03b7zR (12) where \u03b7 is a generic rotation angle. The translation of point P can be described by the translation matrix: PT P P P t t t tP tP tP tP t z y x z y x zz yy xx = = + + + =+ 11000 100 010 001 1 (13) where t is a vector indicating the translation" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002606_0076-6879(89)74042-8-Figure7-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002606_0076-6879(89)74042-8-Figure7-1.png", + "caption": "FIG. 7. Scheme of the measuring chamber (shown actual size). At right, a twice-c~la~ed", + "texts": [ + " Individual cell chains are drawn into the microfunnel from cells sedimented on the slant of an acrylic glass wedge located in the measuring chamber by means of a plastic syringe (5 ml volume) connected to the microfunnel by plastic tubing. The microfunnel is operated by means of a Leitz micromanipulator, and the movement is controlled under a microscope fitted with a long working-distance objective (magnification: \u00d7 25). The measuring chamber is constructed from acrylic glass in such a way as to make a horizontal approach for the microfunnel and the measuring microelectrode possible from each of the two sides of the chamber (Fig. 7). The chamber is covered by a cover glass so that the bathing solution (see below) is kept in the chamber by adhesion despite both sides being open. Yeast cells, in a drop of suspension, are pipetted onto the slant of the wedge and allowed to sediment (within 30 sec). Then, bathing solution from a container located above is slowly circulated through the chamber, eventu- ally dropping into a collecting flask in front of the microscope. Continuous flow is facilitated by connecting the first reservoir, the measuring space, and the collecting flask with filter paper strips in a flat groove" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001713_1.1317233-Figure5-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001713_1.1317233-Figure5-1.png", + "caption": "Fig. 5 The coordinate frame for the finger", + "texts": [ + " 122 \u00d5 605 7 Terms of Use: http://www.asme.org/about-asme/terms-of-use Downloaded F exerted by the bellows. Referring to each bellow, Fi , i51,2,3, are the forces from the voice coil motors and M i , i51,2,3, are the elastic torques depending on the angular deformation. For each finger in the gripper, two different coordinate frames can be considered, the inertial frame and the reference one. The first frame is fixed with respect to the gripper, and the second one is fixed with respect to the finger itself ~refer to Fig. 5 for the definition of the finger reference frame!. The forces Fi exerted by the voice coil motors are all directed along the z axis of the reference frame, and their application 606 \u00d5 Vol. 122, DECEMBER 2000 rom: http://dynamicsystems.asmedigitalcollection.asme.org/ on 08/15/201 points form a simplex of vectors in the frame of the finger. Therefore the vectors t and M of the torques applied to the system result to belong to the ~x,y!-plane within the reference frame, where they form a simplex of vectors", + " The system to be controlled belongs to the class of the systems, such as tendons arms and jet actuated vehicles, for which it appears suitable the use of a robust control technique. A possible solution could be the use of the so-called simplex-based sliding mode technique which has been presented in the literature @7\u20139#; we refer to these works for a general treatment. In this paper, it is presented and proved the convergence of the method for the considered case. The vector t of the torques, which can be applied to the finger, results to belong to the ~x,y!-plane of the reference frame ~Fig. 5! where these external torques form a simplex of vectors. Therefore one has that t5@tx ,ty,0#T and that due to the structure of the system 608 \u00d5 Vol. 122, DECEMBER 2000 rom: http://dynamicsystems.asmedigitalcollection.asme.org/ on 08/15/201 Ftx ty G5F 1 2 1 2 2 1 2 0 ) 2 2 ) 2 G dFF1 F2 F3 G (35) Deviating one motor at a time, three vectors are obtained t15Ft1x t1y G t25Ft2x t2y G t35Ft3x t3y G (36) which form a simplex in R2 containing the origin of R2 that is ( i51 3 m it i50, ( i51 3 m i51, m i" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002432_bf01574849-Figure7-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002432_bf01574849-Figure7-1.png", + "caption": "Figo 7~ A sinusoidally distributed winding in a non sinusoidal flux density distribution", + "texts": [ + " The fundamental component of the induced rotor voltage as a function of the stator tic-current 10\" 4 Airgap flux linkage of a winding in a saturated machine I t must be noted that any theory incorporating saturation of the airgap flux must allow for the presence of harmonics in the air gap flux density wave. Saturation in the main flux iron paths is invariably a source of harmonics and will manifest itself by a flattening of the flux density wave in the airgap. For reasons already accounted for, a sinusoidal winding distribution will be assumed. Figure 7 shows such a winding in a machine, where the flux density wave B(O) is distorted due to saturation. B(O) is produced by all the sinusoidal current sheets giving a resultant ram/dis t r ibut ion F(O) with its crest at an angle ~ in the machine and this angle also defines the position of the fundamental B1 in B(O). I t is important to note that F(O) represent the gross ram/created by the currents in the machine. I t is not the mini across the airgap. The air gap ram/has obviously the same characteristic flat topped shape as the flux density wave as there is a linear relation between ram/and flux density in the airg~p", + " (5) with all lc: except ]~1 put equal to zero. The axis of the winding is at an angle fl and thus, the flux T linked with the winding is where the effective number of turns N with respect to the fundamental flux density component B1 has been defined as ~- ]~1\" ~w (7) Contrary to the linear theory, the amplitude B~ of the fundamental component in the flux density distribution is now a non linear function of the resultant airgap current i. Two identical sinusoidally distributed windings, with their magnetic axes d and q in Fig. 7 located at fl ~ 0 and fi = ~/2, respectively, are defined. The currents i d and iq in these coils are giving a resultant airgap current i and it follows i s = i - c o s ~ (9) iq = i . sin ~ (10) Further on, to shorten notations we introduce ~,n = 2 \" N \" l \" r \" B~ (11) By putt ing fl----0 air. fl = ~/2 in Eq. (6) and using Eqs. (9)--(11) the flux linkages Te and ~Pq are written as = ie 02 ) i Tq = --gZm iq (13) i I t is already obvious from these two equations that if the currents i s and iq both exist, a change of the current in one coil will not only produce a change of flux linkage in the s~me coil but in the other coil as well" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000585_rob.4620121004-Figure2-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000585_rob.4620121004-Figure2-1.png", + "caption": "Figure 2. The 6-4 and (5-5),, general-geometry fully parallel manipulators", + "texts": [ + " ~ Moreover, the analytical-form DPA solution for the planar 6-6 Stewart platform has disclosed5 the same number of configurations as for the 6-6 Stewart platform. Unfortunately, procedures apt to solve planar arrangements cannot be extended to their general-geometry counterparts. By taking the total number of connection points on base and platform as a rough measure of the structural complexity of a fully parallel arrangement, it can be asserted that all general-geometry fully parallel manipulators whose DPAs have until now been accomplished analytically are of complexity not greater than ten. Precisely, they are the 6-4 arrangement (see Fig. 2(a)), and one out of the two existing 5-5 arrangements (see Fig. 2(b)). The analytical-form DPA of the general-geometry 6-4 fully parallel manpulator6 has provided 32 assembly configurations, as many as those obtained with the planarity assumption.' On the other hand, the analytical-form DPA of the solved 5-5 general-geometry manipulator, here labeled as (5-5)* , has provided 40 assembly configurations.8 It is worthwhile to consider that synthetic methods do not always succeed in foreseeing the number of assembly configura- Innocenti: Analytical-Form Direct Kinematics 663 tions, as happened in the case of the 6-4 arrangement" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001255_20.767379-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001255_20.767379-Figure1-1.png", + "caption": "Fig. 1. 3D view of an electromagnetic valve used in ABS systems", + "texts": [ + " This in turn increases the memory consumption and the CPU time required for the the system of equations. Another problem has when using the BEM-FEM coupling. If the and in consequence their boundary elements of separate subdomains approach each other closely during the motion so called nearly singular integrals occur in the BEM. These integrals have to be evaluated using special techniques 1151, [16]. VI. NUMERICAL EXAMPLE To compare both numerical approaches for the analysis of electromechanical devices an axisymmetric electromagnetic valve used in ABS systems has been investigated, Fig. 1, Fig. 2. The yoke (St4U), the armature and the armature counterpart (both Vacoflux50) consist of different conducting and magnetic nonlinear media (QSUU = 1.818 * 106S/m, qacoflux = 7.693.106S/m), Fig. 3. The stroke of the armature is 0.3 mm. The driving coil of the valve has of 546 turns. The prescribed current in the driving coil is depicted in Fig. 4. The constant hydraulic load acting on the valve needle is 15N. The bias force of the spring is 5.93N. Effects due to friction have not been taken into account" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002999_095965180121500411-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002999_095965180121500411-Figure1-1.png", + "caption": "Fig. 1 Elastic system of (k+1) rigid bodies", + "texts": [ + " m1 =m2 = \u00b7\u00b7 \u00b7=m k , d01 =d02 = \u00b7 \u00b7 \u00b7=d 0k and k01 =k02 = \u00b7 \u00b7 \u00b7=k 0k , where m i (i=1, 2, .. . , k) denotes the mass of the ith local rigid where m i (i=1, 2, . .. , k) denotes the mass of the ith local body at the contact between the ith robot end eVector rigid body at the contact between the ith robot end and the object and d 0i (i=1, 2, .. . , k) and k 0i eVector and the object (note that only the manipulated (i=1, 2, .. . , k) denote the damping and stiVness object with mass m0 is connected to each of the k local coeYcients respectively (Fig. 1). rigid bodies with masses m i (i=1, 2, . . . , k); i.e. the k local rigid bodies with masses m i (i=1, 2, . .. , k) are not Assumption 2 mutually interconnected (Fig. 1). Two states of dynamic system of (k+1) bodies may be de ned: the unloadedThe constraint environment is xed (it is not externally state, i.e. the state when no force system acts on thedriven), i.e. the constraint equations do not depend on system of (k+1) elastically connected rigid bodies, andtime t explicitly. the loaded state, i.e. the state when a system of forces acts on the dynamic system of (k+1) bodies, and theyAssumption 3 are moving from the unloaded state. Using the Lagrange Each robotic mechanism is non-redundant and all equations of motion, the mathematical model of multiple coordinated robots have the same number of joints", + " The non-redundant rigid robot manipulators performing robots can have diVerent types of joint. cooperative work on a single dynamical object of which the motion is constrained by a dynamical environmentThe manipulated object has k connections with the has been derived in reference [10]. Here, the equationsmultiple manipulators and one connection with the condescribing the multiple compliant manipulation modelstraint environment. The manipulated object together are brie y reported:with (k+1) connections can be represented by a dynamic system of (k+1) rigid bodies (Fig. 1); i.e. each N(q)q\u0308+n(q, q\u00c7 , Y0a, Y\u00c70a)=\u00f4\u00d7R6k\u00d61 (2)of the k connections between the manipulated object and the k manipulators can be represented as a local rigid Y\u0308ca =J\u00c1 (q)q\u00c7 +J (q)q\u0308 (3)body and, in the mass centre of each local rigid body contact, gravitational, damping and elasticity forces act W0a(Y0a )Y\u03080a +w0a (q, q\u00c7 , Y0a, Y\u00c70a )=Fe (4)as external forces. Clearly, at the contact between the environment and the object there is a contact force Fe W ca(Yca )Y\u0308ca +w ca(q, q\u00c7 , Yca , Y\u00c7 ca )=Fc (5)which acts on the manipulated object" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003261_2001-gt-0255-Figure3-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003261_2001-gt-0255-Figure3-1.png", + "caption": "Figure 3: Contact Geometry", + "texts": [ + " Note that rp, the radial co-ordinate of the nominal center of curvature of the inner race, is not related to Dp. For the j\u2019th ball, the displacement of the inner race center of curvature, u, is related to the displacement of the inner race, d (treated as a known input quantity for the forcing function) according to: { } [ ] [ ]{ }TT r Z x yu u u x y z \u0398= = = \u03b8 \u03b8u R d R (9) where, [ ] j j p j p j p j p j j j 3 5 cos sin 0 z sin z cos 0 0 1 r sin r cos 0 0 0 sin cos \u00d7 \u03c6 \u03c6 \u2212 \u03c6 \u03c6 = \u03c6 \u2212 \u03c6 \u2212 \u03c6 \u03c6 R (10) All surfaces are assumed to be curves with only one radius of curvature. Figure 3a shows the geometry of the centers of curvatures of the inner race, of the outer race and of the center of the rolling element, before and after the known deflection of the inner race u. Hence: b 0i i i Dr c 2 = \u2212 \u2212l , b 0e e e Dr c 2 = \u2212 \u2212l (11) ( ) ( ) 0i 0e 0 z 0i 0e 0 r sin u atan cos u b + \u03b1 + \u03b1 = = + \u03b1 + l l l l (12) and 2 2 i e a b+ = +l l . (13) Copyright \u00a9 2001 by ASME s of Use: http://www.asme.org/about-asme/terms-of-use Down Comparing li + le with the nominal distance and the nominal clearance yields the overall contact deformation: i e 0i 0e i ec c\u03b4 = + \u2212 \u2212l l - l - l (14) If the contact deformation is positive, then the contact load is calculated via the standard Hertzian contact relationship, otherwise contact is lost and no load is transmitted, so that: ( ) ( ) 1", + " Five degrees of freedom model with inertia (5DOF+i) Inclusion of rolling element centrifugal load means that inner and outer race contacts may not occur with the same contact angle. The equations of dynamic equilibrium become: 2 re c p r i i e e z i i e e m Dg 0Q cos Q cos 2g 0 Q sin Q sin \u03c9 \u03b1 \u2212 \u03b1 + = = \u03b1 \u2212 \u03b1 (18) 4 loaded From: http://proceedings.asmedigitalcollection.asme.org/ on 02/03/2016 Term where n i pi i iQ K (for 0)= \u03b4 \u03b4 > (19) and n e pe e eQ K (for 0)= \u03b4 \u03b4 > (20) From Fig. 3b one has the geometric relationships: 0i 0 r r 0i 0 z z i i i cos u v sin u v cos sin \u03b1 + \u2212 \u03b1 + \u2212= = \u03b1 \u03b1 l ll , (21) 0e 0 r 0e 0 z e e e cos v sin v cos sin \u03b1 + \u03b1 += = \u03b1 \u03b1 l ll (22) and i i 0i ic\u03b4 = \u2212 \u2212l l , e e 0e ec\u03b4 = \u2212 \u2212l l . (23) For a given inner ring position (u) one now has two nonlinear equations for the two unknown rolling element positions vr and vZ. These may again be solved by the Newton-Ralphson technique extended to two unknowns, i.e. 1 r r r zr r r z z zz zi 1 i i r z i g g v vv v g v v gg g v v \u2212 + \u2202 \u2202 \u2202 \u2202 = = \u2212 \u2202 \u2202 \u2202 \u2202 v (24) Ref" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003054_6.2002-533-Figure4-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003054_6.2002-533-Figure4-1.png", + "caption": "Figure 4: Plastic Top Housing", + "texts": [], + "surrounding_texts": [ + "work for cases where the flows of interest are already well understood, but they are not reliable for complex flows such as 3D and/or unsteady cases, cases with rough or curved walls, flows with injection or suction, mixing with foreign fluid injection or high-speed flows especially those with impinging shock waves, high enthalpies and/or combustion. A good discussion of such indirect methods is in Nitsche et. al [2].\nThe purpose of this investigation was to develop a non-intrusive, direct skin friction-sensing device for high vibration and acceleration loads without compromising the resolution of the data acquired. This new concept was needed to alleviate problems encountered by existing sensors under such conditions. To meet this goal, various damping techniques had to be studied and applied in the design of the skin friction gage.\nAfter the analysis of several damping methods, it was decided to employ visco-elastic damping through the use of a rubber RTV sheet on the gage surface (See Fig. 1). Viscous damping through the use of oil inside the gage housing would provide further damping. Previous skin friction gages at Virginia Tech [3] used viscous damping due to several inherent advantages. By filling the volume around the sensing head and flexure, pressure gradient effects are minimized. The fluid also acts as a temperature stabilizer, minimizing the temperature gradients that the gage is exposed to during the flight test. However, gages that previously employed this method of damping deteriorated with time due to fluid loss through the small gap around the sensing element.\nA proposed solution to this problem was to fill the gap with a substance that would check this flow. Previous studies had tried using silicone rubber in the entire inner volume of the gage, but loss in gage resolution due to increased resistance forces limited further development [4].\nA study showed that a rubber RTV compound used to cover the sensing element and gap with a thin sheet as illustrated in Figure 1 could be a possible solution [5]. The apparent advantage of this method was that fluid could be retained in the gage cavity, thus keeping the inherent benefits while eliminating the maintenance requirements. However, several aspects needed further investigation, including, but not limited to, RTV adhesion characteristics, trapped air in the fluid volume and the scaling of the design to meet sensitivity requirements of this work.\nThis study was performed to help the successful integration of a rubber RTV compound sheet into a skin friction gage design capable of meeting NASA flight test guidelines of vibration conditions up to 8.0 g^ over a 15 - 2,000 Hz frequency range, requiring performance at altitudes ranging from 15,000 to 45,000\nfeet, Mach numbers ranging from 0.6 to 0.99 resulting in shear values of 0.3 to 1.5 psf. Flow Direction Rubber RTV\nIn order to meet the vibration requirements, a primarily viscoelastic damping method was employed on a non-nulling gage design. RTV 566, a rubber compound obtained from GE Silicones [6] was applied on the gage surface to contain the fluid. Several benefits were realized by the use of this compound. In addition to containing the fluid, it provided sufficient damping itself, minimizing the need for the use of a fluid in the volume cavity. In some cases, glycerin was used to fill the inner volume of the gage :to provide additional viscous damping.\nThe use of the rubber compound presented new challenges as well. First, the RTV did not bond well to materials such as aluminum despite the use of a metal primer. It was suspected that an oxide layer formed on the aluminum surface preventing adhesion. Often the test fixtures on flight test vehicles are made of aluminum. However, the RTV compound successfully bonded to polyethersulfone (PES), a high temperature plastic so that could be used for the gage housing. Second, the RTV sheet reduced gage sensitivity, because the rubber sheet bears part of the load due to shear on the head. This sensitivity reduction was greatly aggravated by any accumulation of the compound in the gap between the sensing head and housing, requiring the modification of the earlier procedure used in the application of the rubber RTV compound onto the gage surface.\nAmerican Institute of Aeronautics and Astronautics", + "SKIN FRICTION GAGE\nThe skin friction gage used in this study consisted of six components: 1) sensing head, 2) flexure, 3) base, 4) upper housing, 5) lower housing and 6) connector. See the schematic in Figure 1.\nSensing Head The sensing head was made of polyethersulfone (PES). This reduced mass compared to a metal head and aided in bonding the rubber sheet. It had a head diameter of 0.75 inches, chosen in conjunction with the flexure described below to meet the design shear levels. The primary feature of the component was its form as a truncated cone with tapered edges from the bottom, a quality that reduced the entrapment of air bubbles during the filling of the volume cavity. A 0.25 inch threaded hole allowed for the fastening of the component to the flexure. See Figure 2.\nFlexure and Base The flexure, made of aluminum, had a shaft of length 1.538 inches and a diameter of 0.150 inches using beam theory. Aluminum was chosen because of its low density, which reduced vibration effects. The high heat conductivity of aluminum would act as a heat sink for the semi-conductor strain gages, which are temperature sensitive. To account for temperature effects on the strain gages, a type K thermocouple was buried into an opening created at the center of the flexure.\nMicron Instruments installed a matched set of eight SS-060-033-500P-S(4) semi-conductor strain gages on the base of the flexure. These allowed for the data acquisition on two perpendicular axes, a trait necessary in measuring skin friction in a 3D-flow regime. An illustration of the flexure and base can be seen in Figure 2.\nA primary goal in the gage design was robustness, which is required for flight application. To achieve this, protection of the strain gages was necessary. The first step was the introduction of a 10-pin connector that isolated the gage wiring from the data acquisition equipment. Next, solder pads located on the base were used to isolate the delicate strain gage wires from the wire leads from the connector. Further, the wires were looped around to provide strain relief and anchored to the base via epoxy. The base had two fill holes that allowed the filling of the inner cavity with a viscous liquid.\nGage Housings A two-piece housing was used in the skin friction gage design. The bottom piece was constructed of aluminum and housed the base, flexure and sensing element. To improve the design robustness, a deflection constraint was incorporated into the top surface of the aluminum housing. PES was the material of choice for the top piece of the housing. This allowed for easier application of the rubber compound onto the gage surface. Further detail on the gage housings can be, seen in Figures 3 and 4. Gage assembly is depicted in Figure 5.\nAmerican Institute of Aeronautics and Astronautics", + "Connector A standard 10-pin plug connector, supplied by Spacecraft Components Corporation [7], was used to interface the skin friction gage with the data acquisition equipment. Each bridge circuit used four pins from the connector. The remaining two pins, made of chromel and alumel were used for the thermocouple buried in the flexure. A matching socket was used to route the wires to the data acquisition equipment. A second thermocouple was buried in the rubber sheet. The connector was secured to the base of the flexure base as demonstrated by Figure 5.\nGage Electronics In a full-bridge circuit, all four strain gages in the Wheatstone bridge are placed on the base of the cantilever beam. In each axis, a strain gage pair with the same vertical alignment is placed in the tensile region while a similarly aligned pair is placed in the compressive region 180\u00b0 from the first pair. For a dualaxis scheme, a similar strain gage layout is done perpendicular to the primary axis. Figure 2 displays a strain gage pair attached to the base of the aluminum flexure.\nGAGE PREPARATION\nA major problem encountered in previous studies with the use of the rubber RT V was its adhesion to the gage surface during testing due to the aggressive nature of the supersonic wind tunnel \"starting\"[5]. It was determined that bonding could be improved by creating grooves on the bonding surface and machining down the outer lip of the housing allowing the rubber compound to wrap around the outer edge.\nA new problem encountered in this study was the unwanted filling of the gap between the sensing head and the plastic housing with the rubber RTV mixture causing loss in gage sensitivity. The previous application procedure required the inverted gage surface to be placed onto the liquid RTV mixture without any support of the gage while the RTV set.\nThis allowed the weight of the assembly to squeeze the RTV mixture into the gap leaving a thin layer of the RTV on the actual bonding surface and a ring of RTV in the gap. This was only a minor problem with the earlier gages, which were small and made entirely of plastic. The gages designed for the present application were larger and heavier, thus greatly amplifying the problem.\nTo counter this problem, a gage adapter was designed. It had an inner cavity into which the gage assembly could be inserted and fastened. The adapter had a shank designed for insertion into a milling machine. This setup allowed the gage to be held at a desired height above the surface upon which the liquid RTV mixture was spread, and this reduced gap filling with rubber RTV resulting in increased gage sensitivity.\nA known characteristic of rubber compounds is the change of modulus of elasticity with temperature. For wind tunnel testing, the exposure time to the flow is on the order of seconds, negating this effect. However, for flight-testing, the testing time varies from 30 minutes to several hours at varying altitudes making the temperature effect a major concern. To account for this effect, a thermocouple was buried into the silicone RTV sheet over the gap between the sensing head and the housing. This enabled the calibration of the unit at different flow temperatures, emulating the final test environment the gage would experience.\nCALIBRATION\nIn this study, two calibration methods were used: a) Point loading b) Distributed Shear\nPoint Loading Method A known weight is placed parallel to the direction of flow and perpendicular to the sensing element. This is usually achieved by hanging a paper cone by thread attached to the floating element with clear tape as illustrated in Figure 6a. A tare value of the cone and string is taken to zero the balance. Different weights ranging from 50 milligrams to 5 grams are placed in the cone while the corresponding output is recorded. The gage is then rotated 180 degrees and the procedure repeated. The whole process is repeated on the crossstream axis. This is the preferred method of calibration for oil-filled gages due to its simplicity. It is also a highly accurate method due to the negligible shear stress contribution of the oil in the gage cavity\nThe mass calibration is then related to shear through:\nCD where K is the conversion factor from mass to force\nunits, M is the mass of the calibration weight and A is\nAmerican Institute of Aeronautics and Astronautics" + ] + }, + { + "image_filename": "designv11_65_0000768_0957-4158(95)00074-7-Figure3-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000768_0957-4158(95)00074-7-Figure3-1.png", + "caption": "Fig. 3. Geometrical relationship between piston, cylinder and swash plate.", + "texts": [ + " The need for accurate calculation increases with high operating pressures. The viscous friction force ~p is obtained by integrating shear stress on the cylinder caused by the shearing of the oil film. The mathematical expression for the viscous friction [1] is given by s;{ [ ~p = r [ (Ps - p r ) h l ( h l + t) - 60ulc/] + 4flu In 1 + dO, (10) [2hl + t] where r/ is the viscosity of oil and lc/ is the length of the piston inside the cylinder bore at any angular position (/7) of the cylinder. When /3 -- 0 \u00b0 (Fig. 3), the length of the piston inside the bore is maximum (/c/max) and when /3 = 180 \u00b0, the length of the piston inside the bore is minimum (Ic/min). The difference between /c/max and lc/mi n corresponds to the piston stroke length indicated in Table 1. /C/max of 20 mm is used in the computations. Similarly, for the leakage flow q an expression can be derived based on pressure distribution and considering the equilibrium of forces on a free flow particle in the 1. Geometric displacement 2. Swash plate angle 3", + " 4, O is the centre of the bore, O1 is the centre of the three lobe piston when one of the three lobes is in contact with the bore and the piston is in the non-tilted state and 02 is the centre of the piston when two lobes are in contact with the bore as shown. Then, el is the eccentricity of the piston inside the bore, which is equal to the clearance, and e2 is the maximum possible eccentricity, which is greater than el. 2.3. Mean torque and eff iciency In one revolution of the motor shaft, the piston moves forward converting pressure into mechanical work. During the return stroke no work is performed. From the geometry (Fig. 3), the piston velocity u can be deduced as follows: u = totan o~gp sin (tot1). (12) The side load on the piston, F z, due to the eccentric nature of the drive, which causes the piston to run at a certain eccentricity, is given by Fz = A ( P s - Pr) tan 0:, (13) where A is the piston area, (Ps - Pr) is the differential pressure and o~ is the swash plate angle. The torque T produced by one piston is given by T = [A(Ps - Pr) - ~P] Rp tan t~ sin 0. (14) The total torque T~ produced by all pistons can be obtained by summing up the torque produced by individual pistons with a phase of 2~r/z shifted from each other [2]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003227_iros.1995.525807-Figure5-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003227_iros.1995.525807-Figure5-1.png", + "caption": "Figure 5: Genot,ype with variable length and crossover", + "texts": [ + " For an exa,mple, in case that, a genotype is [3.4,1]. sequence of subgoals is sta.Tt -+ 116 --f p 5 -+ yon7 in Fig.2-(b), therefore, its phenot,ype becomes to he [p6,p5,goal]. But it is a case of simple environment. our environments are not so simple usually. &o, we don't know if t.he genotype is long enough t o arrive a t goal after starting from start point, or not). So, we caii not decide t,he length of geiiot>ype at, initrial state. In order to solve the problem, we decide that, the genotype is a ring type with variable lengt,li(Fig.5, and it, can be applied repeatedly in order to arrive at, goal. In other word, the genotype is repeatrd until ineeting goal or re-entering the same sub-goal, aiid at next generation the genotype becomes to he long enough to be able to express the path sequence as only one turning. So that, the length of genotype is always regulated that the sequeiice from start point to goal caii be represent,ed as only one repetitlion. Our genetic algoritliin is composed of three operat'ors(Fig.7): (1) ca1culat)ion of Fit>iiess, (2) selection, (3) reproduction", + " In our select,ion scheme, the population numbers are dividecl into subgroups, and numbers with the best fitness among the subgroups are selected for reproduction. The size of subgroups is four. Aftm finishing the above selection stage, we go to reprodiictioii stage aiid operate crossover. First, the newly sdected genotypes are paired together at random. Second, two integer positions \"11\" and \"ni\" along every pair of genotypes is selected uniformly at random. Finally, based on a probability of crossover, the paired strings are exchanged over the range between position \"n\" and \"m\" as Fig.5. Although the crossover operation is a randomized event, when tombined with selection, it becomes an effective means of exchanging information and conibiniiig portions of good quality solutions. And then, we operate mutation such as Fig.6. The mutation is siitiply an occasional random change of a genotype position, based on probability of mutation. Our mutation changes gene of a position to random nuniber. This niutation operator helps in svoiding the possibility of mistaking a local minimum for a global minimumi" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003455_jmes_jour_1980_022_045_02-Figure2-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003455_jmes_jour_1980_022_045_02-Figure2-1.png", + "caption": "Fig. 2. A prism which is obtained from folding a sheet", + "texts": [ + " 1). By assuming that the sheet of metal is rigid-perfectly plastic, then the plastic work done in folding, i.e., the energy dissipated plastically in the folding line, is w= MpBL (1) where M p = aot2/4 is the plastic bending moment per unit length of the sheet of metal, uo the yield stress of the material, t the thickness of the sheet, /3 the angle of fold and L the length of the folding line. 2.2 Folding up into a prism Now consider an initially flat sheet of metal folded up into a prism (see Fig. 2(a)). Assume that its section is an arbitrary convex polygon of n sides and that the seam (or join) of the sheet is on one of the side surfaces of the prism, then the plastic work done in folding is w=Mp(pl+pZ+. . .+pn)L=2TMpL (2) where L is the length of the prism (see Fig. 2(b)). Journal Mechanical Engineering Science 6 IMechE 1980 0022-2542/80/ 1 0 W 2 3 3 $02.00 Vol22 No 5 1980 at University of Bath - The Library on June 5, 2016jms.sagepub.comDownloaded from 234 W. JOHNSON AND T. X. YU The expression W= 2nMpL, involving It+ co, applies particularly to the case of folding a flat sheet to a circular cylinder, and also to the folding of a flat sheet to take up the form of any arbitrarily convex closed cylinder (see Fig. 3). 2.3 Folding up into a frustum Assuming that a right folded frustum has a vertex angle 201, its base being a regular polygon, as shown in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000726_cbo9780511530173.007-Figure5.2-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000726_cbo9780511530173.007-Figure5.2-1.png", + "caption": "Figure 5.2 Motion of the platform with respect to a reference frame in the fixed platform.", + "texts": [ + " Assume that the moving platform is in equilibrium with an externally applied force with coordinates w and magnitude/is applied to it on a line $. Then w = f\\s\\ + fih + hh> (eq. 2-60) where/i,/2, and/3 are the magnitudes of the resultant forces in the connectors and \u00a31, \u00a32, and \u00a33 are the line coordinates of the connectors. A small change 8w in the applied force will cause the upper platform to move with an infinitesimal rotation with coordinates 8D on an axis perpendicular to the page through a point G, as illustrated by Figure 5.2. These quantities are related by a 3 X 3 stiffness matrix [K] which we will determine next. Assume that the free lengths of the springs toi and the stiffness constants ki are known, together with the coordinates w of the applied force and the coordinates s\\, $2, and \u00a33 of the lines $1, $2, and $3. Assume also that the movable lamina is initially loaded and the spring lengths are \u20ac1, \u20ac2, and \u20ac3. Then substitute the relationships/ = kt (\u00a3t - \u00a3oi) (i = 1, 2, 3) into (2.60), where kt are the spring constants and (\u20ac,- \u2014 (,Oi) is the difference between the current and free length of an ith spring, to yield 153 Cambridge Books Online \u00a9 Cambridge University Press, 2009Core terms of use, available at https://www.cambridge.org/core/terms. https://doi.org/10.1017/CBO9780511530173.007 Downloaded from https://www.cambridge.org/core. Columbia University - Law Library, on 13 Aug 2019 at 15:09:03, subject to the Cambridge The total derivative of (5.1) can be expressed in the form ds 1 /\u0302vo p \\__J_\u00a3fi _i_ h. (p D \\ *-<\"Jz of\\ dO\\ 2 \u00b02 J0 2 where pt = (Oi^i and dstlddi = siB are the coordinates of a line $;# perpendicular to $i that passes through a fixed pivot 2?,- (see Fig. 5.2 and equation (4.56)). Equation 5.2 can now be expressed in matrix form as Sw = [si s2 S3] [k] S\u20ac2 \u20ac1861 \u20ac2S02 J38d3_ (5.3) Cambridge Books Online \u00a9 Cambridge University Press, 2009Core terms of use, available at https://www.cambridge.org/core/terms. https://doi.org/10.1017/CBO9780511530173.007 Downloaded from https://www.cambridge.org/core. Columbia University - Law Library, on 13 Aug 2019 at 15:09:03, subject to the Cambridge where [k] and [k(l \u2014 p)] are the (3 X 3) diagonal matrices [*] = Jfci 0 0 0 k2 0 ", + " The following notation was employed in Chapters 2 and 4: 7 = [h h s3] = Ci C 2 C _P\\ Pi P3_ is the matrix of the coordinates of the lines $i, $2, and $3, and [B] = [$IB S2B \u2014 S\\ ~S2 ~S3 c\\ c2 c3 Cambridge Books Online \u00a9 Cambridge University Press, 2009Core terms of use, available at https://www.cambridge.org/core/terms. https://doi.org/10.1017/CBO9780511530173.007 Downloaded from https://www.cambridge.org/core. Columbia University - Law Library, on 13 Aug 2019 at 15:09:03, subject to the Cambridge is the matrix of the coordinates of the lines $1B, $2B, and $3B (see Fig. 5.2). Also, from (4.48) and (4.50): = jT8D and \u20ac1861 = [C]T8D, where C\\ -<1\\C -s2 qic S3~ C3 q3C_ [C] = [sic he hc\\ = is the matrix of the coordinates of the lines $iC, $20 and $3C (see Fig. 5.2). Substituting these expressions, (5.3) can be expressed in the abbreviated form 8w = {j [k] j T + [B] [k (1 - p)] [C]T}SD, and, finally, (5.4) can be expressed as 8w = [K] 8 D, where the required stiffness matrix [K] is given by [K]=j[kyT + [B][k(l - p)][C]T. (5.4) (5.5) (5.6) Clearly, [K] is symmetrical only at the unloaded position for which pt = 1 or tt = \u20acoi and [K\\=j[k]f. It is easy to deduce from (3.33) that under the action of the Euclidean group, a twist 8D\" expressed in a new coordinate system is related to the same twist quantified in the original coordinate system by 8D = [E] 8D\"" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000477_0921-8890(95)00032-b-Figure4-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000477_0921-8890(95)00032-b-Figure4-1.png", + "caption": "Fig. 4. Cost function to reach a tentative or confirmed feature.", + "texts": [ + " Costs In this approach costs are defined to be the expected increase of the configuration uncertainty caused by the desired manoeuvres of the subtasks. As a first step, the cost of a manoeuvre only depends on the length of the driven path. However, there is no limitation on the use of more sophisticated approaches taking into account effects like rotational costs. For all (1 . . . n) confirmed and all (1 . . . m) tentative features the estimated costs Cc(\u00a2) and Cf(\u00a2) depend upon the length of the path d(\u00a2) needed to manoeuvre the robot into the visibility region of the feature (see Fig. 4). { d(4Q if d(~b) n+1 = (vn + (3)/(l + vn(3), which is greater than vn. This vn+\\ must also be the speed with which it leaves the wall at the beginning of trip n + 1, and thus also its speed relative to the Lorentz coordinates during trip n + 1. Thus, on each successive trip, the rightward geodesic is boosted to a higher speed. Finally, after an infinite number of trips through the wall but a finite amount of proper time, it is moving at the speed of light and has asymptoted to the chronology horizon. (This and most other statements made in this paper can be verified fairly easily by elementary, special relativistic calculations.) The seemingly pathological rightward geodesies are not a set of measure zero. Rather, as should become evident below, half the geodesies in the chronal region are of the rightward type and half of the leftward type. What can possibly happen to a rightward family of observers, who move along the rightward geodesies, after they asymptote to the chronology horizon? They have lived only a finite amount of proper time. They have not encountered any spacetime curvature and thus cannot have been killed by infinite tidal forces. So where do they go? As Misner showed in his seminal lecture [Mis67], they pass through a chronology horizon of their own (distinct and different from that of the leftward observers), and Cambridge Books Online \u00a9 Cambridge University Press, 2010https://www.cambridge.org/core/terms. https://doi.org/10.1017/CB 9780511628863.029 Downloaded from https://www.cambridge.org/core. University of Warwick, on 18 Sep 2018 at 06:12:08, subject to the Cambridge Core terms of use, available at into a non-chronal region of their own (distinct and different from the leftward one). This pathological behavior can be understood using the covering space of Misner space [HaE73]; Figure 2. This covering space is constructed by lining up a sequence of copies of Misner space, side by side, each one boosted by speed (3 relative to the last one. The copies of the (physically irrelevant) wall are labeled \\VQ, W\\, W2, etc. in Figure 2; and each Misner space is labeled \"copy 1\", \"copy 2\", etc. A representative event P in Misner space is shown in each of the copies. There is actually an infinite number of copies of Misner space and of the wall and of the point P , with the highorder copies asymptoting to the rightward chronology horizon. The typical leftward geodesic L and typical rightward geodesic R are shown in the covering space, along with the two distinct chronology horizons through which they pass. From this covering space it should be clear that there is a complete symmetry between the leftward observers and the rightward ones. Just as the leftward observers, as they near their leftward chronology horizon, see the rightward observers circle around and around Misner space an infinite number of times, approaching the speed of light, so also the rightward observers, as they near their rightward chronology horizon, must see the leftward ones circle infinitely and approach the speed of light. Cambridge Books Online \u00a9 Cambridge University Press, 2010https://www.cambridge.org/core/terms. https://doi.org/10.1017/CB 9780511628863.029 Downloaded from https://www.cambridge.org/core. University of Warwick, on 18 Sep 2018 at 06:12:08, subject to the Cambridge Core terms of use, available at Intriguingly, there is a third family of timelike geodesies and observers that are intermediate between the leftward and rightward ones. This family consists of geodesies that hit the covering space's origin (the event Q in Fig. 2); an example is the geodesic / . Note that in the covering space there is only one copy of the event Q, whereas there is an infinite number of copies of every other event, e.g. P. This presumably is related to the following remarkable pathology. Although the event Q exists in the covering space, it does not exist in Misner space. There is no way to include it in the spacetime, if one insists that the spacetime be a manifold and one includes the chronal region, and both the left and right chronology horizons, and both the left and right non-chronal regions. Those regions cannot be meshed smoothly with Q; and the impossibility of meshing makes Misner space geodesically incomplete: the intermediate geodesies (e.g. / ) all terminate after finite proper time just before reaching the non-existent event Q. As pathological as this may seem, it would have been much more pathological if the terminating geodesies were not a set of measure zero. The intermediate geodesies can be used as the time lines for a coordinate system that treats leftward and rightward geodesies on an equal footing. This coordinate system (T, X, y, z) is related to the Lorentz coordinates (\u00a3, x,y, z) of the covering space (Figure 2) by x = TsinhX; (1) and correspondingly, the metric in this coordinate system is ds2 = -dT2 + T2dX2 + dy2 + dz2 . (2) The boost-related points that are identified to produce Misner space (e.g. the points P in Figure 2) all are on the same hyperboloid i2 - x2 = constant, and therefore are all at the same T, y, z, but different X. The boost that takes one of the P's into the next one is simply a displacement in X by tanh\"1 (3. Therefore, the n'th copy of P is at Xn = XQ + n tanh\"1 /?, and Misner space can be regarded as the space of Equation (2) with X periodic with period tanh\"1 /?. As seen by the intermediate observers, who sit at fixed (X, y, z), the leftward observers circle leftward around Misner space an infinite number of times as they approach their chronology horizon, and similarly for the rightward observers. The intermediate observers never see either family, leftward or rightward, reach its chronology horizon, because, as they watch and wait, they all come crashing together and cease to exist just before non-event Q, i.e. just before the \"moment\" T = 0 of the two chronology horizons. 3. The Classical and Quantum Instability of Misner Space The infinite relative circling of leftward and rightward geodesies produces severe instabilities in Misner space. Anything (dust particles, electromagnetic waves, gravitons, Cambridge Books Online \u00a9 Cambridge University Press, 2010https://www.cambridge.org/core/terms. https://doi.org/10.1017/CB 9780511628863.029 Downloaded from https://www.cambridge.org/core. University of Warwick, on 18 Sep 2018 at 06:12:08, subject to the Cambridge Core terms of use, available at etc.) moving in the rightward manner (e.g. along the geodesic R) will become infinitely energetic as seen by the leftward observers, as they approach their leftward chronology horizon; and the resulting infinite energy density presumably must act back on the spacetime, via Einstein's equations, to change it radically. Change it how? Nobody knows for sure, but most likely to prevent the creation of CTCs at the leftward horizon, i.e. to enforce chronology protection (a phrase coined by Hawking [Haw92]). Similarly, anything moving in the leftward manner will become infinitely energetic on the rightward horizon, and probably act back to prevent the creation of CTCs there. One might think it possible to stabilize the spacetime along one of the chronology horizons, e.g. the leftward one, by ensuring that nothing (no particles, no waves, no gravitons, ...) in Misner space moves in the opposite direction (rightward). However, there is one sort of thing that cannot be so controlled: Vacuum fluctuations of quantum fields. So long as the \"universe\" (Misner space) is truly closed up in the x-direction and truly shrinking in size, there is no way to prevent vacuum fluctuations of very short wavelength from traveling rightward and thereby piling up on themselves at the leftward chronology horizon. The pileup makes each rightward mode of any quantum field appear more than once at the same location in spacetime near the leftward chronology horizon. The shorter the mode's wavelength, the nearer the horizon it must approach for this to happen, but even a mode of arbitrarily short wavelength will pile up on itself an arbitrarily large number of times when it approaches arbitrarily close to the horizon. When this pileup begins, the mode's half quantum also piles up on itself, thereby endowing the mode with more than a single half quantum; and, as a result, when one renormalizes the mode's energy one winds up with a finite amount rather than zero. This finite energy in each rightward mode gives rise to a diverging renormalized energy density as one approaches the leftward chronology horizon\u2014as Bill Hiscock and Deborah Konkowski showed by a rigorous \"point-splitting\" calculation, when they were in Misner's research group in the early 1980s [HiK82]. Thus, although Misner space is a solution to Einstein's classical vacuum field equations, it cannot be a solution to the equations of semi-classical gravity in which spacetime is treated as classical and all matter fields are quantized; in semiclassical gravity the quantum fields probably distort the spacetime geometry away from that of Misner, as one approaches the chronology horizon, and thereby probably enforce chronology protection." + ] + }, + { + "image_filename": "designv11_65_0003483_imece2004-59492-Figure3-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003483_imece2004-59492-Figure3-1.png", + "caption": "Fig. 3: Extremes values of and \u03c6.", + "texts": [ + "s (19) only the following solution is acceptable: ( ) \u2212+ \u2212 \u2212= 22 arccos \u03c6 \u03c6 \u03b8 Rxy Rx CC C (27) Substituting (27) into (5) yields ( ) ( ) ( ) ( ) ( ) ( ) cos sin cos sin x A B y B A \u03c6 \u03c6 \u03c6 \u03c6 \u03c6 \u03c6 = + = + (28) 4 Copyright \u00a9 2004 by ASME rms of Use: http://www.asme.org/about-asme/terms-of-use D ( ) ( ) ( )2 C 2 C 2 C C 2 C 2 C 2 C C Rxy Rxr xRB Rxy y ryRA \u03c6\u2212+ \u03c6\u2212 \u2212\u2212\u03c6= \u03c6\u2212+ \u2212+= (29) Equations (23) and (28) represent the curves that generate the tooth profile from the rack cutter envelope. Limits\u03c6 values The three curves r1, r2 and r3 generate three profiles t1, t2 and t3, see fig. 3. The boundaries of the curves r1, r2 and r3, in terms of their parameters \u03b81, \u03b82 and \u03b83 are reflected on the boundaries of the envelopes t1, t2 and t3, in terms of \u03c61, \u03c62 and \u03c63. Using mesh eq. (22) and eq. (27) and written in terms of \u03c6 yields R r ctgx R cscx C t T e \u03b8\u03b8\u2212 =\u03c6 \u03b1\u03b8+=\u03c6 (30) The first equation refers to the straight line segments r2 and r3 (see fig. 3) of the rack, while the second one refers to the tooth fillet. The limits \u03c6 can be obtained by substituting into (30) the values for found in (19). Evaluation of intersection points: a numerical technique. By analysing fig. 3 one can argue that some parts of the envelope curves can be spurious, because of undercutting, i.e. the theoretical curves cannot be practically built. The actual intersection point A is given by: ownloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 02/07/2016 Ter ( ) ( ) ( ) ( ) 1 1 3 3 1 1 3 3 t t t t x x y y \u03c6 \u03c6 \u03c6 \u03c6 = = (31) in which xt1 and yt1 are the expressions (28) and xt3 and yt3 are the expressions (23) applied to segment r3 of the rack. Using the Newton technique eq. (31) can be solved iteratively", + " obtain the stationary point; a first trial value for the vector )0( \u03c6 is choosen. The stationary points can be minimum, maximum or saddle points. The purpose is to find the minimum; therefore, the following check is considered (k is the number of iterations): ( ) ( )( ) 21k 3 1k 1 ,f \u03b5<\u03c6\u03c6 ++ (34) For what concern )0( \u03c6 , geometric considerations yield to choose: ( )0 32 12 \u03c6 \u03c6 \u03c6 = (35) Undercutting High values of negative addendum modification give rise to undercutting at the tooth base, i.e. close to the intersection between t1 and t2, see fig. 3. 5 Copyright \u00a9 2004 by ASME ms of Use: http://www.asme.org/about-asme/terms-of-use Do Figure 4 shows how the rack generates spurious envelope branches if it is too close to the wheel. By considering figs. 3 and 4 one finds that a singularity happens for: 32 u\u03c6 \u03c6< (36) The inequality means that the arc profile r3 begins to envelop spurious branch before reaching \u03c6u, at this point the singularity is produced. \u03c6u can be evaluated using eq. (8); moreover, such singular value can be found graphically by observing fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000162_0890-6955(94)00131-3-Figure4-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000162_0890-6955(94)00131-3-Figure4-1.png", + "caption": "Fig. 4. Dimensional details of the fabricated cutter.", + "texts": [ + " The aim of the experiments is to compare the effectiveness of the designed nonuniformly spaced inserts cutter in reducing vibration with that of the standard cutter available with similar geometry. The standard cutter available for study has eight inserts 1442 S .K. Choudhury and J. Mathew S t a n d a r d c u t t e r _ _ _ D e s i g n e d c u t t e r As per the angle specification obtained in design I (Fig. 2) a cutter has been fabricated. The dimensional details and drawing of the fabricated cutter are given in Fig. 4. The proposed cutter is identical in all dimensions to that of the standard cutter except in tooth ,;pacing. The following range of cutting parameters is chosen for experiments: Speed, 200-300rpm; Feed, 2 4 - 3 8 m m / m i n and Depth of cut, 0 .1-0.3 mm. The material of the workpiece is EN-8 steel. The same carbide tips are used in the proposed design of the milling cutter. A block of 47 cm \u00d7 30 cm \u00d7 11 cm is cut on a horizontal milling machine and the vibration is recorded with the help of a vibration pick-up" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000949_1.2802323-Figure2-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000949_1.2802323-Figure2-1.png", + "caption": "Fig. 2 Newton's method", + "texts": [ + " 1), which is a nonlinear term and disables interpreta tion of iteration (17) as an \"integrator of cr\" (although it can be understood as a \"limited integrator of a\"). 5.2 Newton's Method. The second control law corre sponds to famous Newton's method (frequently referred to as Newton-Raphson method). It is characterized by F(u, t) = u - \\a(u, t)/(da(u, t)ldu), where \\ is again a constant parame ter, and daldu * 0. From (7) we have daldu = \u2014 dfldu, and this leads to the recursive control law \"(4+1) = M(f\u00ab) + X o-(4) 5 / du (4) \u00ab(4+i) e [M, mm* '^maxJ (19) illustrated in Fig. 2. Notice that a simple modification of the method is introduced here; the textbook version assumes X = 1. We introduce \\ which is not necessarily equal to unity in order to account for uncertainties in dfldu and provide a way to change the mode of convergence. Bounds on X are fixed by the convergence condition. Condition (b) of Theorem 1 yields \u201e , (dflduf - ad^fldu^ ^ 0 < X^-^^ IT < 2 {dfldu)^ which, in vicinity of u = 0, becomes 0 < X < 2 (20) (21) cr(4) in (19) is calculated by (6), i.e. from measurements" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003338_2004-01-1784-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003338_2004-01-1784-Figure1-1.png", + "caption": "Figure 1: Schematic of the 3-axle planar model as derived.", + "texts": [ + " A three-axle articulated model is expanded to contain five axles to avoid lumping the parameters for the drive and semitrailer tandems. Compromises inherent in using the linearized models are discussed and evaluated. Finally, a nonlinear tire cornering force model is coupled with the 5-axle model, and its ability to simulate a jackknife event is demonstrated. The model is shown to be valid over a wide range of inputs, up to and including loss of control, on low-and-medium-\u00b5 surfaces. 3-AXLE PLANAR ARTICULATED VEHICLE MODEL DERIVATION The system to be modeled is depicted in Figure 1. As seen in the system sketch in Figure 1, the articulated vehicle has a forward velocity, denoted by 1ue , and lateral velocity, denoted by 2ve . Note that from this point forward the under-tilde ( ie ), used to designate vectors, will be dropped for convenience. Also note that as shown in Figure 1, positive forward velocity for CG1 results in a negative rotational velocity, \u03b8 . The angle \u03c8 denotes the prime mover yaw angle, and the angle \u03b3 denotes the hitch articulation angle. For this derivation only the front axle is allowed to steer and per usual convention, the steer angle, relative to the body centerline, is denoted by \u03b4. KINEMATIC EQUATIONS AND CONSTRAINTS The only constraint on the system is the hitch constraint requiring the position, velocity, and acceleration of both tractor and semitrailer to be identical at the hitch point", + " q q d T T VL Q dt q q q \u2202 \u2202 \u2202= \u2212 + = \u2202 \u2202 \u2202 (1) 2 where: T = system Kinetic Energy term, V = system Potential Energy term, Qq = potential work expression for the generalized variable, qk, q = generalized displacement variable. The generalized coordinates (qk) will be: y = centripetal location of CG1 with respect to the center of the turn, = rotational location of CG1 with respect to the fixed basis, Ei, \u03c8 = rotational orientation of the tractor (CG1, yaw), and \u03b3 = rotation orientation of semitrailer (CG2) with respect to the tractor (i.e., hitch articulation). The reader can verify that the generalized variables above are sufficient to fully describe the system represented in Figure 1 with respect to the inertial reference frame. Equations (2) and (3) are for the position vectors for the respective CG\u2019s (\u03c11 and \u03c12), expressed in the most convenient bases. 1 1 20e ye\u03c1 \u2032 \u2032= + (2) 2 1 2( cos ) sinc d e d e\u03c1 \u03b3 \u03b3= \u2212 + \u2212 (3) where: \u03c11 = position vector of CG1, \u03c12 = position vector of CG2. 3 The respective velocity expressions can be derived by either using the relative velocity in a rotating reference frame or by using the change of bases in equations (47) through (49) (located in the Appendix) to transfer the velocity vectors to the standard (fixed) reference basis {Ei}, then simply differentiating the position equations with respect to time" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000076_0045-7825(93)90131-g-Figure4-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000076_0045-7825(93)90131-g-Figure4-1.png", + "caption": "Fig, 4, lsovalues of ~Vcq,", + "texts": [ + " There is not enough time to evacuate heat towards the head of the sample by conduction. For isothermal tests, the stress maximum is always at the periphery of the cylindrical part. On the contrary, because of the conjugated effects of the orthoradial velocity and temperature, this maximum is localized inside the sample (the stress increases with r but decreases with T). The numerical prediction shows this phenomenon (Fig. 10). $. Example of simulation of a torsion-tension test We consider the same geometry as in Fig. 4. The material obeys the rheoiogical law of (49). The rotational velocity is low (ltr/min). A tension velocity is prescribed at the other extremity of the sample (1 mm/min). The thermal computation shows that deformation is nearly isothermal. Time steps are constant and equal to 5 s. As can be found in the analytical solution of the tension test, the dependency of the longitudinal velocity on z is linear (Fig. 11). The distribution of equivalent stress (Fig. 12) 432 A. Moal et al., A simulation of' the torsion and torsion-tension tests A" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002211_0-306-46956-1_9-Figure9-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002211_0-306-46956-1_9-Figure9-1.png", + "caption": "Figure 9 Plan view of longitudinal, lateral, yaw car model.", + "texts": [ + " Usually, the braking and drive systems distribute torques equally to left and right sides, and the steering and turning geometries are such that there are hardly any differences between right and left sides except for the wheel loads. If the influences of load transfer on the tire forces of a pair of wheels belonging to one axle are ignored, which is approximately the case for mild maneuvering, it is as if the two wheels of an axle were joined together at the vehicle center plane. Capitalizing on the absence of car-position-based forces in the problem, one can describe the motions conventionally by longitudinal, lateral, and yaw velocities, as seen from the car (Fig. 9). These are sufficient to describe tire slip angles and car accelerations and to allow the derivation of the equations of motion. In the absence of the kind of differential longitudinal tire forces used in traction control and active stability control systems (Abe 1998), the equation for the longitudinal translation is only weakly coupled to the lateral and yaw equations, and the latter two describe the handling problem at this level. In linear form, they are (Dixon 1996) Solving the equations of motion for steady turning reveals that the curvature response to steer angle ratio is given by the character of which depends on whether K is positive or negative (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001321_s0026-2692(98)00038-x-Figure5-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001321_s0026-2692(98)00038-x-Figure5-1.png", + "caption": "Fig. 5. The bridge structure of a Hall plate.", + "texts": [ + " The sensitivity is alfected by all the above-mentioned non-linear effects as well as temperature. During the model development, we have built the macro-model incorporating all these effects. Each effect is modelled separately in such a way that sensitivity S introduced by Eq. (2) in the macro-model becomes an almost independent quantity. Now, the macro-model sensitivity depends only on the sensor geometry and temperature. In order to simplify a description of the modelling procedure, we shall take as an example a conventional Hall plate (Fig. 5). As has already been mentioned, any vertical Hall sensor corresponds to an equivalent Hall plate. Fig. 5 shows a bridge model of a Hall plate. The four branches of the bridge are situated between each pair of the neighbouring n+ contacts. The ohmic resistances of the contacts are also shown. The active area of the Hall plate is situated between the current supplying contacts Ic and 12. The border lines containing sense contacts Vt and V~ are the limits of the depletion zones of the pn-junctions. These pn-junctions correspond to the ones between the active volume and the p-well ring of the vertical Hall sensor. There is also a pnjunction in the third dimension due to the shallow p-shield in the vertical Hall sensor. All pn-junctions are distributed among four branches shown in Fig. 5. Regarding the recommendation in Ref. [6], each branch in a bridge representation of a vertical Hall device is modelled by the use of the circuit shown in Fig. 6. It is an ELDO [7] model capable of simulating material and geometry nonlinearity as well as the junction field effect. The junction field effect is modelled by means of a JFET. All JFET parameters needed are extracted from the data related to the potential and current distribution within the active volume, and obtained by the device simulation in MEDICI" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003638_s021812740401103x-Figure4-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003638_s021812740401103x-Figure4-1.png", + "caption": "Fig. 4. Examples of Poisson spheres with contour lines of the effective potential Ul(\u03b3). The point \u03b3 = e1 points backward to the upper left. The three cases on left belong to the \u03b1-range I with \u03b1 = 0.505. From left to right: l = 1.82 (values of the potential increase from red to magenta to red), l = 1.85 (red to magenta to yellow), l = 2.1 (red to yellow). Picture on right: \u03b1 = 1.5 (range III), l = 3 (green to red).", + "texts": [ + " Colors correspond to topological types of energy surfaces E 3 h,l, their changes indicate bifurcations. The two pictures on left in Fig. 3 represent the \u03b1-range I, corresponding to a disk-like (oblate) mass distribution. The bifurcation diagram defines four (h, l)-ranges corresponding to different topological types of energy surfaces: one three-sphere S3 (red), two disjoint three-spheres (yellow), the projective space RP 3 (blue), and the union of a threesphere with the direct product S1 \u00d7 S2 (magenta). These assertions are derived from an analysis of the effective potential (6) of which Fig. 4 gives four examples. (The situation of low l-values is omitted there because in the color code used it would simply be represented by a red sphere.) Let us explain some details, and how the colors on the Poisson spheres are related to the colors in Fig. 3. Notice first that the effective potential on the Poisson sphere Ul(\u03b3)|S2 = l2 2A1(\u03b1 + (1 \u2212 \u03b1)\u03b32 1 ) \u2212 \u03b31 (10) depends on \u03b31 only; the equipotential lines are circles \u03b31 = const. on the Poisson sphere. A schematic view of their organization is given in Fig", + " Given a red point (h, l) in the bifurcation diagram, the projection Uh,l of E3 h,l is a disk D2 \u223c S2\\D2 on the \u03b3-sphere which is bounded by an equipotential line in the red region; the corresponding type of E3 h,l is S3. For sufficiently low values of l2, only the red and blue types of energy surfaces occur; cf. the left picture in Fig. 5. For l-values in a small range around 1.8, see the blow-up in Fig. 3, the potential Ul(\u03b3) has a degenerate relative minimum (h = 1.807 for l = 1.82) along the right black circle in the magenta region of the leftmost picture in Fig. 4. With these values of h and l, the Lagrange top is in a stable relative equilibrium, rotating at a fixed inclination with respect to the direction of gravity. With slightly higher values of h, there exists, in addition to the disk with center at \u03b3 = e1, an annulus S2\\2D2 of accessible \u03b3-values around that circle; hence the energy surface has two disjoint components: a three-sphere and a direct product S1 \u00d7S2 (magenta). As energy increases towards h = 1.81825, a degenerate relative maximum along the left black \u03b3-circle is reached" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000322_218013.218068-Figure15-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000322_218013.218068-Figure15-1.png", + "caption": "Figure 15: Examples of set operation", + "texts": [ + " If an intersection line crosses with these edges, we consider that the intersection poil)t is incident to an uncertain vertex region and make the sector test for faces which are picked from a global vertex cycle traversing only the small edges. This corresponds to using the adaptive tolerance [4]. After we obtain the intersection table for (he loop, we calculate the local structure again. Using this metl]ocl, we do not make small faces, instead we generate a consistent local structure composed of face sectors. 4.4 Examples In Fig. 15, we show some simple examples of set operat ions made by the system which employ t lie proposerI algorithms. Although the system I)as not been completed, we have confirmed that our algorithms are effective. Figure (a) shows a solid made by subtraction operations to a cube. Figure (b) shows a solid made by union of two regular tetrahedral. It includes vertices which are made by intersection of two edges. Figure (c) shows a solid made by union of two regular octahedra. It includes coincidence of vertices" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002416_871645-Figure3-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002416_871645-Figure3-1.png", + "caption": "Figure 3. Points Checked on Tooth Surface", + "texts": [ + " GRAPHICAL OUTPUT An important part of each of the coordinate inspection methods is the pattern of points on the tooth to be measured. The pattern described here has been derived by an evolutionary process with much feedback from people in Engineering, Qual ity Control, and Manufacturing. In a manner analogous to the common practice of plotting profile and lead traces for spur and helical involute gears, the points checked on the bevel gear or pinion are arranged on three profile 1ines and one lengthwise line on the tooth (Figure 3). The plots (Figures 5 and 6) look very similar to the famil iar lead and profile traces for spur and hel ical gears, but, instead of being continuous, they are made up of discrete points connected by straight 1ines. Any number of points can be calcul ated for any number of profiles. For routine qual ity checks, experience has shown that three profile patterns and one lead pattern are sufficient. The straight lines connecting the measured points convey no information about the surface between them" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003670_j.jappmathmech.2004.11.006-Figure2-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003670_j.jappmathmech.2004.11.006-Figure2-1.png", + "caption": "Fig. 2", + "texts": [ + " The corresponding phase trajectory asymptotically approaches the point / +/ = - - \" 2 ~ 2 2 V* 4 ~ tg[~* = 42izTo_mob Kz (2.7) Stationary solution (2.7) represents the uniform rotation of the platform around a vertical axis. The pointA in this case moves with constant velocity V* along a circle of radius bctg~*. If the initial conditions for the differential equations are such that ~ T0 < system (1.4) will not have stationary solutions. Possible types of trajectories of the point A, obtained by numerical integration of system (1.3), are shown in Fig. 2 in the case when inequality (2.8) is satisfied (a) and when the system reaches a steady state (2.7) (b). 3. THE MOTION OF THE CARRIAGE WHEN THERE IS AN ELASTIC MOMENT We will consider the case when the moment of the pair of forces is proportional to the angle of rotation [3 of the front wheeled pair of wheels with respect to the platform M = - K ~ (3A) Here K is the stiffness of the corresponding spring. Equations (1.4) can then be written in the form K~t0(13) m 2 V~, V - Kl]tgl3 mltg[3 Vo~ (3.2) = co, 6J - 12~tl(~) b~l(~)cos [~ b~l(~) ~tl(~)cos2 ~ Remark 2" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000544_20.767200-Figure4-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000544_20.767200-Figure4-1.png", + "caption": "Fig. 4. Magnetic flux plot for the analyzed model when the rotor position is at 90 *", + "texts": [ + " We can rewrite (10) in integral form as where ST is the cross section area of the rotor steel shell. From (11) we obtain (12) D. System matrix B. Circuit equations After all, i f (4), (9) and (12) are combined, a system matrix of which unknown variables are A. Z and v~ can be constructed as follows~ Since the BLDC motor is driven by a voltage source I304 Cldt 0 0- = Qldt Lldt 0 I Sldt 0 0. K+ Clot K A \u2019 - j f + A f I + . V 4 - t . ( 13) IV. ANALYSIS RESULTS A. Magnetization distribution and back-emf wave form Fig. 4 shows the magnetic flux plot of the analysis model when the rotor position is a t 90 electrical degree. The magnetization distribution in the permanent magnet is obtained by analyzing the magnetization process of the bonded Nd-Fe-B material using the finite element method (51 and it is shown in Fig. 5. Actually, this magnetization distribution is used by interpolating the curve using cubic spline method in the finite element analysis. 5 c Y o 9 0 180 360 Rotating electrical angle(deg.) Fig. 6. Back-emf wave form compensating for fringing flux due to the transverse edge of the magnet" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002410_2000-gt-0643-Figure4-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002410_2000-gt-0643-Figure4-1.png", + "caption": "Figure 4: Scheme of the operational principle of the active lubrication.", + "texts": [ + " 2 illustrates the interior of the bearing, built by 4 pads which are orthogonally arranged in the y and z directions. It is possible to see the ori ces machined on the pad surface. Fig. 3 shows the pipeline connection of the electronic injection and two different pad geometries: with 5 and 15 ori ces. The pads are built by two parts, which attached to each other by screws create a small sealed reservoir ful lled by pressurised oil. These reservoirs are shown in the scheme of operational principle (see Fig. 4, active injection). In Fig. 1 the test rig is presented, focusing on the servovalves, and pipelines attached to the bearing. The oil (lubricant) coming from the servovalves is controlled using a feedback control law. The feedback law is designed using displacement and velocity signals as shown in the scheme presented in Fig. 4. It is important to point out that the in uence of the piping injection system on the rotational motion of the tilting-pad is very small. The piping line contributes with a very low torsional sti ness which can easily be calculated using the piping line dimensions. Comparing the torsional sti ness 2 Copyright (C) 2000 by ASME Terms of Use: http://www.asme.org/about-asme/terms-of-use D of the pad resulted from the oil lm to the piping sti ness the rst one can be neglected. High response servovalves of MOOG are used in this application" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002717_auv.1996.532403-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002717_auv.1996.532403-Figure1-1.png", + "caption": "Fig. 1", + "texts": [ + ". INTRODUCTION The development of underwater vehicles that can navigate autonomously in the hostile environment of open sea requires enormous investments both in the hardware and software development. The Autonomous Underwater Vehicle (AUV) Twin-Burger (Fig. 1) [Fujii, Ura 1995, also see http:// underwater.iis.u-tokyo.ac.jp/robot/twinburger-e. html] has been developed as a testbed for various new technologies in underwater robot ics . I t has 2 large rectangular , hydrodynamically bulky hulls to house various electronic equipments. They are placed on top of a heavy cylindrical battery compartment. This configuration allows experiments be conveniently carried out in a testing pool and electronic equipments be easily accessible. Open frame structure is adopted for the main structural elements to make the attachment of additional equipment (such as lights or acoustic beacons) simple", + " EXPERIMENTAL RESULTS OF SURGE, YAW AND HEAVE TRACKING The performance of the controller is verified by a series of experiments carried out in a fresh water test pool in the Institute of Industrial Science of the University of Tokyo. Discrete Butterworth low pass filter of cut-off frequency 1Hz is used throughout the controller. The system ran at 10Hz. The performance of simultaneous heaving and surging tracking is shown in Fig.8 and Fig. 9, respectively. In Fig. 8(a), the components U,, u2, ui and the resultant output at the thrusters 4 and 5 (refer to Fig. 1) are shown. It should be noted that output at thrusters 4 and 5 are identical. In the initial 8 seconds fixed command sequence sent the robot to a depth of about 1 meter. When the controller was set active thereafter, the large initial error caused a big integral error in the buoyance estimation (positive for floating, negative for sinking). But soon afterward a good estimation of buoyance have been obtained. If the estimated buoyance curve is carefully examined it could be found that the buoyance of the Twin-Burger fluctuated at a 25 second cycle, which is in fact the time duration of its heave motion" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002350_robot.1998.680983-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002350_robot.1998.680983-Figure1-1.png", + "caption": "Figure 1: Image of casting manipulator", + "texts": [ + " Extending the link length or adding extra degrees-offreedomE1, 21 will produce other problems, such as an increase in weight or inefficiency of operation. Mobile manipulators have been also studied in order to enlarge the work space[3]. However mobile manipulators cannot reach some places, such as a deep depression or a steep slope. In order to solve such problems we proposed a new type of manipulator, which we call a casting manipulator[4]. A casting manipulator consists of rigid links, a flexible string, and a gripper as shown in Figure 1. By throwing the gripper to a target by changing the length of the flexible string, a distant target can be reached without having to move the body. In spite of its small size and simple mechanism, the manipulator can have a large work space by making effective use of its dynamics compared with the conventional manipulators. The manipulator uses substantially less energy to move and does not suffer moving tumbles. Therefore it is possible to use this manipulator to collect objects such as garbage floating on the sea, or in agriculture for collecting various kinds of Casting manipulation consists of the following five phases", + " These computers, the DOS-V machine and Tracking Vision, are controlled by VxWorks in order to keep a short constant cycle time. The cycle time of the DD motor control is 1 msec. The cycle time of the Tracking Vision is 33 msec. One workstation SPARC Station is used for the man-machine interface. 4.2 Swing Control For the preparation of swing motion control, estimation of the joint angle 0 2 was investigated. In the case of using the encoder or potentiometer, friction of joint 2 becomes a serious problem. Then we use the force-torque sensor shown as Figure 1 1 . The joint angle 0 2 is estimated from the force information f, and f,, with the following equations: e? = atan2 (Fy -hc , Fx -$w) + n t2 m is the total mass of the sensor plate and the stopper. With the compensation of the inertial force and the gravitational force this method gives a fairly good estimation. The swing experiments about generation of constant swing (91=50 degree) from initial stationary state were carried out. This experiments corresponded to the simulation described in the previous Section" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001984_jsvi.2001.3643-Figure7-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001984_jsvi.2001.3643-Figure7-1.png", + "caption": "Figure 7. Slotted composite shaft model.", + "texts": [ + " Using the same theory, analysis has been carried out for the slotted composite shafts in the present study. Figure 5 shows a composite shaft created using a four-noded shell element (CQUAD4) which is shown in Figure 6. The CQUAD4 element is a four-noded, bilinear, isoparametric element capable of representing membrane, bending (with transverse shear e!ects) and membrane-bending coupling behaviour. The property of the laminae is supplied using PCOMP property cards. A two-dimensional orthotropic material is used for the analysis using MAT8 card. Figure 7 shows the discretized model of the slotted composite shaft with 1100 elements. Four layers of equal thickness (0)004 m) with di!erent stacking sequence are considered for this problem. Dimensions of the shaft are the same as shown in Figure 1. Simply supported boundary conditions are considered for the shaft ends. The eigenvalue analysis is carried out in NASTRAN in a similar way as was done for the isotropic shaft. TABLE 1 E+ect of longitudinal slots and inertia slots on eigenfrequencies of slotted isotropic shaft Mode no" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003054_6.2002-533-Figure3-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003054_6.2002-533-Figure3-1.png", + "caption": "Figure 3: Aluminum Bottom Housing -01.7500", + "texts": [], + "surrounding_texts": [ + "work for cases where the flows of interest are already well understood, but they are not reliable for complex flows such as 3D and/or unsteady cases, cases with rough or curved walls, flows with injection or suction, mixing with foreign fluid injection or high-speed flows especially those with impinging shock waves, high enthalpies and/or combustion. A good discussion of such indirect methods is in Nitsche et. al [2].\nThe purpose of this investigation was to develop a non-intrusive, direct skin friction-sensing device for high vibration and acceleration loads without compromising the resolution of the data acquired. This new concept was needed to alleviate problems encountered by existing sensors under such conditions. To meet this goal, various damping techniques had to be studied and applied in the design of the skin friction gage.\nAfter the analysis of several damping methods, it was decided to employ visco-elastic damping through the use of a rubber RTV sheet on the gage surface (See Fig. 1). Viscous damping through the use of oil inside the gage housing would provide further damping. Previous skin friction gages at Virginia Tech [3] used viscous damping due to several inherent advantages. By filling the volume around the sensing head and flexure, pressure gradient effects are minimized. The fluid also acts as a temperature stabilizer, minimizing the temperature gradients that the gage is exposed to during the flight test. However, gages that previously employed this method of damping deteriorated with time due to fluid loss through the small gap around the sensing element.\nA proposed solution to this problem was to fill the gap with a substance that would check this flow. Previous studies had tried using silicone rubber in the entire inner volume of the gage, but loss in gage resolution due to increased resistance forces limited further development [4].\nA study showed that a rubber RTV compound used to cover the sensing element and gap with a thin sheet as illustrated in Figure 1 could be a possible solution [5]. The apparent advantage of this method was that fluid could be retained in the gage cavity, thus keeping the inherent benefits while eliminating the maintenance requirements. However, several aspects needed further investigation, including, but not limited to, RTV adhesion characteristics, trapped air in the fluid volume and the scaling of the design to meet sensitivity requirements of this work.\nThis study was performed to help the successful integration of a rubber RTV compound sheet into a skin friction gage design capable of meeting NASA flight test guidelines of vibration conditions up to 8.0 g^ over a 15 - 2,000 Hz frequency range, requiring performance at altitudes ranging from 15,000 to 45,000\nfeet, Mach numbers ranging from 0.6 to 0.99 resulting in shear values of 0.3 to 1.5 psf. Flow Direction Rubber RTV\nIn order to meet the vibration requirements, a primarily viscoelastic damping method was employed on a non-nulling gage design. RTV 566, a rubber compound obtained from GE Silicones [6] was applied on the gage surface to contain the fluid. Several benefits were realized by the use of this compound. In addition to containing the fluid, it provided sufficient damping itself, minimizing the need for the use of a fluid in the volume cavity. In some cases, glycerin was used to fill the inner volume of the gage :to provide additional viscous damping.\nThe use of the rubber compound presented new challenges as well. First, the RTV did not bond well to materials such as aluminum despite the use of a metal primer. It was suspected that an oxide layer formed on the aluminum surface preventing adhesion. Often the test fixtures on flight test vehicles are made of aluminum. However, the RTV compound successfully bonded to polyethersulfone (PES), a high temperature plastic so that could be used for the gage housing. Second, the RTV sheet reduced gage sensitivity, because the rubber sheet bears part of the load due to shear on the head. This sensitivity reduction was greatly aggravated by any accumulation of the compound in the gap between the sensing head and housing, requiring the modification of the earlier procedure used in the application of the rubber RTV compound onto the gage surface.\nAmerican Institute of Aeronautics and Astronautics", + "SKIN FRICTION GAGE\nThe skin friction gage used in this study consisted of six components: 1) sensing head, 2) flexure, 3) base, 4) upper housing, 5) lower housing and 6) connector. See the schematic in Figure 1.\nSensing Head The sensing head was made of polyethersulfone (PES). This reduced mass compared to a metal head and aided in bonding the rubber sheet. It had a head diameter of 0.75 inches, chosen in conjunction with the flexure described below to meet the design shear levels. The primary feature of the component was its form as a truncated cone with tapered edges from the bottom, a quality that reduced the entrapment of air bubbles during the filling of the volume cavity. A 0.25 inch threaded hole allowed for the fastening of the component to the flexure. See Figure 2.\nFlexure and Base The flexure, made of aluminum, had a shaft of length 1.538 inches and a diameter of 0.150 inches using beam theory. Aluminum was chosen because of its low density, which reduced vibration effects. The high heat conductivity of aluminum would act as a heat sink for the semi-conductor strain gages, which are temperature sensitive. To account for temperature effects on the strain gages, a type K thermocouple was buried into an opening created at the center of the flexure.\nMicron Instruments installed a matched set of eight SS-060-033-500P-S(4) semi-conductor strain gages on the base of the flexure. These allowed for the data acquisition on two perpendicular axes, a trait necessary in measuring skin friction in a 3D-flow regime. An illustration of the flexure and base can be seen in Figure 2.\nA primary goal in the gage design was robustness, which is required for flight application. To achieve this, protection of the strain gages was necessary. The first step was the introduction of a 10-pin connector that isolated the gage wiring from the data acquisition equipment. Next, solder pads located on the base were used to isolate the delicate strain gage wires from the wire leads from the connector. Further, the wires were looped around to provide strain relief and anchored to the base via epoxy. The base had two fill holes that allowed the filling of the inner cavity with a viscous liquid.\nGage Housings A two-piece housing was used in the skin friction gage design. The bottom piece was constructed of aluminum and housed the base, flexure and sensing element. To improve the design robustness, a deflection constraint was incorporated into the top surface of the aluminum housing. PES was the material of choice for the top piece of the housing. This allowed for easier application of the rubber compound onto the gage surface. Further detail on the gage housings can be, seen in Figures 3 and 4. Gage assembly is depicted in Figure 5.\nAmerican Institute of Aeronautics and Astronautics", + "Connector A standard 10-pin plug connector, supplied by Spacecraft Components Corporation [7], was used to interface the skin friction gage with the data acquisition equipment. Each bridge circuit used four pins from the connector. The remaining two pins, made of chromel and alumel were used for the thermocouple buried in the flexure. A matching socket was used to route the wires to the data acquisition equipment. A second thermocouple was buried in the rubber sheet. The connector was secured to the base of the flexure base as demonstrated by Figure 5.\nGage Electronics In a full-bridge circuit, all four strain gages in the Wheatstone bridge are placed on the base of the cantilever beam. In each axis, a strain gage pair with the same vertical alignment is placed in the tensile region while a similarly aligned pair is placed in the compressive region 180\u00b0 from the first pair. For a dualaxis scheme, a similar strain gage layout is done perpendicular to the primary axis. Figure 2 displays a strain gage pair attached to the base of the aluminum flexure.\nGAGE PREPARATION\nA major problem encountered in previous studies with the use of the rubber RT V was its adhesion to the gage surface during testing due to the aggressive nature of the supersonic wind tunnel \"starting\"[5]. It was determined that bonding could be improved by creating grooves on the bonding surface and machining down the outer lip of the housing allowing the rubber compound to wrap around the outer edge.\nA new problem encountered in this study was the unwanted filling of the gap between the sensing head and the plastic housing with the rubber RTV mixture causing loss in gage sensitivity. The previous application procedure required the inverted gage surface to be placed onto the liquid RTV mixture without any support of the gage while the RTV set.\nThis allowed the weight of the assembly to squeeze the RTV mixture into the gap leaving a thin layer of the RTV on the actual bonding surface and a ring of RTV in the gap. This was only a minor problem with the earlier gages, which were small and made entirely of plastic. The gages designed for the present application were larger and heavier, thus greatly amplifying the problem.\nTo counter this problem, a gage adapter was designed. It had an inner cavity into which the gage assembly could be inserted and fastened. The adapter had a shank designed for insertion into a milling machine. This setup allowed the gage to be held at a desired height above the surface upon which the liquid RTV mixture was spread, and this reduced gap filling with rubber RTV resulting in increased gage sensitivity.\nA known characteristic of rubber compounds is the change of modulus of elasticity with temperature. For wind tunnel testing, the exposure time to the flow is on the order of seconds, negating this effect. However, for flight-testing, the testing time varies from 30 minutes to several hours at varying altitudes making the temperature effect a major concern. To account for this effect, a thermocouple was buried into the silicone RTV sheet over the gap between the sensing head and the housing. This enabled the calibration of the unit at different flow temperatures, emulating the final test environment the gage would experience.\nCALIBRATION\nIn this study, two calibration methods were used: a) Point loading b) Distributed Shear\nPoint Loading Method A known weight is placed parallel to the direction of flow and perpendicular to the sensing element. This is usually achieved by hanging a paper cone by thread attached to the floating element with clear tape as illustrated in Figure 6a. A tare value of the cone and string is taken to zero the balance. Different weights ranging from 50 milligrams to 5 grams are placed in the cone while the corresponding output is recorded. The gage is then rotated 180 degrees and the procedure repeated. The whole process is repeated on the crossstream axis. This is the preferred method of calibration for oil-filled gages due to its simplicity. It is also a highly accurate method due to the negligible shear stress contribution of the oil in the gage cavity\nThe mass calibration is then related to shear through:\nCD where K is the conversion factor from mass to force\nunits, M is the mass of the calibration weight and A is\nAmerican Institute of Aeronautics and Astronautics" + ] + }, + { + "image_filename": "designv11_65_0002843_12.421041-Figure7-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002843_12.421041-Figure7-1.png", + "caption": "Figure 7: Detection of a crack in an aluminium plate and the resultant damaged rivet joints.", + "texts": [ + " The rivet joints hidden under black paint had the same geometrical dimensions, so it was impossible to distinguish one from the other. But each of the two blind rivet systems provides a different compressive stress because of the varying material combinations. This effect can be visualized with ULT. The phase image taken at 0.5 Hz and a power of 600 W demonstrates that it is possible to distinguish the copper-brass rivet from the aluminum-steel system. The detection of a crack along a row of rivets as an example for the maintenance and inspection of safety relevant structures is shown in Figure 7. The crack length had been known from eddy current inspection (figure 7a). At first we investigate the riveted aluminium structure using OLT. In figure 7b it is obvious that the amplitude image taken at 0.11 Hz is sensitive to the non-uniform intensity distribution. As the phase image (figure 7c) is insensitive to all kinds of perturbations, it shows essentially the thermal features from the surface down to a depth of about two times the thermal diffusions length \u00b5. One can recognize clearly the reinforcement of the aluminium plate on the right and the apparent intact riveting. No damage and no crack could be detected. But using ultrasonic excitation a bright area was found with a significantly larger extension (figure 7d). These rivets provide a reduced compressive stress so that the integrity of the riveting is no longer sure. The ULT measurement revealed that the damaged area is larger than expected from the eddy current results. To detect the crack only and to reaffirm the eddy current results all rivets were removed and an amplitude (figure 7e) and a phase image (figure 7f) were taken again with ULT. As there was no more any rubbing contact to the rivets or the rib, only the tip of the crack caused hysteresis losses whose locations are identical with the result of the eddy current measurement. The hot spot on the left is no defect, it was caused by the rubbing contact between rib and plate. This example shows how efficiently ULT can be applied for the selective imaging of fatigue cracks. Proc. SPIE Vol. 4360 571 Downloaded From: http://proceedings.spiedigitallibrary" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003068_nafips.2001.943669-Figure2-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003068_nafips.2001.943669-Figure2-1.png", + "caption": "Figure 2: Membership functions for [ z o , ~ ] .", + "texts": [ + " Then COs(Z6) cos(z,), taking into account the bounds from above, can be bounded by two constant functions: 0.7071 < cos(x6) < 1 and 0.7071 < cos(z7) < 1, which gives 0.5 < cos(z6) cos(z7) < 1. Now the three nonlinear terms can be fully described by the upper and lower bounds derived above in the following manner: where F!, F: E (011, Ff = 1 - F: and Fi = 1 - F,'. By solving the above equations for Ft, F:, F,' and F: we obtain the following membership functions The graphs of the membership functions Ft and F: are shown in Fig. 1 and the graphs of Fi and F; in Fig 2. The TS rule base is then expressed as follows: 1 : IF 211 IS Ff and [26 ,27] IS Fi THEN X = A ~ z + BU 2 : IF 211 IS Ff and [ 2 6 , ~ 7 ] IS F i THEN X = A ~ x + BU 3 : IF 211 IS F: and [ZS, 271 IS F i THEN X = A32 + BU 4 : IF 211 IS F: and [26,27] IS F . THEN X = A42 + BU In the above rules the matrix A1 is the Jacobian obtained by Taylor series expansion of (2) for values of 26, 27, and zll such that F:(zll) = 1, and F,j((z6,27) = 1. The rest of A2, AS, and A4 are obtained in the same manner" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002685_robot.1999.772524-Figure3-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002685_robot.1999.772524-Figure3-1.png", + "caption": "Figure 3: The \u201c321\u201d parallel manipulator architecture.", + "texts": [ + " The best-known examples are the serial 6R robots (with six revolute joints) and the parallel StewartGough platforms, [8, 11, 191, (with six prismatic joints). Indeed, the pairing based on the reciprocity operation (Sec. 2.1) maps a twist along a revolute joint into a wrench along a prismatic joint. Definition 4 (Projective dual) A parallel and serial manipulator are projectively dual i f they are topologically dual, and corresponding joint axes intersect in the same way. Examples of this are the Puma-like serial manipulator, and the \u201c321\u201d parallel manipulator, [13] (Fig. 3). The former has parallel second and third revolute joint axes (parallel means: intersecting at infinity), and intersecting fourth, fifth and sixth joint axes; the latter has a set of three intersecting prismatic joints, and another set of two intersecting prismatic joints. Definition 5 (Metric dual) A parallel and serial manipulator are metrically dual i f they are projectively dual, and their corresponding joints are each other\u2019s dual. Note that both projective duality and metric duality have weaker, instantaneous forms, and stronger, global forms (\u201cglobal\u201d means that the duality remains valid over a range of joint values, not just at one single configuration)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003139_tencon.2000.892289-Figure6-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003139_tencon.2000.892289-Figure6-1.png", + "caption": "Fig. 6 The fuzzy surface of the simulated system with a AMRFC controller and added noise after 500 seconds", + "texts": [ + " Implementation and Results In order to estimate the performance of the AMRFC, it needs to be compared with conventional controllers, from which the proportional-integral controller (PI) and the model-reference adaptive controller (MRAC) are chosen. The various control algorithms are first tested in simulation using MATLAB with noiseless and noisy models of the water level system. Fig. 4 shows a typical simulated step response of the noisy system starting from vacuous initial conditions, Fig. 5 shows the corresponding controller action, and Fig. 6 shows the fuzzy surface u(u,,y) after training the system for 500 seconds. Then real-time implementations with the aid of the microcontroller board are performed. Fig. 7 shows a typical real-time response of the system after training the system for 1900 seconds, Fig. 8 shows the corresponding controller action, and Fig. 9 shows the fuzzy surface u(uC, y ) after training the system for 500 seconds. The board is based on Motorolas 68HC11Al with an RS232 communication port that facilitates the communication with a normal PC, which has the command and monitoring software required for monitoring the microcontroller's performance", + " 111-362 To cancel the effect of the transient period, only results between the 1000 seconds and 2000 seconds are used in calculating these indices, which are: Table 3 gives them for real-time implementations. Comparing the results obtained from simulation and realtime implementations, it is found that the MRAC and AMRFC performance is much better than the PI controller. AMRFC performance is as good as the MRAC, but it is expected that the AMRFC would have performed better if the plant were a nonlinear one. Finally, the difference in the fuzzy surface between simulation, Fig. 6, and and real-time implementation, Fig. 9, is due to the fact that the real system is not symmetrical because the pump pumps the water in and gravity drains it out, whereas in simulation the plant is symmetrical. VI. References [I] C.C. Lee, \"Fuzzy Logic in Control Systems: Fuzzy Logic Controller, Part I and II\", IEEE Transactions on Systems, Man, and Cybernetics, Vol. 20, No. 2, G. Gateau, P. Maussoin, and J . Faucher, \"Investigation on Adaptive Fuzzy Controllers\", Proceedings of the Fourth IEEE Conference on Control Applications, pp" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000087_107754639900500506-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000087_107754639900500506-Figure1-1.png", + "caption": "Figure 1. Geometry of a helicopter rotor blade.", + "texts": [ + "comDownloaded from 763 moment trim condition. It can be seen that the algorithms are simple, behave well, and do not show numerical instabilities. Specifically, the procedures do not have multistage design processes and therefore no iterations involved. A more elaborate control of a full flap-lag motion of the individual helicopter rotor blade will be the topic of discussion of our future paper. 2. ROTOR BLADE MODEL The equations of motion for a centrally hinged, spring-restrained, rigid, single helicopter rotor blade (c.f. Figure 1) in forward flight can be written as (Johnson, 1980; Papavassiliou, Friedmann, and B€nkatcsan, 1994) where,8 is the flap angle and 0 is the pitch angle of the rotor blade. In the above equation, prime refers to differentiation with respect to y The elements of the various matrices of equation (1) are at MICHIGAN STATE UNIV LIBRARIES on April 15, 2015jvc.sagepub.comDownloaded from 764 The various parameters in the above elements are defined in the following waYOJF is the nonrotating flap frequency, y is the lock number defined as the ratio of the aerodynamic to inertia forces of the blade, If" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002448_6.1989-1202-Figure3-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002448_6.1989-1202-Figure3-1.png", + "caption": "Fig. 3: Schematic of two-link Planar Manipulator", + "texts": [ + " The motion equations for the two-link flexible planar manipulator, consistently linearized in small elastic deflections and speeds, were programmed in FORTRAN and implemented in a VAXstation 2000. The code allows for a maximum of four cantilevered modes per beam. The mass matrix is inverted using LU decomposition and a Runge-Kutta fourth-order scheme with adaptive time step is utilized for the time integration [17]. Energy and angular momentum checks arc built into the simulation. Table 2 shows the physical properties assumed for the arm (see also Fig. 3). These numbers were chosen to mimic an actual experimental testbed presently under construction at Martin Marietta by Dr. Eric Schmitz. D ow nl oa de d by P E K IN G U N IV E R SI T Y o n Ju ne 2 2, 2 01 6 | h ttp :// ar c. ai aa .o rg | D O I: 1 0. 25 14 /6 .1 98 9- 12 02 The smooth reference slew maneuver is obtained by specifymg initial and final rigid body configurations, speeds, and accelerations using a quintic equation [4]. The open loop torques for the reference maneuver are then obtained using computed torques (see Figs", + " It must be stated that at the time this paper is submitted, it is still not clear whether this failure is a correct prediction of the inconsistent model or a numerical failure. However, for the other two models, energy and angular momentum checks seem to indicate that correct results were obtained for the nominal trajectory. Physical Ropaties of Planar Manipulator with Two Flexible Links M a s s of shoulder body (kgl 20 .0 Lengthoflink 1 (m) 0.9144 Mdss density of link 1 W m ) 1.33937 Length of link 2 (m) 0.9144 Mass of elbow body (kg) 14.0 M a s s density of link 2 (kg/m) 0.669685 Other lengths (sae Fig. 3): Mass of tip body (kg) 2.0 bl (m) 0.0762 bzl (m) 0.0762 Moments of Inertia (about axis perpendicular to plane): bz2 (m) 0.0127 should^ body (kgm2) 0.01 4 (m) 0.0508 Elbow body (kgmz) 0.03 np body (kgm2) 0.01 Fust four \"narural frequenc~es\" of the arm (configuration dependent): Elbow angle (deg) Narural frequencies (radlsec) Angle: 0 Mode 1 42.90 Angle: 45 Mode 1 42.79 Mode 2 68.06 Mode2 68.03 Mode 3 153.66 Mode3 153.55 Mode 4 224.26 Modc 4 224.26 Angle: 90 Mode 1 42.71 Mode2 68.01 Mode3 153.46 Mode4 224" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002529_00037-1-Figure6-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002529_00037-1-Figure6-1.png", + "caption": "Figure 6 Wyoming modified celanese compression test fixture.", + "texts": [], + "surrounding_texts": [ + "dized by ASTM, Hofer and Rao (1977) of the Illinois Institute of Technology Research Institute (IITRI) introduced what became known as the IITRI compression test fixture, shown in Figure 4. The two major deficiencies of the Celanese fixture were eliminated. Flat wedge grips, like tensile wedge grips but inverted, were used. These could accommodate a range (about 2.5mm) of specimen thickness, while always maintaining full contact with the\nholders. Also, the \u00aaalignment\u00ba sleeve was replaced by posts and linear ball bushings, which truly do maintain alignment. It is essentially impossible to bind up linear bushings.\nThe deficiency of the IITRI fixture is its extreme weight, typically about 45 kg (WTF, 2000). In comparison, the Celanese fixture weighs just about one-tenth as much, about 4.5 kg. This makes the IITRI fixture very cumbersome to handle. Being so large, and thus requiring considerable machining, it is also very expensive to fabricate. A typical selling price of the least expensive version of the IITRI fixture is about US$8000 (WTF, 2000). For example, this is more than twice as costly as the Celanese fixture, which itself is not an inexpensive fixture.\nOne reason for the large mass of the IITRI fixture is that the rectangular holder blocks do not carry the forces induced by the wedge grips as efficiently as the circular holders of the Celanese fixture. Another reason is that most IITRI fixtures are capable of testing a specimen up to 38mm wide, and 15mm thick in the tabbed regions. It will be noted that the standard Celanese fixture is designed to accommodate a specimen only 6.3mm wide and 4mm thick. The much higher forces required to fail the larger specimen require a sturdier fixture.\nThe IITRI test fixture was added to ASTMD 3410 in 1987, as Method B. It will be noted that this was a full 10 years after it was first introduced into the open literature. This is not untypical for new test methods and new test fixtures.\nTo reduce both weight and cost, a smaller version of the IITRI fixture is available (Adams and Odom, 1991; WTF, 2000). Shown in Figure 5, this fixture can accommodate a", + "12.7mm wide specimen up to 7mm thick. It weighs only 10.5 kg, about one-quarter that of a standard IITRI fixture. Its cost is similar to that of the Celanese fixture.\nThere are at least two so-called modifications of the Celanese fixture also. The Wyomingmodified Celanese fixture, shown in Figures 5 and 6, retains the efficient circular shape of the holders, but the alignment sleeve has been replaced by posts and linear bushings, like the\nIITRI fixture. Also, the wedge grips are tapered circular cylinders rather than cones. Thus, they make full contact with the holders independent of the specimen thickness, just like flat wedges. However, the circular cylinder wedges distribute the clamping force reactions more uniformly. The standard design utilizes a tabbed specimen only 114mm long, i.e., 25mm shorter than the IITRI and Celanese fixtures, demonstrating that the 64mm long tabs commonly used with those fixtures are really longer than necessary (Irion and Adams, 1981; Berg and Adams, 1989; Adams and Odom, 1991). Specimens up to 12.7mm wide and 6mm thick can be accommodated. The fixture weighs 4.5 kg, just about the same as the standard Celanese fixture, and is only about 70% as expensive. Thus, it has become a very popular alternative to the standard Celanese and IITRI compression fixtures (WTF, 2000).\nAnother modification is the German Modified Celanese Fixture (DIN Standard 65 380, 1991), shown in Figure 7. This fixture uses flat wedges like the IITRI, but circular holders like the Celanese. Unfortunately, it incorporates an alignment sleeve like the Celanese also. This fixture has experienced somewhat limited use to date, primarily in western Europe.", + "5.06.6.4.2 End-loading fixtures\nBy far the most popular end-loading compression test fixture at the present time is the socalled Modified ASTM D 695 compression test fixture. It actually does not conform to that ASTM standard, and is not an ASTM standard in itself. It was developed by the Boeing Company in conjunction with Hercules, Inc. in 1979 (Berg and Adams, 1988), and then included in Boeing Specification Support Standard BSS 7260, first issued in 1982 (Boeing BSS 7260, 1988). Thus, it is often called the Boeing Modified ASTM D 695 compression test fixture. It was later adopted by the Suppliers of Advanced Composite Materials Association in 1989 as SACMA Recommended Method SRM 1\u00b188 (SACMA SRM 1\u00b188).\nThe fixture is shown in Figure 8. It incorporates I-shaped lateral supports like the ASTM D 695 fixture (ASTM D 695, 1996), but that is where the similarity ends. The specimen, rather than being an untabbed dogboned specimen, is straight-sided and tabbed. Actually an untabbed straight-sided specimen is used to measure modulus, and the tabbed specimen for determining compressive strength. The specified 4.8mm gage length between tabs is too short to accommodate a strain gage. The gage\nlength is very short to prevent gross buckling since a specimen only 1mm thick is specified. The untabbed specimen cannot be loaded to failure because it will end-crush prematurely. Having to test two specimens rather than one is inefficient. Also, a complete stress\u00b1strain curve to failure cannot be obtained. Nevertheless, this test method is currently very popular, the fixture being very small and relatively inexpensive, and the strength and modulus results obtained being very comparable to those measured using the shear-loading fixtures (Adams and Lewis, 1991; Westberg and Abdallah, 1987; Adams, 1995). Also, there is no inherent reason why a thicker specimen with a longer gage length cannot be tested in this fixture, as demonstrated by Adams and Lewis (1991) and Westberg and Abdallah (1987).\n5.06.6.4.3 Sandwich specimens\nThe sandwich beam loaded in four-point bending, the face sheet on the compressive side being the test coupon, has already been mentioned. This configuration has been used for many years, at least since the late 1950s, but by just a few groups. General Dynamics Corporation, Fort Worth, Texas, has long" + ] + }, + { + "image_filename": "designv11_65_0003437_0954406042690498-Figure2-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003437_0954406042690498-Figure2-1.png", + "caption": "Fig. 2 Generation of gear and pinion with curvilinear tooth traces", + "texts": [ + " By utilizing Newton\u2019s rootfinding method and computer program, gear designers can efficiently determine the values for parameters of cutting tools to meet a user-defined magnitude of parabolic kinematic error and dimension of the major axis of contact ellipse. With the verification of numerical examples, the goal of designing a robust cylindrical gear set generated by two blade-discs has been fulfilled. To manufacture teeth with symmetric curvilinear tooth traces, the tooth surfaces of two sides of adjacent teeth are generated simultaneously by two blade-discs as shown in Fig. 2, where Fig. 2a represents a gear with convex\u2013convex tooth traces and generated by two bladediscs of which the strips formed by cutting blades are overlapped and Fig. 2b represents a pinion with concave\u2013concave tooth traces and generated by two blade-discs of which the strips formed by cutting blades are crossed. The reason for designing the pinion as C04304 # IMechE 2004 Proc. Instn Mech. Engrs Vol. 218 Part C: J. Mechanical Engineering Science possessing concave\u2013concave tooth traces is to be able to consider reducing the bending stress. To avoid collision of the cutting blades on one blade-disc with those on the other blade-disc, two blade-discs have to rotate synchronously" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002350_robot.1998.680983-Figure5-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002350_robot.1998.680983-Figure5-1.png", + "caption": "Figure 5: Swing states of excitation and suppression", + "texts": [ + " On this orbit the manipulator swings like a pendulum. In this paper we call the orbit an elliptic orbit, and the motion of the manipulator on the elliptic orbit as pendulum motion, for convenience. 2.3 Reference Trajectory of joint 2 The amplitude of pendulum swing motion corresponds to the size of elliptic orbit. In other words, if we excite the amplitude of pendulum swing motion, a transition from one elliptic orbit to a larger orbit occurs. In this section we describe how to change the amplitude of pendulum swing motion. Figure 5 shows counter-clockwise rotational swing. When we move the position of joint 2 forward to the straight line connecting the centers of both joint 1 and the gripper, then pulling the string helps the gripper to accelerate. The swing is thus excited and the swing amplitude becomes larger. On the other hand, when we move the position of joint 2 backward to the straight line connecting the centers of both joint 1 and the gripper, then pulling the string helps the gripper to decelerate. The swing is thus damped and the swing amplitude becomes smaller. In order to realize the excitation and suppression of the pendulum swing shown in Figure 5, we try to change the angle of joint 2 periodically. We use a cycloidal curve for the reference trajectory of joint 2, which is described as follows: 2t 1 4m t, 2n t, 2 &,=Aml{---sin(-)) ( O < t l f l ) 2t 1 4m f. re 2~ re 2 6 Z d =Amz(--+-sin(-)}+2Amz (-3.0.co;2-b-Figure3-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002066_1099-0887(200007)16:7<497::aid-cnm352>3.0.co;2-b-Figure3-1.png", + "caption": "Figure 3. Compressed circular beam.", + "texts": [ + " However, this is only a minor drawback compared to the computational e$ciency of the stress resultant approach. The computing times for the elastic and elasto-plastic cases with the convergence criterion Eqk`1 n`1 !qk n`1 E(TOLEqk#1 n#1!q0 n#1E (19) with TOL\"10~6, were 0.55 and 0.74 s, respectively. In the above formula (19) q is the nodal displacement vector, subscript denotes the step number and superscript denotes the iteration number. 4.2. Compressed ring In the second example we analyse a circular beam loaded by two point forces as shown in Figure 3. Due to symmetry only one-half of the ring is discretized by 40 three-noded elements. The resulting equilibrium paths for elastic and elasto-plastic analyses are depicted in Figure 4. Computing times for the elastic and elasto-plastic cases were 4.75 and 9.85 s, respectively, showing the e$ciency of the stress resultant approach. In Figure 5 some deformed shapes of the ring are shown with both the compressive and oppositely directed forces. The use of stress resultant based elements in fully non-linear structural analysis o" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003702_gt2004-53860-Figure8-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003702_gt2004-53860-Figure8-1.png", + "caption": "Fig. 8. Computer automated foil bearing and damper load deflection test rig", + "texts": [ + " The displacement of the bump foil assembly, due to the applied external load, was measured using two high precision mechanical indicators (precision as high as 50pm). Additionally, in various tests, an eddy current displacement sensor was used to measure inner ring motion to verify the readings taken with the dial indicators. Previous testing of a smaller 150 mm diameter compliant foil damper assembly was completed using a computer controlled and automated load deflection tester 4 Copyright \u00a9 2004 by ASME l=/data/conferences/gt2004/71222/ on 07/07/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use D as shown in Fig. 8. This rig was comprised of a stationary outer base, an inner hydrostatically levitated slide table with a post sized to accommodate the 150 mm diameter damper assembly; a strain gage load cell, an eddy current displacement sensor, a counter weight and an air cylinder to affect the desired range of motion. While this automated tester was not configured for the 216 mm diameter damper, data from the 150 mm damper assemblies were useful in validating the static load deflections of the 216 mm diameter damper taken when installed in the dynamic test rig weldment" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002371_iros.1991.174511-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002371_iros.1991.174511-Figure1-1.png", + "caption": "Figure 1: The local real space normal, sliding tangent, and radius vectors of a generic contact.", + "texts": [ + " It is convenient to choose the reference point of the object at its center of mass (or more generally at its center of compliance), and to represent the object\u2019s motions in terms of generalized coordinates. The generalized coordinates are ( x , y , q ) , viewed as elements of the manifold R2 x Sa. Here S i is the circle of radius p, and p is taken to be the radius of gyration of the object. Thus the relationship between the usual representation of orientation as an angle and the generalized coordinate q is q = /I B . Consider Figure 1, which depicts an abstraction of a planar object in one-point contact with some other object, Two different contacts that could give rise to this same picture are shown in Figure 2. We denote by no the unit real-space normal at the point of contact, and write = (n,.. ny). We let r = (rx, r,.) denote the vector from the point of contact to the moving object\u2019s reference point. Suppose now that we permit the moving object to rotate and translate while maintaining single-point contact with the immobile object. The legal motions of the object thus constrained have two degrees of freedom. We may describe these legal motions as atwo-dimensional surface in the (x , y, q) configuration space of the moving object. At any point on this surface we may construct an outward unit normal to the surface, denoted by n. Referring to Figure 1, we may write with 1, = ( p 2 + ni) ' !2/p, and tig to be determined. It is no coincidence that the first two components of h s vector are directly parallel to the real space normal nu. We can see that this must be so, merely from force considerations. Intuitively, one should think of the configuration space normal n as specifying the direction of a generalized reaction force that arises in response to a frictionless applied force acting on the surface. Consequently, since we have chosen the reference point at the center of mass, and chosen p to be the radius of gyration, ny must simply be the torque induced about the reference point by a unit reaction force at the point of contact: n4 = no x w r where x 2~ is the two-dimensional cross product", + " We can think of friction as acting tangenoally to the physical edge of contact. Let t, denote the unit tangent to the edge of contact. Then t, must be of the Friction acts along this tangent through the point of contact. For a unit frictional reaction force, the induced torque about the center of mass is therefore vq, with vq = t, A D r form t, = *(ny, -nx). Observe that vq = ~ ( 1 1 , rx + ny ry ) . L,et us now write down the equations of motion. Figure 3 depicts a force-body diagram for the contact of Figure 1. Lnt FA = (Fx3 F y . F s ) be a generalized applied force. In other words, the applied Cartesian force is (F,.F,,) and the applied torque is 7 = ,)Fy. This force is applied at the center of mass, that is, at the reference point. Let the normal reaction force at the point of contact have signed magnitude f n r and let the frictional reaction force have signed magnitudef,. We measure f along the outward normal no, and f along the tangent vector t,. For the choice of t , as in Figure 1, the equations of motion are: (m is the mass of the moving object) f n n x + f , n , + F , = ma, f n n y - f , n , + F , = may (1) j n ny + f , vy + 7 = ni p 2 t i , with the restriction 0 5 $,I 5 p f Let us now view the vectors no, t,, and r as elements of !R,3. In other words, they are just like before, but with a thrd coordmate that is zero. Then we can write the configuration space normal n implicitly as A n = no + (no /- r)/,). Clearly .Inn describes the direction of the generalized normal reaction force" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001820_1522-2683(200107)22:12<2606::aid-elps2606>3.0.co;2-i-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001820_1522-2683(200107)22:12<2606::aid-elps2606>3.0.co;2-i-Figure1-1.png", + "caption": "Figure 1. Packing device for the CEC packed column.", + "texts": [ + " The reservoir, containing slurry of the stationary phase in acetone, was sonicated throughout, and methanol was used as the packing solvent. The capillary was typically packed within the first few minutes, after which the pressure was maintained at 400 kg/cm2 for a further 1 h to ensure a firmly packed bed. During the packing, the end of the fused-silica capillary was connected to a Valco union with a metal screen (0.5 m pores) to prevent the material being expelled from the capillary. The packing device is shown as Fig. 1. Frit preparation was according to Yamamoto\u2019s method [27]. CEC columns were conditioned from storage by a low pressure capillary rising device (Fig. 2) with the relevant electrolyte and they were then installed in the instrument for voltage conditioning until a steady current was obtained. CEC experiments were performed with capillaries of total length around 78 cm. The separations were carried out on coupled capillaries connected via a PTFE sleeve. The CEC experiments were performed with a SpectraPHORESIS 100 electrophoresis system (Thermo Separation Products, Fremont, CA, USA) equipped with an UV absorbance detector (Spectra Focus Scanning CE detector) using PC 1000 software V" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001951_s0389-4304(01)00102-3-Figure4-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001951_s0389-4304(01)00102-3-Figure4-1.png", + "caption": "Fig. 4. Ratio change principle.", + "texts": [ + " In addition, the two variators are positioned symmetrically to the center of the output gearset, making it possible to cancel out the opposing thrust forces that act on the discs. Coupling the two input discs via the shaft forms a single thrust generation mechanism that allows equal thrust force to be applied to both variators [2]. 2.3. Ratio change principle of toroidal CVT Transmission ratio changes are accomplished by tilting the power rollers around the axis of rotation of the trunnions that support the rollers; that changes the ratio of the radii of the circles traced by the contact points between the input/output discs and the power rollers (Fig. 4). The tilting of the power rollers is accomplished by applying hydraulic pressure to offset the axis of rotation of the rollers from that of the discs. When the transmission ratio is constant, the axis of rotation of the power rollers and that of the discs intersect in the same plane. As shown in Fig. 5, when a ratio change is executed, the power rollers are offset vertically to produce a difference in rotational speed between them and the discs at their contact points, thereby generating traction force in the tilt direction that causes the power rollers to tilt" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001092_cnm.1630080604-Figure3-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001092_cnm.1630080604-Figure3-1.png", + "caption": "Figure 3. Helicoidal beam: E = 2.1 x 10' kg/cm2, Y = 0", + "texts": [ + " It comes as no surprise that the displacement results for reduced VIBRATION ANALYSIS OF COIL SPRINGS 377 integration (URI) are better than those for full integration (FI). The internal force accuracy is, however, largely dependent on the quadrature rule. This fact is clearly illustrated by comparing transverse shear forces. For URI the transverse shear forces are exact, namely 1 .OOOO, while for FI there is significant oscillation, and even for eight elements the accuracy is poor. The exact displacement was determined by using the exact differential equations for circularly curved beams as in Reference 2. Example 2. Helicoidal beam The helicoidal beam problem (Figure 3) further illustrates the benefits of URI. The beam is subjected to self-weight y = 2-5t/m3. For the analysis, the angle \\I. is constant and chosen as 0-0. Results are presented in Table 11. Displacement convergence is better for URI, and force accuracy is dramatically better. For coarse meshes, results are also better than those of Reference 7 except for MY min. Example 3. Pretwisted beam To illustrate the ability of the element to represent a strongly pretwisted beam, the Tabarrok pretwisted ~an t i l eve r~~\" with shear load at the tip was used (see Figure 4)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000726_cbo9780511530173.007-Figure5.4-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000726_cbo9780511530173.007-Figure5.4-1.png", + "caption": "Figure 5.4 A passive two-parameter spring actuated by a P-P manipulator.", + "texts": [ + "28) (see (5.2)) and make the necessary substitutions by simplifying the results from Chapter 4. Now assume that a wheel is connected to a platform using a two-spring Cambridge Books Online \u00a9 Cambridge University Press, 2009Core terms of use, available at https://www.cambridge.org/core/terms. https://doi.org/10.1017/CBO9780511530173.007 Downloaded from https://www.cambridge.org/core. Columbia University - Law Library, on 13 Aug 2019 at 15:09:03, subject to the Cambridge system, as illustrated in Figure 5.4. The platform is connected to the ground by a serial pair of actuated prismatic joints that are tuned for fine position control. The wheel maintains contact with a rigid wall. From (5.6), [K] is a symmetric 2 X 2 matrix and (5.29) (5.30) where and where [K]-- KJ [Kp] Pi = \u20ac, = [Koi + = \\~Sl i p ov i' [Kpl [*'(1o Ci \"Pi) -s2 c2 , (5.31) Cambridge Books Online \u00a9 Cambridge University Press, 2009Core terms of use, available at https://www.cambridge.org/core/terms. https://doi.org/10.1017/CBO9780511530173", + "34) by computing the proper force error-reducing displacement (or force-correcting displacement) SDC = 8dci i + 8dc2 j . Assume now that the wheel is at rest, but that it is loaded with an exces- Cambridge Books Online \u00a9 Cambridge University Press, 2009Core terms of use, available at https://www.cambridge.org/core/terms. https://doi.org/10.1017/CBO9780511530173.007 Downloaded from https://www.cambridge.org/core. Columbia University - Law Library, on 13 Aug 2019 at 15:09:03, subject to the Cambridge sive normal force parallel to the vector un (see Fig. 5.4). This force, which is reciprocal to the wheel motions, can thus be reduced without moving the wheel provided that _ *f rcos 4 5 i _ *f r \u00b0 - 7 o 7 \" Sf\" [sin 45J - 8f\" [o.7Ov where 8fn is the desired change in the normal force. Substitute this result in (5.34) to give ;H[~ 1 4 1 4 i\"1 So, for this example, 8dc\\ = -0.1414 dfn and 8dc2 = 0. Therefore, a displacement in the negative x direction (for a positive 8fn) reduces a compressive normal force that passes through C A compressive contact force /\u201e is negative and hence a positive 8fn reduces an excessively compressive contact force" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003038_s0094-5765(01)00002-9-Figure5-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003038_s0094-5765(01)00002-9-Figure5-1.png", + "caption": "Fig. 5. Schematic diagram of a multibody system in a chain topology showing coordinate frames and vectors used to deAne an elemental mass dmi on body i.", + "texts": [ + " origin of the frame F2) with respect to the center of mass of the platform. To keep the formulation general, such relative translational motion between adjacent bodies is permitted throughout the chain, although normally it will be present only between bodies 1 and 2. The relative translational motion, when not present, can be eliminated quite readily by the introduction of constraints through Lagrange multipliers. With these introductory comments, consider the chain-type manipulator with N bodies as shown in Fig. 5. As before, body 1 represents the platform, while the remaining bodies (2 to N ) correspond to the manipulator modules. Thus, the second body represents the Arst module of the manipulator, while the body N corresponds to the (N th\u22121) module. Note, the lengths of bodies 2 to N can vary with time. Moreover, each body is free to rotate and translate with respect to its neighbors. As in the case of one module system (Fig. 4), the xi axis is along the length li of the body i; yi is perpendicular to xi in the orbital plane; while zi completes the orthogonal triad", + " The approach here is to evaluate the position vector Rdmi deAning location of the mass element with respect to the inertial frame F0. Its time derivative will give the required velocity. To that end, the body i is isolated from the rest of the system and permitted to translate as well as rotate in three dimensions thus simplifying the calculation of the kinetic energy directly with respect to the inertial frame F0. Later, constraint relations are introduced to establish connections of the body i with the neighboring bodies. Thus two diIerent coordinate systems are used. From Fig. 5, the vector Rdmi to the mass element dmi on the ith body can be written as Rdmi =Di + Ti(ri + i i); (3) where Di refers to the inertial position of the frame Fi; ri=[xi; yi; zi]T is the position vector to dmi with respect to the frame Fi (in absence of deformation of the body i) and fi(ri)= i i is the Oexible deformation at ri. Here i are the admissible shape functions and i represents generalized coordinates. The matrix Ti in eqn (3) denotes a rotational transformation from the body Axed frame Fi to the inertial frame F0, i", + " Thesis, The University of British Columbia, Vancouver, Canada, 1999. 39. Meirovitch, L., Elements of Vibration Analysis, 2nd edn. McGraw-Hill Book Company, New York, U.S.A., 1986, pp. 255\u2013256, 282\u2013290. 40. Pradhan, S., Modi, V. J. and Misra, A. K., A Collection of Technical Papers, AIAA=AAS Astrodynamics Specialist Conference, San Diego, California, U.S.A., July 1996, Paper No. 96-3624, pp. 480\u2013490. APPENDIX Nomenclature di translation vector of the frame Fi from the tip of the (ith \u2212 1) body, Fig. 5 Di inertial position vector to the frame Fi, Fig. 5 dmi mass of the inAnitesimal element located on the ith body, Fig. 5 EAi product of the Young\u2019s modulus of the ith body with its cross-sectional area fi(ri) displacement of the mass element located at ri due to body Oexibility, Fig. 5 F vector containing the terms associated with the centrifugal, Coriolis, gravitational, elastic, and internal dissipative forces, eqn (1) F0 inertial reference frame Fi reference frame attached to the ith body gi position vector to a mass element relative to the frame Fi accounting for deformation of the body i In n\u00d7 n identity matrix K StiIness of the revolute joint li length of the ith body mi mass of the ith body M coupled system mass matrix M\u0303 decoupled system mass matrix M\u0303i decoupled mass matrix of the ith body N number of bodies (i" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000854_s0389-4304(99)00024-7-Figure11-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000854_s0389-4304(99)00024-7-Figure11-1.png", + "caption": "Fig. 11. Reference \"gure for understanding bush friction.", + "texts": [ + " The bottom of the damping force characteristics diagram shows the waveform of the relative displacement during the \"nal 1 s of damping after the vertical up-anddown motion has ended and the technician's hand has been released from the vehicle. The spring constant and wheel load are also output. When spring constants are measured, it is desirable to eliminate the e!ects of static bush friction. To achieve this, the measures described below were added to the measuring process. d Lift up the vehicle body slightly; then release your hand from the body. d In other words, this condition corresponds to point A in Fig. 11. d Press down to cause a shift from point A to point B. As a result, the spring constant can be determined from the gradient between A and B, which nearly eliminates the e!ects of the bush friction. Conversely, the size of the vehicle's bush friction can be determined from the di!erence in values obtained between cases executed from a direction that eliminates the e!ects of such bush friction and cases when the vehicle body is simply pushed downward from the initial condition. 3.3. Vehicle application requiring leverage consideration (such as double-wishbone suspension vehicles) In such cases, application is possible by adding the correction equation described below" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001966_ijmtm.2002.001439-Figure11-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001966_ijmtm.2002.001439-Figure11-1.png", + "caption": "Figure 11 CSV test mode (see online version for colours)", + "texts": [ + " Typically, a corkscrew test produces one- or two-quarter-turn vehicular rolls, that is, 90\u00b0 or 180\u00b0, respectively. Source: Wilson and Gannon (1972). Source: Berg et al. (2002). Vehicular roll and non-roll conditions can be generated using different ramp configurations, such as: ramp length ramp height ramp surfaces: flat vs. spiral longitudinal velocity entrance to the ramp: straight-in vs. curved-in. Therefore, this test mode can provide both roll and non-roll data for rollover sensing algorithm development. The CSV test method, shown in Figure 11, was developed for vehicular roll propensity testing by NHTSA (1991). The CSV test is quasi-static and is relatively easy to perform. In this test mode, the test vehicle is placed at the top of a ramp, which can be adjusted to any angle smaller than the critical static angle of the test vehicle (defined as [90\u00b0 \u2013 tan 1 (2 hcg/T)], where hcg is the vehicular cg height and T, the track width). The wheels of the vehicle sit on \u2018frictionless padding\u2019, which are placed in the guide rails on the ramp. The vehicle slides down the ramp laterally under its own gravity when the ramp angle becomes large enough. Vehicular rollover is initiated when the tires impact the flange located at the bottom of the ramp, as shown in Figure 11. This test mode has been shown to be capable of providing mid-range lateral acceleration levels (5\u20137 g), depending on the ramp angle (for a given ramp length) for both roll and non-roll conditions. Some sensor suppliers and test institutions have adapted this test methodology as a method to generate low lateral g and vehicle roll rates for rollover sensor/algorithm development. However, signals generated from this mode have posed some difficulties to sensor engineers in discriminating rolls from non-roll conditions" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000628_s0001924000066161-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000628_s0001924000066161-Figure1-1.png", + "caption": "Figure 1. Schematic front view of a ram-air parachute deployment (spreading stage). (A) shows the slider initially near the parachute and", + "texts": [ + " This so-called ideal parachute model predicts opening shock and fill ing time, among a wealth of parachute deployment observables, solely based on a parachute's design characteristics and its speed prior to cell inflation. The inflation of (square) ram-air parachutes is very different from that of circular parachutes since each uses different reefing methods*5' 8> 9>. In the case of sport and military personnel ramair parachutes, a slider is used as the principal device to soften open ing shock. By being packed-in at the base of the canopy, the slider initially constrains the parachute into a roughly hemispherical shape (Fig. 1). As each cell of the parachute inflates through ram-air action, the canopy gradually spreads out into its rectangular shape, pushing the slider down the suspension lines in the process. The canopy's inflation rate is controlled by the slider since, by construction, the rate of canopy expansion is determined by the speed of the slider down the suspension lines. In turn, such a speed results from the force balance between the slider's own drag, tension on the suspen sion lines and friction force between the lines and slider; and ultimately, suspension line tension depends on the canopy's own force of drag. This paper focuses on the role of the slider in controlling the infla tion of ram-air parachutes. The deployment stage under considera tion is the so-called wing spreading stage, one during which each parachute cell is inflating and the slider is moving downward as the canopy spreads (mostly) spanwise as sketched in Fig. 1. This partic ular stage is the last of the four stages characterising ram-air para chute deployment and inflation*611). In order of occurrence, the first three stages are: line stretch, when the suspension lines have been fully extended; bag-strip, when the parachute is extracted out of the container bag and remains in a limp state; and initial cell pressurisation, a stage during which the cells are beginning to inflate but to a volume smaller than their design volume. (To clarify the nomencla ture, the first two stages take place during the so-called deployment phase and the last two during the inflation phase", + " For this reason, it is proposed that the second derivative of the drag area be given by the following differential equation l\" = Kv2 ...(2) Here K, or parachute configuration scaling number, is a dimensionless constant which at first sight should depend on packing style, slider and canopy design, suspension line lengths, etc. As shown be low, this equation can be motivated by the fact that the parachute's drag area depends to a large measure on the location of the slider along the suspension lines as shown in Fig. 1, implying that the rate of change of the drag area should depend strongly on the slider's velocity and acceleration. 2.2 A derivation of the configuration number A specific form for the configuration number K can be argued, but not rigorously demonstrated, by using the slider's own Newtonian equa tion of motion. The latter depends on the force of drag generated by the slider, slider weight, suspension lines friction and tension. With respect to a reference frame fixed to the ground, such an equation of motion has the following form _ pslider _ u/ _ 7\u0302 ", + " Finally the force of line friction being generated by the line tension against the slider's grommets, is approximated by Ffrktion~\\LkCosn\u00b1plv2 ...(6) where |it is the coefficient of dynamical friction for nylon rubbing against metal. The next step of the derivation involves rewriting aslider in terms of the parachute's drag area . Considering that today's ram-air chutes feature aspect ratios spanning a 2-3 range, that the chord-wise length of the spreading canopy is approximately constant in time during spreading and that the ratio of the slider-distance-to-canopy (or X) to the span-wise extension of the canopy (Y) (see Fig. 1) is also approximately a constant, one writes 2, = SCD ~ LchardYCD ~ Lchord\u2014 XCD \u2022\u2022\u2022(/) Lline Writing the slider's acceleration with respect to the canopy (or X\") in terms of its acceleration with respect to the ground (aslider), one has -X\" = aslider -v'; now using the fact that the drag coefficient is only changing by a small amount, \u00a3\" ~ ~CDLchord\u2014 \\aslider - v ) . . . (8) Lline combining Equations (8), (6), (5) and (4) into (3) gives y\" ~ ~ chord^pP span f l ^ - ^ l - S i n Q J + ^ C o s a i - ^ ^ - l l v 2 This result shows that \u00a3\"is indeed proportional to the square of the speed as advertised earlier", + " This means that although having a minimal effect on the value of K, packing style will still have a noticable effect on opening shock because of the latter's sensitivity on initial speed as shown below. 2.3 Calculating the value of K The motion of the slider is caused by the sum of large aerodynamic forces acting in opposite directions. According to Equation (9), the value of the configuration number K is determined by the slider's mass and surface area, as well as the parachute's initial drag area (after setting \u00a3 = E0), rigging angle and suspension line friction. The first problem with this formula is that the rigging angle \u00a32 (see Fig. 1) is difficult to measure in the laboratory given the extreme flexibility of a parachute. Moreover, the value of K may depend on dimensionless geometry constants which were not included in the derivation of Equation (9). Finally, although the initial parachute surface area S(t = 0) can be estimated as two to four times that of the slider sur face area, this may not be accurate enough for use in Equation (9). For these reasons, it is unlikely that an accurate value of K will ever be calculated from Equation (9)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002762_s0094-114x(00)00039-2-Figure3-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002762_s0094-114x(00)00039-2-Figure3-1.png", + "caption": "Fig. 3. 10-link mechanism derived from chain No. 118.", + "texts": [ + " Proceeding as before, the successive elimination of variables h7; h5 and h10 can be expressed in condensed form as F1 h6; ~h10 F2 ~h5; h6; h10 F3 h5; h6; ~h7; h10 F4 h5; h6; ~h7; h10 F34 ~h5; h6; h10 9=;F2;34 h6; ~h10 9>=>;R h6 28 The I/O polynomial R 2 k h6 is given as R X26 n 1 an cosn h6 bn cosn\u00ff1 h6 sin h6 a0 X52 n 0 cntn 6 0; 29 where t6 tan h6=2 . Since tangent half-angle substitutions are used repeatedly, Theorem 3 was used to verify that R is devoid of any extraneous solutions. Therefore for a given value of h3; the I/O polynomial is of 52nd degree in t6 and the mechanism of Fig. 3 has 52 possible assembly con\u00aegurations. For the numerical data given below and h3 37:0\u00b0, the coe cients of this 52nd degree polynomial including the 52 possible solutions for t6 are given in Table 4. The three cases considered above in Examples 1 and 2 are the only true 10-link mechanisms which can be obtained from chain No. 118. As discussed earlier, the displacement analysis problem for r1 3:80 r2 8:50 r3 3:25 r4 6:00 r5 3:75 r6 3:00 r7 3:30 r8 1:80 r9 2:15 r10 3:10 r3a 2:65 r3b 4:70 r5a 2:30 r6a 3:75 r6b 2:88 r7a 1:80 r10a 5:95 a1 145:0\u00b0 a2 16:0\u00b0 b 33:0\u00b0 c 49:0\u00b0 / 37:0\u00b0 d 30:0\u00b0 g 51:80 remaining mechanisms (obtained by \u00aexing links 3\u00b110) is no more complex than that of a 8-link mechanism" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003112_iemdc.1999.769044-Figure4-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003112_iemdc.1999.769044-Figure4-1.png", + "caption": "Fig. 4 Flux plot when the motor with solid rotor operates on full load", + "texts": [ + " EXAMPLE The proposed direct coupled FEM modeling has been used to simulate the steady-state operation and dynamic operation of a permanent magnet synchronous motor (2.5 kW / 220 V, 50 Hz, 4 poles, 36 stator slots, A connection, NbFeB magnet for excitation). The time step size is 0.0345 ms. The solution domain of the FEM is one pole pitch. The 2-D FEM mesh of each slice has 3 183 nodes and 5691 elements. The average CPU time to solve the FEM equations at each time step on a Pentium I1 / 266 MHz is 5.5 s. A typical flux plot of the motor with solid rotor on full-load is given in Fig. 4. The transient processes of the electromagnetic torque and the rotor electrical angle (relative to the coordinate system rotating at synchronous speed), when the load torque changes suddenly from 15 Nm to 5 Nm, for a laminated rotor and a solid rotor are shown, respectively, in Figs. 5 ,6 ,7 and 8. One can notice that the motor with a solid rotor has better transient response than the motor with a laminated rotor. V. CONCLUSION The paper presents a direct coupled FEM model of PM synchronous machines which is directly based on the stator circuit equations, FEM equations and torque equation" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002750_bf02326366-Figure3-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002750_bf02326366-Figure3-1.png", + "caption": "Fig. 3--Angular measurements to determine the strain", + "texts": [ + " The components of displacement will then be given by ( ~ l , v ) i ( n , m ) = ( X o __ X i ' y O __ y i ) (1) where u represents the tangential displacement in the direction Ox, v the displacement perpendicular in the direction Oy and i represents the level of the deformation. By using a graphical-differentiation procedure, ' ' the derivatives Ou/Ox and Ov/Oy can be obtained. The component y,y is of particular interest and is proport ional to the shear stress. In the plane of the grid 1 yxy = -~ ( Ou/Oy + Ov/Ox) and O~ry = 21XTxy (2) Figure 3 shows the principle of the measurement. At the point M(n ,o ) of the contact, two angles can be defined: (a) angle 01 made by the tangent t, to the cylinder and the direction Ox; (b) angle 02 made by the direction Oy and the tangent t2 to the line of the grid through point M(n ,o) . The measurement of each of these two angles permits the determination of the derivatives Ou/Oy and Ov/Ox. From angles 0, and 02 in Fig. 3, measured in the positive direction, it follows: tan 0, = Ov/Ox and tan 02 = -Ou/Oy (3) from which the value of %y can be obtained: 27 , v = tan 0, - tan 02 (4) Actually, to obtain a good precision, it is necessary to read the angle 02 several times. The sign of \"7~, can be obtained directly by comparing the angle ( t , , t2) to 7r/2. Results The tests reported here were conducted on a model on which the printed grid had two l ines/mm. The dimensions of the model were 26 x 100 x 120 mm. The cylindrical punch was made of Plexiglas and had a diameter of 40 mm" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003211_6.2003-5754-Figure2.3-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003211_6.2003-5754-Figure2.3-1.png", + "caption": "Fig. 2.3 - Schematic representation of a satellite with two rigid solar panels (drawing not to scale).", + "texts": [ + " In Mooij and Wijnands (2002) a detailed discussion can be found on the evaluation and validation of the GAOCS simulator. As an example of the evaluation process, below an additional test will be described, i.e., related to the perturbing effect of the solar radiation. One dynamic test will be executed. In this case, we put the Sun exactly at the negative X-axis, by hardcoding S = (1,0,0)T, for \u03c6 = \u03b8 = \u03c8 = 0\u00b0. Moreover, two solar panels (L x b = 10 x 1 m) will be attached to the body, as indicated in Fig. 2.3. Note that the surface normal of the solar panels is pre-defined in (local) Z direction. The panels are attached to the central body of the satellite at rP,1 = (0,0.5,-0.1)T m and rP,2 = (0,-0.5,-0.1)T m defined with respect to the geometric centre. The orientation of the panels is such that the local ZP-axis is positive in the direction of the Sun. This means that for solar panel #1, the orientation is \u03c6P,1 = -90\u00b0, \u03b8P,1 = 0\u00b0 and \u03c8P,1 = 90\u00b0, whereas for solar panel #2, the orientation is \u03c6P,2 = 90\u00b0, \u03b8P,2 = 0\u00b0 and \u03c8P,2 = -90\u00b0", + " This can be alleviated by a second extension to the adaptive algorithm, i.e., to apply not only feedforward around the plant but also around the reference model. Note that although the feedforward compensators have been implemented, their use has been postponed till later, once a dedicated effort is done to optimise the MRAC design. A more detailed treatment of theoretical aspects of MRAC can be found in Kaufman et al. (1994), whereas Mooij (2003) gives a recent application. The application of MRAC focuses on a satellite, similar to the one depicted in Fig. 2.3. The related geometry and mass properties are as follows. The satellite body is a parallelepiped with a height of 2.5 m and a square top and bottom cover of 1.8 x 1.8 m. The mass of this body is 1,600 kg, assumed to be uniformly distributed, with a corresponding inertia tensor of Ib = diag(1265,1265,864) kgm2. The two solar panels are represented by stiff rectangular flat plates of 9x1.2 m, each having a mass of 40 kg and an inertia tensor of Ip = diag(5,270,275) kgm2, referenced to the panel c" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001124_02783649922066574-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001124_02783649922066574-Figure1-1.png", + "caption": "Fig. 1. A simulated task involving the positioning of a large object.", + "texts": [ + " A changing definition of those camera-space targets will ensure that the manipulated body gradually approaches the target body with a specified, geometrically workable trajectory, until closure is obtained. The determination of the angular joint rotations is performed by minimizing the following scalar function \u03b3 over the angular joint coordinates of the robot, 2: \u03b3 (2)= nc\u2211 i=1 nt (i)\u2211 j=1 { [ xt i j \u2212 fx ( rx i j (2), ry i j (2), rz i j (2); Ci )]2 + [ yt i j \u2212 fy ( rx i j (2), ry i j (2), rz i j (2); Ci )]2 } Wi j , (1) for nc cameras and nt (i) target points in the ith of nc cameras. See Figure 1 for an illustration of target points in a simulated task. Unlike past experiments where all cameras pointed to the same region, this paper introduces an innovation of different cameras pointing to different regions to ensure precision when handling a large manipulable object. The j th target point for the ith camera is defined as (xi tj , y i tj ), and the corresponding physical location is desig- nated as (ri xj , r i yj , r i zj ), which depends on the angular joint configuration of the manipulator, included in 2, according to the nominal forward-kinematic model", + " In general, if the desired three-dimensional terminal location of the target points is defined with respect to a coordinate system fixed on the stationary nonmanipulable body, the intermediate three-dimensional positions of the target points, which constitute the approach trajectory, can be defined by means of a 4 \u00d7 4 transformation matrix Di for the ith camera as follows: t \u2032ixj t \u2032iyj t \u2032izj 1 = Di t ixj t iyj t izj 1 (5) Primed coordinates (t \u2032ixj , t \u2032i yj , t \u2032i zj ) represent the location of the j th target point at an intermediate juncture that is part of the approach trajectory, whereas (t ixj , t i yj , t i zj ) represent the desired location of the same target point as defined with respect to a coordinate system attached to the nonmanipulable body (see Fig. 7). Matrix D has been used in the past (Skaar, Brockman, and Jang 1990) to define the trajectory of the manipulated body toward the workpiece. The present paper introduces a complementary way of defining such an approach at UNIV OF ILLINOIS URBANA on March 10, 2015ijr.sagepub.comDownloaded from trajectory using another 4 \u00d7 4 matrix, called 0, as explained below. Consider now Figure 1, which depicts the general case in which a large object is moved by a robot. As explained above, several cameras are pointed toward different regions of the workspace. If a coordinate system is to be attached to the nonmanipulable body in each of the regions observed by the cameras, then the number of separate coordinate systems will be limited by the number of cameras used to perform the task. The corresponding camera-space targets are evaluated by using the model equations defined in eqs. (2) and (3) and the view parameters included in C i as follows: xt i j = b1 i t \u2032x i j + b2 i t \u2032y i j + b3 i t \u2032z i j + b4 i , (6) yt i j = b5 i t \u2032x i j + b6 i t \u2032y i j + b7 i t \u2032z i j + b8 i , (7) where b1 i , b2 i , " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003918_j.jmatprotec.2005.02.163-Figure8-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003918_j.jmatprotec.2005.02.163-Figure8-1.png", + "caption": "Fig. 8. Sample plugs and related inlets.", + "texts": [ + " If everything works fine and the control program is finished, all plugs can be removed. By turning off the main source, the cylinder opens the fixation and the system turns back into its initial state, as shown in Fig. 6. The network can be removed by the operator. ombined switch and light is figuratively represented by the rawing of the lamp and the switch P LOCK. If the testing device is in use, the plug is pressed into the nlet and the switching contact P LOCK is closed by the plug. In a complex car network exists many different kinds of plugs, Fig. 8 shows only a few examples. First of all the new solution should be suitable for all different kinds and shapes of plugs. It should be able to fix the plug with or without the air tightness test. The required space for the testing element should be standardised to the size of the front platform in Fig. 5. The available testing device needs twice as much of space, as shown in a later chapter in Fig. 12. To achieve the aim of universal usage it is not possible to be dependent on the special shape of each plug", + " For example, it would be possible to fix the plug with bigger shapes by using a kind of balloon, which increases its volume inside of the plug. Another concern is to reduce the problem of fixation to a simple requirement: a solid body with one edge has to be fixed somehow, as shown in Fig. 9. The size of the plug does not matter because if it is required, more than one element could be used to fix it. If this goal would we realised, the f If the plug is fixed inside of the inlet, the only additional task to do is attach contacts to test the connections. Three different kinds of inlets are shown in Fig. 8. The air tightness test is solved in the best and cheapest way. Something has to get into the empty space inside of the plug and air is readily available and cheap as well. Other mediums would possibly leave something behind. Therefore this system requires no improvement and remains in place. Some plugs were given by the company to get an idea of their variety. Fig. 10 provides a wide range of shapes and sizes. To achieve the aim of universality, the biggest plug with the dimension of 39 mm \u00d7 22 mm \u00d7 22 mm was chosen" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002341_cdc.1990.203964-Figure2-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002341_cdc.1990.203964-Figure2-1.png", + "caption": "Figure 2: Finger i in contact with object.", + "texts": [ + " The expression relating the surface parameter U E 92' to the position/orientation of the object, is given by Define the object, velocity by Y e [voT wOTlT, where v, = &(t) and w, x R, = R, . Also note that there exists a matrix PI() such that wo = PI(^,)+^ Thus for X , = [zf y f I T there exist some matrix P satisfying Y = PX, . Differentiate equation 8 to obtain the following constraint equation between L and Y . J'L = UY (9) where we define Note that U is a non-singular square matrix. We now consider contact between the robot hand and the object. Consider a hand of m fingers, each finger making contact with the object through a point contact with friction as shown in figure 2. Fix a coordinate frame C f i to the object at the contact point with finger i such that its z-axis is aligned with the last link of finger i. Let the orientation of the frame C f i relative to the object frame CO be specified by Ri(t) E SO(3). Suppose the m finger contact points on the object have coordinates c1,. . ., c, respectively relative to the object frame Co. The velocity w; of the contact point i, is given in terms of the object velocity Y by vi = U; [ ;] f O T i = l ...m (10) where Vi = [ I - (Ri ( t )c i )x ] E Px6", + " Define FC, the friction cone for the hand, to be the Cartesian product of all the individual friction cones at each of the contact points. We say that a grasp is force closure if G maps FC onto @. See [19] and [18]. Each finger is a manipulator with a specific forward kinematic map. Suppose finger i has ni joints with generalized coordinates denoted by q; E Pi. For finger i we define a finger Jacobian matrix Ji E vXn*. This matrix relates the joint velocities ( q i E Rn*) to the translational velocity of the fingertips (vi E p). Since finger i maintains contact with the object at the point ci shown in Figure 2, it follows that the fingertip velocity is equal to 'U; given by equation 10. Thus vi = J;q; f o r i = 1.. .m. (12) Substituting each of these m equations into equation 11, we obtain an expression relating the joint velocities to the object velocity of the form Jq = GTY (13) where J = Block diag[Ji . . . Jm] E p m x m n i and q = [qr . . . qZlT. We will use fingers of three joints each (i.e. n; = 3) so that each finger Jacobian matrix is square. 3.2 System Force Relations It is our d m to provide a dynamic control law for the hand so it is necessary to determine what forces act on the system to produce the desired object motion" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001065_(sici)1099-1115(199812)12:8<623::aid-acs517>3.0.co;2-2-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001065_(sici)1099-1115(199812)12:8<623::aid-acs517>3.0.co;2-2-Figure1-1.png", + "caption": "Figure 1. Reference systems, ship motions and angle of incidence of disturbances", + "texts": [ + " Section 4 presents the solution for implementation problems including adaption, slew-rate limiting and integral wind-up. In Section 5 results are presented and analysed and conclusions are drawn in Section 6. The motions of a ship can be separated into two categories, namely, symmetrical (pitch, surge, and heave) and asymmetrical (roll, yaw and sway) motions. The different categories can be assumed to be uncoupled but within one category the motions are coupled. All motions will be referred to the body axes of the vessel, as shown in Figure 1. A second earth-fixed Cartesian axes may be used to describe the vehicles position and orientation. The dynamic equations of motion of a marine vehicle may be described by considering the ship as a rigid body which translates and rotates about three body-fixed Cartesian axes, shown in Figure 1. The body-fixed axes is free to translate and rotate in the earth-fixed axes. This enables all the six-degrees of ship motion to be described. Two non-linear ship models were used in this work. The first one corresponds to the model of a supply vessel with 68)88 m length.4 The second one corresponds to a 171)80 m length Mariner class vessel.5 The models consist only of the ships\u2019 surface motions (surge, sway and yaw). The Int. J. Adapt. Control Signal Process. 12, 623\u2014648 (1998)( 1998 John Wiley & Sons, Ltd" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003864_50002-2-Figure2.6-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003864_50002-2-Figure2.6-1.png", + "caption": "Fig. 2.6. Rotational slip resulting from path curvature and wheel camber (slip angle = 0).", + "texts": [ + " In the normal case of an approximately horizontal road surface, the wheel speed of revolution ~ may be defined as the angular speed of the wheel body (rim) seen with respect to a vertical plane that passes through the wheel spindle axis. On a flat level road, the angular speeA of rolling ,O r and the speed of revolution of the wheel s are equal to each other. The absolute speed of rotation of the wheel about the spindle axis co, will be different from -s when the wheel is cambered and a yaw rate occurs of the plane through the spindle axis and normal to the road about the normal to the road. Then (cf. Fig. 2.6) co - - ~ + ~b siny (2.7) This equation forms a correct basis for a general definition of O also on nonlevel road surfaces. Its computation is straight forward if o9, is available from wheel dynamics calculations. The longitudinal running speed V~x is defined as the longitudinal component of the velocity of propagation of the imaginary point C* (on radius vector r) in the direction of the x-axis (vector l). In case the wheel is moved in such a way that the same point remains in contact with the road we would have Vcx = V,x", + " The definitions of the slip components then reduce to: V x - - sx (2.11) V c x V t a n a - - sy (2.12) V CX The slip velocities V~x and V~y form the components of the slip speed vector V~ and x and tana the components of the slip vector s~. We have: / sx] Vs- gsy ) and s s - (2.14) t a n a The 'spin' slip (p is defined as the component -to z of the absolute speed of rotation vector o9 of the wheel body along the normal to the road plane n divided by the forward running speed. We obtain the expression in terms of yaw rate and camber angle ~, (cf. Fig.2.6): % ~b - t2 sin?, Vcx v* c x (2.15) The minus sign is introduced again to remain consistent with the definitions of longitudinal and lateral slip (2.11, 2.12). Then, we will have as a result of a positive ~o a positive moment M z. It turns out that then also the resulting side force Fy is positive. The yaw rate ~ is defined as the speed of rotation of the line of intersection (unit vector l) about the z axis normal to the road (cf. Fig.2.3). If side slip does not occur (a --- 0) and the wheel moves over a flat road, equation (2" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002669_fuzzy.1995.410052-Figure4-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002669_fuzzy.1995.410052-Figure4-1.png", + "caption": "Fig. 4 Optimization of Combination of Angles, w.", + "texts": [ + " GA uses population of individuals and genetic operations of selection, crossover and mutation. CA advances by stages of generation and evaluates the population to select good individuals at each generation. The evaluation function is used for this selection as a fitness function. The less evaluation value, the more the individual suits to the problem. Each individual is a string of deviation from the initial rotational angle w of the tool. The initial value is generated at random. In the case that there are n points along a path as in Fig. 4, the individual is expressed as follows: dHi ) is the deviation at i-th point. The population is m. The initial population is generated two ways: 1) Each d H i ) is generated at random where IdH i ) lS a. 2) Every \"genes\" are the same (0, -p, or p). Selection, crossover, and mutation operations are used to produce new individuals. In the selection, some of the population is selected while using the evaluation function obtained by the GMDH. In the crossover operation, two individuals are randomly chosen from the selected population at random and one locus of a string is also chosen at random" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001066_bf00055000-Figure7-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001066_bf00055000-Figure7-1.png", + "caption": "Figure 7. The equations constraining the r(w) variables are stated assuming, for simplicity of notation, that walks comprising two of the segments of a Y- or E-junctions are oriented so as to leave the third segment on its left side as in (A). Figure (B) and (C) clarify the meaning of the r(w) variable: In both cases, toj is paneled, but r(wj) ~. 1 in (B), whereas r(w]) = 0 in (C).", + "texts": [ + " w~.} of paneled walks. The following proposition has therefore been proved: Proposition 3. There exists a polynomial time algorithm which determines whether a line drawing \u00a3 is panelable and finds a legal paneling i f it exists. See Section 4.5 for examples of paneling. 4.3. Labeling a Paneled Line Drawing Once a legal paneling for \u00a3 is found, the labeling task is greatly simplified. Once a walk is known to be paneled, only the side on which the panel lies is unknown (see for example walk wj in Fig. 7. Once this is determined for all panels, a unique labeling for the line drawing is automatically available. Let (/2, P) be a paneled line drawing. Let W' be the set of paneled walks associated with (\u00a3, P) . A boolean 152 Parodi variable r(w') is associated with every element w' of W' such that, given an arbitrary orientation for the walk, r(w') = / 1 if w' is paneled on the right side if w' is paneled on the left side / U The variables thus defined are not independent of one another. The constraints come from the junctions. Let us assume, for simplicity of notation, that in the case of 3-degree junctions a walk comprising two of the three segments of the junction under examination leaves the third segment on its left side, as in Fig. 7(A). I f this is not the case for walk w, it is enough to replace r(w) with r(w). These are the constraints on paneled junctions: 1. Y(-)-junctions. At most one panel can lie on the side containing the third segment (as wj in Fig. 7(C)). This means that, if all walks 1/3i, W j , W k associated with the junction are paneled (P(wi) = P(wj ) = P(wk) = 1) then it must be (r(wi) + r(wj))(r(wj) -t- r(tok)) (r(tOk) -I- r(wi)) : 1 If for example only wi and wj are paneled then it must only be r(wi) q- r(wj) = 1 2. Y( +)-junctions. No panel can lie on the side containing the third segment. Then if all walks wi, wj, Wk associated with the junction are paneled ( P(wi) = P(wj) = P(wk) = 1) it must be r(wi) : r(wj) = r(wk) = 1 One of these three conditions disappears whenever one of the three walks is not paneled" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002448_6.1989-1202-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002448_6.1989-1202-Figure1-1.png", + "caption": "Fig. 1: Single Beam Example", + "texts": [ + " Using a similar argument as above, it is easy to see that since the partial velocity has to be c o m t to first order in the generalized coordinates and speeds, and again this term is obtained through differentiation of the velocity with respect to the generalized speeds, the velocity has to be correct to second order in qi and ui until we form the partial velocities. This is necessary if we want our equations of motion to be consistently linear in the generalized coordinates and speeds. To conclude this section we present a simple yet very important example which will help to point out the practical consequences of the above discussion. Consider a simple uniform beam cantilevered to a solid body free to rotate in the plane but pinned at point C which is also its mass center (see Fig. 1). The frame N, defined by the unit vectors nl, n2, n3, is mertial, and we introduce the rotating frame A defined by the unit vectors al, a2, a3, attached to body A and whose a1 axis lies along the undeformed neutral axis of the beam B initially. Let us derive the consistently linearized equations of motion using both Kanc's and Lagrange's methods. We will ignore shear and rotary inertia effects (i.e., slender beam assumption). For simplicity, we assume no external forces act on the system. 2.1 Laerange's of motion: First we form the kinetic and potential energies of our system: where IA is the moment of inertia of the rigid body B about its mass center; vP1 is the velocity with respect to the inertial frame of the centroid of a rigid differential element P1 of the beam; p and El are the mass per unit length and bending stiffness of the beam respectively; and, ul(x,t) and u2(x,t) are the axial and transverse displacements respectively of the element P1 a distance x along a1 from point 0", + " The \"inconsistent\" model shall be that obtained through linear kinematics of deformation, that is, one whose equations omit the superscript 2 terms mentioned above. Finally, a \"ruthlessly linearized,\" or simply \"ruthless\" model, shall be one in which the non-linear terms which include elastic coordinates and speeds are ignored, including those terms in the mass matrix which depend on elastic coordinates. In other words, in our \"ruthless\" model equations we ignore terms bij and &,i, and we assume the mass matrix depends on rigid body configuration only. Consider a simple, slender, uniform beam cantilevered to a rigid base (see Fig. 1). This time we allow the base to move freely in the plane, but we again assume small elastic deflections. We have used Kane's dynamical equations together with non-linear strain-displacement relations to obtain the following equations of motion, exact to first-order in generalized elastic coordinates and speeds [24] : In deriving the above equations, we have again assumed no external forces act on the system for simplicity. m~ is the mass of the rigid body A; ul and uz are translational speeds of body A in the directions of a1 and a2 respectively; us is the rotational speed of body A; qi are the n generalized elastic coordinates, where we have proceeded as in section 2" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001578_ao.41.002879-Figure11-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001578_ao.41.002879-Figure11-1.png", + "caption": "Fig. 11. Schematic of the LED-based prototype sensor system using a single-reflection configuration. The apertures A in front of the LED and photodetector allow only the light within the angular width of 8\u00b0 to interrogate the sensing layer and be detected.", + "texts": [ + "7\u00b0 . Optimized Configuration The experiment discussed in Section 3 demonstrates that, for the sensing configuration used, it is possible to achieve a significant optimization of sensor response in a single-reflection configuration. Furthermore, as mentioned in Subection 2.C, functions R 20 May 2002 Vol. 41, No. 15 APPLIED OPTICS 2885 and S\u0303 do not change significantly if an LED is used as a light source. Consequently a prototype sensor unit incorporating these features was designed, as shown in Fig. 11. The LED was chosen so as to match the absorption spectrum of the BCP-doped solgel film that was prepared as described in Subsection 3.A. The emission spectrum of the LED is shown in the inset of Fig. 7. The emission maximum, located at max 570 nm, corresponds closely to the absorption peak of the BCP-doped solgel film that is located around 580 nm. The core feature of the prototype sensor unit is a metal head that provides rigid support for the LED and Si photodiode. Orientations of the LED and detector are selected to ensure that light interacts with the sensing layer at the optimum incident angle. Furthermore the dimensions of the LED and the detector apertures see Fig. 11 were designed so as to allow only light within a narrow angular width the full angular width was approximately 8\u00b0 to interrogate the sensing layer and be detected by the detector. The orientations of both the LED and the detector were chosen so as to ensure that the center angle of interrogation corresponded to the point of maximum sensitivity obtained from Fig. 10 b , i.e., max 66\u00b0. It can be seen from Fig. 10 b that the light of this angular width is not incident solely in the region of optimum sensitivity but also in a region where sensitivity decreases to approximately half of its maximum value" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001198_0022-0728(96)04535-4-Figure3-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001198_0022-0728(96)04535-4-Figure3-1.png", + "caption": "Fig. 3. Schematic outline of the membrane-covered electrode used in the experiments.", + "texts": [ + " 2, which shows the decay of the current density j (so that we can use the same scale for different cathode sizes) over the first half-second after switch-on. The various parameter values used in calculating these theoretical curves are given in Table 1 assuming an aqueous electrolyte. Although not shown, all electrode sizes give the same initial current density, the smaller cathodes decaying less rapidly as oxygen is drawn in radially from the more diffusive electrolyte layer, whilst the larger cathodes deplete the electrolyte of oxygen very All studies were conducted with a sensor constructed like an ordinary Clark Oxygen electrode, as shown in Fig. 3 and described by us previously [25]. The working electrode was either a 10 p.m diameter gold microdisc electrode sealed in a glass rod (EG&G Princetown Applied Research Corporation, Princetown, USA) with a diameter of 4 mm and a length of 84 mm, or an 80 /xm diameter gold microdisc electrode, also sealed in a glass rod of 5.4 mm diameter and of length 82 mm, and fabricated at La Trobe University, Australia [26]. The diameter of each working electrode was ascertained by the oxidation of ferrocene in acetontrile, again as described previously [27]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002980_00004356-200109000-00012-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002980_00004356-200109000-00012-Figure1-1.png", + "caption": "Fig. 1. The arm-lever propelled three-wheeled chair. A, arm lever; B, brake lever; C, crank rod; D, front wheel; E, rear wheel; F, footboard.", + "texts": [ + " Little information exists regarding evaluation of performance of ALpropelled chairs using cardiorespiratory parameters in practical situations encountered by the users in their day-to-day activities. The purpose of this study was to evaluate the physiological stress and locomotor performance of the ALPTWC in actual locomotive conditions by regular users, which may be helpful for recommendations in clinical practice. Address for correspondence: Department of Occupational Health, All India Institute of Hygiene and Public Health, 110 Chittaranjan Avenue, Calcutta-700073, India. The ALPTWC is a tricycle modified into an arm\u017d .lever propelled chair Fig. 1 , built on a rigid frame which is based on a traditional \u2018crank-to-rod\u2019 mechanism. It consists of three wheels, one in front and two at the rear. The two arm-levers are mounted at each side so that the handles of the operating lever arm remain at the shoulder level of the user in a seated position. A connecting rod is attached to two levers and its other end is pinned to the crank bar, which drives the rear wheel. The operation of the lever through this mechanism creates a torque on the rear wheel that provides the rotation for starting and running of the vehicle" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002623_a:1015265514820-Figure10-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002623_a:1015265514820-Figure10-1.png", + "caption": "Figure 10. The placement of the ultrasonics at the front right side of the wheelchair. The figure shows three sonars at 90\u25e6 (sensor n6 in Fig. 8), 65\u25e6 (sensor n5 in Fig. 8) and 25\u25e6 (sensor n4 in Fig. 8).", + "texts": [ + " An aluminum frame that surrounded the chair was used for mounting all the proximity sensors. Five of the ultrasonics (p1\u20133, s1\u20132) worked in a fail-safe mode, thus increasing the reliability of the sensor readings in these directions. The absolute orientation of the robot was extracted using a pair of encoders attached to the inner side of each of the two back wheels of the chair. The encoder outputs were processed by an encoder card installed in the PC. The wheelchair moved using the two back wheels, while the two front wheels were passive (Fig. 10). The existence of two front castor wheels introduced a large degree of uncertainty during the interpretation of the motion commands by the robot controller. For example, it was not possible for the wheelchair to move straight, immediately after a sharp turn. What usually happened in this case was that the wheelchaor started swerving left and right until bot castor wheels achieved the desired forward motion direction. Then the wheelchair started moving in a straight direction. This behavior was potentially dangerous, especially when the robot moved in cluttered spaces that included narrow corridors or stairs" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000726_cbo9780511530173.007-Figure5.1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000726_cbo9780511530173.007-Figure5.1-1.png", + "caption": "Figure 5.1 A planar compliant coupling.", + "texts": [ + "he stiffness mapping for a parallel manipulator Figure 5.1 illustrates an elastically compliant, planar parallel manipulator. The moving and fixed platforms are connected by three RPR serial chains and in each prismatic pair there is a linear spring. Assume that the moving platform is in equilibrium with an externally applied force with coordinates w and magnitude/is applied to it on a line $. Then w = f\\s\\ + fih + hh> (eq. 2-60) where/i,/2, and/3 are the magnitudes of the resultant forces in the connectors and \u00a31, \u00a32, and \u00a33 are the line coordinates of the connectors", + "org/core. Columbia University - Law Library, on 13 Aug 2019 at 15:09:03, subject to the Cambridge Now dim 8D = 1 Therefore, from (5.18) and (5.19): dim ([K] 8D) = FL = dim 8w. The dimensions of (5.18) are thus consistent. (5.19) (5.20) Figure 5.3 is a schematic representation of a pair of RPR serial connectors, with springs in the prismatic joints, and connections to the ground at points B\\ and B2. The free ends are connected to a common turning joint at a point C. This system can be obtained from Figure 5.1 by shrinking the movable platform to a point C and removing the third RPR connector. Equation 5.5 can then be expressed by Sf = [K] SD, (5.21) where Cambridge Books Online \u00a9 Cambridge University Press, 2009Core terms of use, available at https://www.cambridge.org/core/terms. https://doi.org/10.1017/CBO9780511530173.007 Downloaded from https://www.cambridge.org/core. Columbia University - Law Library, on 13 Aug 2019 at 15:09:03, subject to the Cambridge is a small increment of force acting at point C and \u2022ra is a small displacement of point C", + " Stability phenomena that result in sudden erratic behavior can be explained by a mathematical theory called catastrophe theory (see Zeeman 1977, Arnol'd 1992, and Bruce and Giblin 1993). Catastrophe theory is currently being applied to the two-spring system by R. Hines (a research assistant at the University of Florida), D. Marsh (a lecturer in mathematics at Napier University, Edinburgh), and the author. As far as the author is aware there is no study of the stability of the planarthree-spring system in progress. A relevant paper on the planar three-spring system was reported by Griffis and Duffy (1992). EXERCISE 5.1 The two-spring system (see Exr. Fig. 5.1) is in its unloaded configuration. A vertical force F is applied at C Plot the locus of the equilibrium posi- Cambridge Books Online \u00a9 Cambridge University Press, 2009Core terms of use, available at https://www.cambridge.org/core/terms. https://doi.org/10.1017/CBO9780511530173.007 Downloaded from https://www.cambridge.org/core. Columbia University - Law Library, on 13 Aug 2019 at 15:09:03, subject to the Cambridge B2 2 iiin. Bj Q2 = 60 /f02 = 2 in kt=2 Ibf/in f01 =3.46 in kt=2 Ibf/in rFlbf Exercise Figure 5.1 tions for C when the magnitude of the force increases in increments of 0.25 lbf in the range 0 < F < 5 lbf. (a) Assume that the spring matrix is given by c2~\\ \\ki Ol |~ci s{\\ s2\\ [0 k2\\ [c2 s2j (b) Assume that the spring matrix is given by \u2014s\\ \u2014s2 | | /ci(l \u2014 pi) 0 [K] = [Ko] + c\\ c2 0 - P 2 > \\ \\ - s} Cambridge Books Online \u00a9 Cambridge University Press, 2009Core terms of use, available at https://www.cambridge.org/core/terms. https://doi.org/10.1017/CBO9780511530173.007 Downloaded from https://www" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003663_robot.2003.1242186-Figure3-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003663_robot.2003.1242186-Figure3-1.png", + "caption": "Figure 3: Dynamic model of the planar quadruped showing two of the four legs simulated. A compliant leg is simulated when the leg contacts the ground.", + "texts": [ + " (Illustration from [lo].) per stride (hereafter, \"cost oftransport,\" or COT) and the peak forces experienced in the legs The complete specifications for the optimization problem are given in [6]. Due to space limitations, we will assume here that the optimization problem for the body trajectory has been solved, leaving the task of finding periodic leg trajectories, described below. Before discussing this, however, we will first present the dynamic model for the simulation. 4 The Dynamic Model Figure 3 shows the dynamic model of the planar quadruped. Parameters can he found in 161 and are based in large part on 141. The quadruped has a t e tal m a s of 61.3kg, of which 4 6 . 3 k g is comprised of the body. The center of mass is situated towards the front of the machine, similar to quadrupedal mammals, such that d f = 0.25m and d, = 0.57m. A planar model was selected because the velocity component in the direction perpendicular to the plane of progression is typically small for quadrupedal running gaits [Ill" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003684_cdc.2004.1428835-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003684_cdc.2004.1428835-Figure1-1.png", + "caption": "Fig. 1. Kinematic chain model of a limb.", + "texts": [ + " Therefore, \u03a6\u0302k drives the system with parameters \u03a8k, \u03c8k to the configuration [qk, q\u0307k]. Notice that, even if the number of the primitives has to be noticeably increased (from O(n) to O(n4)), the control paradigm has gained in terms of robustness. Different kinematic structures can be controlled to the desired configuration with simple modifications of the time invariant combinators. With the intent of imitating human reaching motions, we have considered a 2DOF (two degrees of freedom) planar model of a limb (see Fig. 1) similar to the one used in [1] and [2]. The dynamics of this model can be expressed in the form (3) with n = 2: let the input be the torques applied at the joints, u = [u1 u2], and the output be the cartesian position of the extremity P , i.e. y = [xP yP ] . The input to output feedback linearization of this model leads to y\u0308 = v. We have then synthesized a set of motion primitives \u03a6k using results exposed in Section IV. In a first simulation, the weighting matrices Q and R have been chosen on the basis of the minimum energy paradigm proposed in [14]: thus Q1 = Q2 = Q3 = On\u00d7n and R = Im" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003242_tencon.1993.320527-Figure4-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003242_tencon.1993.320527-Figure4-1.png", + "caption": "Figure 4 The K , T, and J ( ? ) contour.", + "texts": [ + " The direct contributory factors in the control efforts for a PI regulator is the proportional gain and the reset action. As has been highlighted earlier, the reset action for the given PI control law is a function of K and T , , see equation E9. Therefore, the gradient method should more ap ropriately be expressed in terms of the vector [ K - J ( e ) I , instead of the one T, suggested earlier. The cross coupling of K and T, in the reset attion can produce drastic localised convergence when a local minima exists near the global one, see Figure 4. Changes in K and T, can result in the reset action value reaching a critical point which becomes mapped onto the K, T, and J ( 0 ) plane as a critical ridge, since there is an infinite ccinbination of K and T, values that could be mapped onto a given critical value of the reset action. R Hence, when the gain and the reset ?me values are changed at an mconipatible rate causing the ratio to reach the critical value, i t can produce local convergence of the parameters. This phenomena was investigated on the temperature loop adaptation" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000527_60.136226-Figure2-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000527_60.136226-Figure2-1.png", + "caption": "Figure 2: Field distribution when only the main main winding is excited", + "texts": [ + "rra) 2 r a Therefore, the end winding inductance per pole will be : a + r 2 a a + r l a r r r After carrying out integrations and multiplying by 4 for the four poles, we obtain: I 0.5(al+ a2+2ra) Substituting numerical values in equation ( 5 ) , ( N = al = 22 mm, r = 5 mm), we obtain : Lae= 7.26423 mH b) Zig-zag leakage inductance, a2 = 41 mm, Lazz : ( 5 ) 254, The zig-zag leakage flux is the one which crosses the air-gap, but does not link the rotor bars or end rings. The path of this flux is shown in Figure 2 as the flux passing between the points P - P and P - P 1 2 3 4' The zig-zag flux forms a good part of the total leakage flux and therefore, it has a strong effect on the performance of the machine. Having investigated the correct path of the zig-zag flux, this flux is calculated as the product of the flux density in each element on the flux path and the area perpendicular to the flux path. The flux density on each element is obtained using Finite Element method. Therefore, zig-zag flux linkage is calculated as: 309 (6) were B 1 2 and B are the flux densities over the triangles between the points P 1 2 - P and P3- P4, respectively (see Figure 21, and hl and h2 are the lengths perpendicular to the zig-zag flux paths" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002336_cdc.1996.573729-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002336_cdc.1996.573729-Figure1-1.png", + "caption": "Figure 1: the HILARE robot", + "texts": [ + " S ( t 0 ) = q t f ) = 0 'Here, ri means $ for e being a scalar or a vector. 0-7803-3590-2/96 $5.00 Q 1996 IEEE 3587 The approach we present applies t o a great number of mechanical systems, and particularly to nonholonomic mechanical systems, and it has already been considered for robotic manipulators [ll, 121. We present the general methodology for nonholonomic systems in the broad outline but all the computations and examples are done for the HILARE mobile robot. The HILARE robot is sketched in the figure 1. The configuration is defined by q = (q1 q2 9 ~ ) ~ = (x y in R3. This robot has two driven wheels at the rear and a castor at the front. It is submitted to the nonholonomic rolling without slippzng constraint. Concerning dynamics, we consider that the torques are bounded in the same way for both driven wheels and that angular acceleration of each of them is directly obtained from the torque by multiplying by a scaling factor. First, admissibility of paths is characterized from the nonholonomic constraints" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003476_jrs.1250090313-Figure2-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003476_jrs.1250090313-Figure2-1.png", + "caption": "Figure 2. (a) Microcellule a capillaire et faces planes. (b) Microcellule a capillaire. (c) Microcellule a quatre faces planes.", + "texts": [ + " Les cellules que nous avons mises au point rtpondent 1 trois impiratifs: elles ont quatre faces situCes sur deux directions peryendiculaires, elles sont de f aibles volumes et elles sont 1 circulation pour Cviter l\u2019kchauffement de 1\u2019Cchantillon par le faisceau laser. Trois types de cellules ont C t C envisagkes: Un tube capillaire de 1 mm de diamktre interieur disposC dans le S ~ I I S du faisceau laser et dont les extrEmitCs sont fermkes par deux lames B faces parallkles. L\u2019CtanchCitk est assurCe par deux joints permettant la circulation (Fig. 2(a)) et son volume est infCrieur 1 10 ~ 1 . Un tube capillaire disposC perpendiculairement au plan form6 par les directions d\u2019Cclairement et d\u2019extraction de la lumikre. Le diamktre intkrieur du tube est 2 mm. I1 est suffisant pour que la trace du faisceau laser dans I\u2019kchantillon couvre la totalit6 de la fente d\u2019entrCe du spectromktre. Son volume est d\u2019environ 15 ~ 1 . La circulation se fait en raccordant deux tubulures en tCflon aux extrtmites du capillaire (Fig. 2(b)). Une cellule 1 quatre faces planes dCmontables, d\u2019un volume de 15 pl (Fig. 2(c)). Cette cuve prCsente l\u2019avantage d\u2019Climiner la lumikre parasite due aux impacts du faisceau laser sur les fenetres d\u2019entrke eY de sortie et c\u2019est cette dernikre que nous avons choisie d\u2019utiliser pour nos expCriences. ~~ EXEMPLES D\u2019APPLICATION Le Tableau 1 rCsume les rksultats obtenus avec le sulfure de carbone dans le mkthanol, le thiophkne dans le benzkne, le 2-nitrophknol (2 NP) et les 2,4- et 2,5- dinitrophCno1 (DNP) dans I\u2019eau,6 ainsi que le 4-arsonate 4\u2019-dimCthylaminoazobenzkne en solution aqueuse, Les Figures 3(a) et 3(b) montrent les spectres Raman du 2,5-DNP dans l\u2019eau obtenus 1 diffkrentes concentrations" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001696_s0263-8223(01)00168-4-Figure5-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001696_s0263-8223(01)00168-4-Figure5-1.png", + "caption": "Fig. 5. Spring subject to diametrical compressive surface load in surface-loading condition.", + "texts": [ + " (12), the strain energy expression becomes U \u00bc 1 Do Z w=2 0 MB \u00feF 2 s 2 ds \u00fe 1 Do Z p a a MB cosh \u00feFR 2 cosasinh 2 Rdh \u00fe 1 b Z p a a MB sinh \u00feFR 2 1\u00f0 cosasinh\u00de 2 Rdh \u00fe 1 Do Z w=2 0 MB \u00feF 2 s 2 ds\u00fe 1 b Z w=2 0 FRcosh\u00f0 \u00de2ds: \u00f013\u00de Hence, the strain energy expression becomes U \u00bc F 2R3 1 Do 4/2 sina \u00fe 2/ sin2 a\u00fe 1 3 sin3 a \u00fe 2/2 p 2 a sin2a 2 \u00fe 1 2 cos2 a p 2 a\u00fe sin2a 2 \u00fe 1 b 2/2 p 2 a\u00fe sin2a 2 \u00fe 4/cosa\u00fep 2 a \u00fe 1 2 cos2 a p 2 a\u00fe sin2a 2 \u00fe 4sina ; \u00f014\u00de where / \u00bc 1 Do 1 2 sin2 a 1 b cos a 1 Do 2 sin a \u00fe p 2 a sin 2a 2 \u00fe 1 b p 2 a \u00fe sin 2a 2 ; which leads to dxz \u00bc oU oF \u00bc FR3 Do 4/2 sina \u00fe 2/ sin2 a\u00fe 1 3 sin3 a \u00fe 2/2 p 2 a sin 2a 2 \u00fe 1 2 cos2 a p 2 a \u00fe sin 2a 2 \u00fe FR3 b 2/2 p 2 a \u00fe sin 2a 2 \u00fe 4/ cos a \u00fe p 2 a \u00fe 1 2 cos2 a p 2 a \u00fe sin 2a 2 \u00fe 4 sin a \u00f015\u00de and Kxz \u00bc F dxz \u00bc 1 R3 1 Do 4/2 sina \u00fe 2/ sin2 a\u00fe 1 3 sin3 a \u00fe 2/2 p 2 a sin2a 2 \u00fe 1 2 cos2 a p 2 a\u00fe sin2a 2 \u00fe 1 b 2/2 p 2 a \u00fe sin 2a 2 \u00fe 4/ cos a \u00fe p 2 a \u00fe 1 2 cos2 a p 2 a \u00fe sin 2a 2 \u00fe 4 sin a 1 : \u00f016\u00de 2.2.1. In-plane compressive stiffness (Ky) The composite spring is subject to uniformly distributed compressive surface load in y-direction on the flat surface as shown in Fig. 5. In derivation of the strain energy expressions, it is reasonable to assume that the sandwiched flat contact surfaces are rigid bodies. Therefore, strain energy is mainly stored in the flexible arms when under loading. Since the rigidity of the sandwiched flat contact surfaces is very high, the uniformly distributed load can be simulated by a resultant force F acting at the middle of the sandwiched portions [12]. Hence, the strain energy of the complete spring is U \u00bc L Z n D11D66 D2 16 jDj M2 dn : \u00f017\u00de By substituting the moment resultants into Eq" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002214_robot.1986.1087733-Figure8-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002214_robot.1986.1087733-Figure8-1.png", + "caption": "FIGURE 8 COORDINATE SYSTEM DEFINITION", + "texts": [], + "surrounding_texts": [ + "CCRIACT MU IBCREASES S I i H APPLIED FORCE\nLE COLUNN REAO IWIPUII\nFIGURE 2 TACTILE ARRAY-OPERATIONAL OETAlLS\nd e v e l o p m e n t o f t h e t a c t i l e s e n s o r system a r e a n A p p l e I I e m i c r o c o m p u t e r ( 6 5 0 2 b a s e d ) , a n in te r face board which conta ins two 6522 V e r s a t i l e I n t e r f a c e A d a p t o r s , a n e l e c t r o n i c i n t e r f a c e d e v i c e w h i c h c o n t a i n s a n a n a l o g - t o - d i g i t a l c o n v e r t e r and a compliant t ac t i le sensor pad. A schematic diagram of t h e t a c t i l e s e n s o r system is shown i n F i g u r e 3 .\nT A C T I L E SENSOR\n16 16 ... . . .\nI I 1 I 1 I 1 ! I DIGITAL , I\nM I C R O - PROCESSOR\nFIGURE 3 SENSOR SYSTEM SCHEMATIC\nAn Applesoft BASIC program c a l l e d SENSOR was w r i t t e n w h i c h c a l l s r o u t i n e s t h a t p e r f o r m a number of t a s k s i n c l u d i n g , a c q u i s i t i o n of d a t a from the sensor, graphic displays of t h e d a t a and c a l c u l a t i o n o f t h e p e r i m e t e r a n d a r e a of a n o b j e c t ( s e e l a t e r s e c t i o n s ) . The menu crea ted by program SENSOR showing all t he op t ions ava i l ab le is shown i n F i g u r e 4 .\nF i g u r e 5 shows a r e p r e s e n t a t i o n o f a h e x a d e c i m a l a r r a y d i s p l a y . The l a r g e r t h e hexadec ima l va lue the g rea t e r t he fo rce which is a p p l i e d t o t h e s e n s o r e l e m e n t . A hexadecimal v a l u e o f 00 i n d i c a t e s t h a t no f o r c e i s b e i n g appl ied to the sensor e lement . T h i s p a r t i c u l a r d i s p l a y r e s u l t s f r o m a s l e n d e r o b j e c t p r e s s e d d i a g o n a l l y a c r o s s t h e a r r a y . A l t h o u g h e a c h e l e m e n t of t h e a r r a y p r o d u c e s a n e i g h t b i t f o r c e v a l u e and can theref o r e r e s o l v e f o r c e s t o o n e p a r t i n 2 5 6 , s u c h a\nt h r e s h o l d e d i n t o f o u r l e v e l s s i m p l y f o r e a s e of presenta t ion . F igure 6 shows the equiva len t g rey s c a l e d i s p l a y o f t he da t a shown i n F i g u r e 5.\nTwo add i t iona l sub rou t ines GREEN and BINARY a r e a l s o u s e d t o d i s p l a y d a t a on t h e s c r e e n . These sub rou t ines ope ra t e the same as subrout ine DRAW excep t fo r t he f ac t t ha t t hey do no t d i sp l ay t h e i m a g e i n t h e c e n t e r of t h e s c r e e n . Subr o u t i n e GREEN p l o t s a g r e e n - s c a l e image on t h e l e f t s i d e of t h e s c r e e n a n d s u b r o u t i n e BINARY", + "p l o t s a b ina ry image (g reen and b l ack ) on t h e r i g h t s i d e of t h e s c r e e n ( s e e F i g . 7 ) . These s u b r o u t i n e s a r e u s e d t o d i s p l a y t h e r e s u l t s ob ta ined f rom pose es t imat ion ( see next sec t ion) .\nGray-scale image\nBinary image\nThe r e l a t i o n b e t w e e n t h e c o o r d i n a t e s u, v and x, y of an element of a r e a dA a r e :\nu = xcose + ys ine v = ycose - x s i n e\nS u b s t i t u t i n g f o r u and v i n t h e e x p r e s s i o n f o r I u g ives :\nI = ~v 2 d~ = s(yC0se - x s i n e l 2 dA U\n= COS 2 2 e I y dA - 2sinecose Ixy dA + s i n 2 0\nr x 'd~ taking equat ions (1) into the account leads to:\nI = I cos 8 - 21 sinecose + I s i n e (2) 2 2\nu x XY Y\nS i m i l a r l y , e x p r e s s i o n s f o r I, a n d Lv a r e obtained:\nIv = I, s i n 8 + 2 1 s in8cos8.+ i c o s 0 (3)\nIuv = I, sinecose + I (cos e - s i n e ) - ( 4 )\n2 2 XY Y\n2 2 XY\nFIGURE 7 GRAY SCALE AND BINARY IMAGES I sinecose Y\n4. Pose Estimation\n4.1 I n e r t i a and DrinciDal axes\nC o n s i d e r t h e a r e a A and the coord ina te axes x and y ( F i g u r e 8 ) . I f t he moments and product of i n e r t i a\nIx = I y dA I = I x dA I = Ixy dA 2 2\nY XY (1)\no f t h e t o t a l a rea A a r e known, t h e moments and product of i n e r t i a Iu, I,, I,, of t h e a r e a A with r e s p e c t t o new a x e s u and v can be obtained by r o t a t i n g t h e o r i g i n a l a x e s x and y a b o u t t h e or igin through an angle 8.\ne q u a t i o n s ( 2 ) , ( 3 ) and ( 4 ) may b e w r i t t e n as fol lows:\nIx + I Ix - I I =\nV 2 cos28 + I sin20 (6)\nXY\nSince the product of i n e r t i a w i t h r e s p e c t t o t h e p r i n c i p a l a x e s i s zero, the angle ep, which d e f i n e s t h e o r i e n t a t i o n of t h e p r i n c i p a l a x e s f o r t h e area, may be found by s e t t i n g IUV i n e q u a t i o n (7) equa l t o ze ro and then so lv ing fo r 0.\nTherefore , a t e = O P ,\n- 2 1 t a n 28 = XY\nQ Ix - I Y\n4.2 P o s i t i o n and o r i e n t a t i o n of a n o b i e c t\nOne of t h e o p t i o n s o f t h e m a i n p r o g r a m S E N S O R i s t o f i n d ' t h e p o s e o f a n o b j e c t i n c o n t a c t w i t h t h e s e n s o r pad. A number of BASIC and assembly language subroutines were developed w h i c h d e t e r m i n e t h e o b j e c t ' s p o s i t i o n a n d o r i e n t a t i o n by ca l cu la t ing the cen te r of g r a v i t y a n d t h e p r i n c i p a l a x e s of t h e o b j e c t ' s image. The object f rame xo, yo and the sensor pad frame X , Y t o w h i c h t h e c e n t e r of g r a v i t y a n d t h e p r inc ipa l axes a re r e fe renced a re shown i n F i g u r e 9.", + "The ma in p rog ram SENSOR c a l l s t h e b a s i c s u b r o u t i n e C E N T E R w h i c h i n t u r n c a l l s t h e assembly language subroutine SUM. Subrout ine SUM c r e a t e s a b inary image of t h e o b j e c t . A p i x e l w i t h a c o l o r v a l u e of 1 5 ( g r e e n ) i s considered o f f and a pixel with a co lor va lue of 0 t o 1 4 is c o n s i d e r e d o n a n d i s a s s i g n e d t h e c o l o r 0 ( b l a c k ) . S u b r o u t i n e SUM a l s o c a l c u l a t e s t h e v a l u e s of t he image area, t h e sum of t h e image columns, xSum (x coordinates) and the sum of t h e image rows, ysum (y coordinates). These values a r e t h e n p a s s e d t o s u b r o u t i n e CENTER which c a l c u l a t e s t h e x and y components of t h e c e n t e r of grav i ty , xcg and ycg using the equat ions:\nOnce t h e c e n t e r of grav i ty is ca l cu la t ed the v a l u e s a r e p a s s e d t o program SENSOR and a binary image of t h e o b j e c t i s p lo t t ed on the r igh t s ide of t h e s c r e e n by subrout ine BINARY. The c e n t e r of g r a v i t y i s t h e n p l o t t e d i n t h e c o r r e s p o n d i n g p i x e l l o c a t i o n on the b inary image ( f ig s . 10 and 11) .\nR e p l a c i n g t h e i n t e g r a l w i t h a summation f o r t h e d i s c r e t e image, the genera l equat ions become:\nI = E x ' A Ixl = z y ' A Ix ly l = 2 2\nY'\nwhere A i s t h e a r e a of an element. a l l t h e elements have the same area.\nFrom Figure 9 t h e r e l a t i o n s h i p b\nCx'y'A (10)\nI n t h i s c a s e\netween the known c o o r d i n a t e s x , y , which a r e measured with r e s p e c t t o t h e s e n s o r a x e s and the coord ina tes x' , y ' i s seen to be:\nX I = x - x Y ' = Y - Ycg cg\nS u b s t i t u t i n g t h e a b o v e r e l a t i o n s i n t o t h e equations (10) gives:\nThese equations are analogous to equation (1). E q u a t i o n (8 ) may now be used t o c a l c u l a t e t h e two v a l u e s o f 8 by s u b s t i t u t i n g i n t h e r e l a t i o n s f o r I X t , I I and & l y t g i v e n i n equat ions (11, (12) and ( 1 3 1 ," + ] + }, + { + "image_filename": "designv11_65_0001972_6.2002-4703-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001972_6.2002-4703-Figure1-1.png", + "caption": "Fig. 1 { Three views and perspective view of the UAV.", + "texts": [ + " Introduction In this paper we present the design of a multivariable ight control system for a shrouded-fan uninhabited aerial vehicle (UAV) that is currently under development at the University of Rome \\La Sapienza\" and Polytechnic of Turin. The VTOL vehicle, the characteristics of which are reported in detail in Refs. 1-3, is powered by three twostrokes, air-cooled engines and has a MTOW of 900 N including a 100 N sensor payload for scienti c applications to be accommodated in the doughnut-shaped shroud. Thrust is provided by two counter-rotating, three bladed rotors operating at constant angular speed. The solid model of the aerial platform is shown in Fig. 1. The open-loop dynamics of the vehicle were investigated in Ref. 1. The result showed that, unlike helicopter, coupling of longitudinal and lateral dynamics in forward ight is weak because of plat- 1 Copyright \u00a9 2002 by the American Institute of Aeronautics and Astronautics, Inc. All rights reserved. form axial symmetry. Also, some unfavourable characteristics were highlighted, such as the presence of unstable pendulum modes for both longitudinal and lateral degrees-of-freedom and a relatively high frequency of periodic modes that somewhat prevents approaches to control system design based on time-scale separation" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000292_jsvi.1996.0303-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000292_jsvi.1996.0303-Figure1-1.png", + "caption": "Figure 1. Assumed deformation pattern in the multi-layered conical shell.", + "texts": [ + " The developed computer program has been applied to an N layered shell to obtain results for the vibrations of composite material cylindrical shell (apex angle a 2 =0) and these have been compared with reported results [2]. The program has also been validated for the vibration of elastic three-layered conical shells. Results for the variations of the resonance frequencies and associated system loss factors with fiber orientation angles and cone apex angles for axisymmetric vibrations of cross-ply and angle-ply conical shells are presented. The cross-section of a N layered truncated conical shell is shown in Figure 1. The orthogonal curvilinear co-ordinate system X\u2212f\u2212Z is aligned to the conical geometry as illustrated in Figure 2(a). If there is assumed to be no slip at the interfaces of the layers, the displacements at a point zi from the mid-plane of the ith layer along the X- and f-directions (see Figure 1) are given by uzi =(1/ti )[ui{ti /2\u2212 zi}+ ui+1{ti /2+ zi}], vzi =(1/ti )[vi{ti /2\u2212 zi}+ vi+1{ti /2+ zi}], (1) where ui , vi , ui+1 and vi+1 are the displacements at the radii ROi and ROi+1 of the ith layer along the X and f directions, respectively, and ti is the thickness of the ith layer. The strains at the point are given by (oxx )i =(1/ti )[ui,x{(ti /2)\u2212 zi}+ ui+1,x{(ti /2)+ zi}], (off )i = {1/(ri ti )}[{vi,f + ui sin a}{(ti /2)\u2212 zi} + {vi+1,f + ui+1 sin a}{(ti /2)+ zi}+ tiw cos a], (gfz )i = {1/(ri ti )}[ti w,f \u2212 vi cos a{(ti /2)\u2212 zi}\u2212 vi+1 cos a{(ti /2)+ zi}]+(vi+1 \u2212 vi )/ti , (gzx )i =w,x +(ui+1 \u2212 ui )/ti , (gxf )i = {1/(ri ti )}[{ui,f \u2212 vi sin a}{(ti /2)\u2212 zi}+ {ui+1,f \u2212 vi+1 sin a}{(ti /2)+ zi}] + (1/ti )[vi,x{(ti /2)\u2212 zi}+ vi+1,x{(ti /2)+ zi}], (2) where ri =ROi + x sin a+ zi cos a and w is the transverse displacement" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002201_iemdc.2001.939386-Figure13-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002201_iemdc.2001.939386-Figure13-1.png", + "caption": "Fig. 13. Setup for Flow Visualization in the Stator Channels", + "texts": [ + " Also, for this case the curves converge towards the theoretical value of + = 0.632 for cp + 0. The improvement of the rotor performance is less significant for the case with high axial mass flow rates (Fig. 12) than it is for the case without axial mass flow (Fig. 11). C. Recirculation an Stator Gaps M = m 3 . (T3 . U 3 - T2 . U 2 ) (8) The power transferred to the fluid is then: Windows applied in the stator made it possible to visualize the flow in the stator gaps by a laser sheet technique (see Fig. 13). For this purpose a laser sheet was introduced normal to the axis of rotation in the middle of the stator cooling gap. Smoke was injected through flexible tubing to visualize the flow path. Without the modifications to the stator bars and gap supports a strong recirculation was detected. The coolant air is leaving the rotor-stator gap only through the slot on the windward side of the stator bar. The cooling air then passes through the cooling gap and into the stator plenum. It is then partially reingested P = w " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000167_jsvi.1997.1078-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000167_jsvi.1997.1078-Figure1-1.png", + "caption": "Figure 1. Load-displacement curves for an elasto-plastic body. (a) Kinematic hardening, (b) isotropic hardening.", + "texts": [ + " Some numerical results are shown for free and forced vibrations of the oscillator with kinematic, isotropic and kinematic/isotropic hardening. Elasto\u2013plasticity theory describes the behavior of bodies that, when subjected to a load P, present an elastic response until the elastic limit, defined by the yield surface, is reached. After this limit, the body has an irreversible response which is associated with plastic displacements [1]. The hardening phenomenon, observed on most metals and alloys, can be represented by a combination of kinematic and isotropic hardening. Figure 1 shows qualitative plots of elasto\u2013plastic behavior with kinematic and isotropic hardening. Consider a load history imposed on the body. By increasing the load from the origin of the force\u2013displacement space, there is an elastic domain until the limit Py is reached (point A). Outside the elastic domain, irreversible plastic displacement begins to occur. Displacement continues to increase until point B is reached. At this instant, load begins to decrease and the body presents an elastic reponse. Point C shows the residual displacement imposed by this load process. If the load continues to be decreased, the body keeps presenting an elastic response until the new elastic limit is reached (point D). This limit was altered by the load history. Kinematic hardening requires that the yield surface conserves its original size while it varies its position, whose center is originally at the point O (Figure 1(a)). Isotropic hardening, on the other hand, requires that the yield surface expands but conserves its original center which is initially at the point O (Figure 1(b)). Point D represents the new position of the yield surface limits for both cases. A constitutive model to describe this elasto\u2013plastic behavior is considered assuming an additive decomposition, i.e., the total displacement, x may be divided into an elastic part, xe, and a plastic part, xp. Hence, the force\u2013displacement relation is given by P=K(x\u2212 xp), (1) where K is a stiffness parameter. The isotropic hardening is described by introducing an internal variable, a, referred to as the internal hardening variable" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000024_s0167-8922(08)70489-9-Figure9-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000024_s0167-8922(08)70489-9-Figure9-1.png", + "caption": "Figure 9. Oil ilm pressure and clearance shape along the centre line of the bearing at 3 degree crank angle", + "texts": [ + " In the test, an axial groove is machined on the lining close to the bearing crown in order to disrupt the formation of oil film and thus induce the occurrence of bearing seizure. 646 The optical examination of the tested bearing shows that the bearing surface immediately downstream the axial groove is significantly polished. So is a very narrow band immediately upstream the groove. Interestingly, a column of pits also appears inside the polished narrow band. Figure 8 give a simple illustration of these observed features. EHL analysis of the bearing behaviour was carried out. In Figure 9, the clearance shape and the profile of the oil film pressures along the centre line of the bearing is presented. 'The results are taken at 3 degree crank angle, when the peak oil film pressure peaks. The diagram is in the same format of Figures 1 and 2, with the journal rotating in the clockwise direction. The upward slot of the surface a1 the upper left comer of the bearing represents the axial groove. It is obvious that the oil film thickness in the region downstream to the axial groove is very thin" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002422_2002-01-0587-Figure3-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002422_2002-01-0587-Figure3-1.png", + "caption": "Figure 3 \u2013 Variator Module Test Rig", + "texts": [ + " FE analysis of the contact (Figure 2) shows that traction coefficient of less than 0.1 do not affect the stress distribution greatly. In addition, the influence of the variator contact geometry on reducing the film thickness is much weaker than other parameters e.g. fluid viscosity. Hence the life analysis methods used to assess gear and bearing fatigue durability should be applicable to the IVT variator traction drive contacts. Durability fatigue tests have been carried out on a power re-circulating four-square test rig, shown in Figure 3, which incorporates a 100mm IVT variator test module. Hence real contact geometry and operating conditions could be used. The ratio (Rv) of the input disc speed to the output disc speed was set via the rig gear set, the majority of test being conducted at a nominal Rv=1.0. The rotational speed and torque, of both the module mainshaft (containing the disc and rollers) and the module layshaft, were measured along with the temperature of all six variator rollers. The rms. vibration level of the variator module was measured using an accelerometer" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002108_12.7972960-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002108_12.7972960-Figure1-1.png", + "caption": "Fig. 1. Schematic diagram showing the shadow moire contouring arrangement.", + "texts": [ + " Thus, it is important to coat the target (if it is too dark or too reflective) with a material which is a good scatterer of light: a thinned white paint is often effective, and so too is nickel evaporated onto the surface in a poor vacuum (10 -2 Torr instead of the customary 10 -6 Torr). The coating should be of uniform thickness. The advantage of evaporating over painting is that a thin coating of controlled thickness (typically 10 -7 m) can be applied. The coating can be removed afterwards. Figure 1 is a schematic diagram showing the shadow moire arrangement. The point source of light illuminates the prepared surface through a \"coarse\" grating (i.e., a grating possessing a line spacing too large for use as a diffraction grating). A shadow of the grating is cast onto the surface. When the surface is viewed through the same grating, moire fringes can be seen superimposed on the surface. Takasaki9 has shown that if the light source is placed at the same height above the grating as the lens of the observer, then the moire fringes become contour lines of equal depth below the grating plane" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001307_960083-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001307_960083-Figure1-1.png", + "caption": "Figure 1: The view on the investigated engine showing the prechamber.mnm", + "texts": [ + " On the basis of the results of investigations (7,8,9), it can be concluded that installation of a catalytic insert in the prechamber will facilitate ignition, extend the flammability limits and ensure pure combustion of lean mixture in the main chamber. The engine operating with a high compression Otto cycle fed by a lean air-fuel mixture was built based on a four-cylinder engine Ford FSD 425.The geometry of its combustion chamber established a high charge swirl ratio and a relatively high compression ratio. The modification of a Diesel engine to the Otto cycle was performed in a simple way, i.e. by replacing the fuel injectors with prechambers equipped with ignition plugs (Fig. 1). The catalytic insert, which is a platinum wire (coil l[mm] in diameter and 400[mm] in lenght), was suspended on a special hanger screwed into the prechamber. The prechambers were cooled with water drawn from the cooling circuit of the engine. Table 1 presents engine data for both variations (Diesel and Otto engines). In the Otto version the engine was tested with prechamber (the capacity of which was 24% of the capacity of the main chamber - see Fig. 2). For this configuration of the engine the geometric compression ratio was \u03b5=16" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000533_bf02192245-Figure2-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000533_bf02192245-Figure2-1.png", + "caption": "Fig. 2.", + "texts": [ + " The objective of player 2 is F~e C[P 1 x p2, R] and is chosen arbitrarily to be F2(p l , p2) = 0.75(p2 - 1)2 + 1.5(pj - 3) 2 + 0.75(pl - 3)(p2 - 1). The level curves Fi(l(c) are pictorially depicted in Fig. 1 (left), where the x-axis represents the decision pl e p1 c R of player 1, the y-axis represents the decision P2 ~/,2 c R of player 2, and each closed contour corresponds to the set of points (pl , Pz) such that F l ( p l , P2 )= c for c ~ R . The level curves F~2 I(c) are similarly depicted in Fig. 1 (right). In Fig. 2 (left), the two sets of level curves are superimposed. The Nash solution is computed to be (p*, p*) = (5/3, 5/3). With the update strategies A l ( p ) ( f f l ) =if1 - r iDp~Fl(pt . . . . . P i -1 ,P~ , Pi+ 1 . . . . . P u ) I : , , A 2 ( p ) ( : 2 ) =-~Oz- r2Op2F2(pl . . . . . P , - i , Pg, P,+1 . . . . . P u )[:2, the parallel gradient algorithm for this case is specified as follows: ~+1 = A~'(p')(p~), ~+1 = A~'(p')(p~). For a particular set of r~, rz, {7'}~=I, the game evolves as shown by the dots in Fig. 2 (right). Each dot with its associated stage t indicates the status of the game for that stage. It is observed that the evolution of the game corresponds to a very fast decaying oscillation around the Nash equilibrium point (p*, p*). Proposition 2.1 states that there is upper limit on the stepsizes of the players when they are resorting to parallel gradient descent. Figure 3 (left) JOTA: VOL. 90, NO. 1, JULY 1996 55 (L) Nash equilibrium; (R) parallel gradient descent with r~ = 0.5, r2 = 0.~, 7' = 5" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001007_jsvi.1997.0998-Figure5-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001007_jsvi.1997.0998-Figure5-1.png", + "caption": "Figure 5. Two loops connected by two pulleys, l and m.", + "texts": [ + " (4) The equations of motion for the other pulleys are obtained in the same manner, and the equations of motion for all pulleys can be expressed in matrix form in the same way as equation (3). In equation (4), the coefficients of um and u m are the summation of the coefficients of um and u m in the equation of motion which is derived separately with respect to each loop. Therefore, the matrices K and C of the equation of motion for a belt system which is connected by one pulley in each loop can be obtained by superposing the matrices K and C of the equation of motion for each loop. If two loops are connected by the pulley l and the pulley m, as shown in Figure 5, the equation of motion for the pulley m is Imu m \u2212 clmrlrmu l +(clm + cmn )r2 mu m \u2212 cmnrmrnu n \u2212 c'lmrlrmu l +(c'lm + cmn')r2 mu m \u2212 cmn'rmrn'u n' \u2212 klmrlrmul +(klm + kmn )r2 mum \u2212 kmnrmrnun \u2212 k'lmrlrmul +(k'lm + kmn')r2 mum \u2212 kmn'rmrn'un' =Tm . (5) In equation (5), the coefficients of um and u m are the summation of the coefficients of um and u m in the equations of motion which are derived separately with respect to each loop, and the coefficients of ul and u l are obtained in the same way. Therefore, the matrices K and C of the equation of motion for the belt system which is connected by two pulleys in each loop can be obtained by superposing the matrices K and C of the equation of motion for each loop" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003278_ias.1989.96735-Figure3-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003278_ias.1989.96735-Figure3-1.png", + "caption": "Fig. 3. Typical .current trajectory with a hysteresis current controller.", + "texts": [ + " With this assumption the instantaneous voltage equation for the phase 'a' of the IPM motor can be obtained from the equivalent circuit of Fig. 2. (9) di, dt Va = R,i , + Laa- + E a where, V, = terminal voltage of phase a i, = current through phase a Laa = synchronous inductance of phase a E, = the back-EMF of phase a R , = the resistance of phase a = Laa - Lab Similar equation can be easily written for the phases b and c. Inductance Calculation fiom the Hysteresis Current Controller The phase inductance of an IPM motor can be calculated analytically from the instantaneous voltage and current information. Fig. 3 shows the current waveform obtained when a hysteresis current regulator is used. The salient features of the hysteresis current controller has been discussed very well in the literature [ll-141. The mathematical equations for calculating tor. the phase inductance has been discussed by the authors recently [15]. These equations are repeated here for convenience and to facilitate comparisons to be made with the other methods discussed in this paper. In order to obtain a closed form expression for the phase inductance the following assumptions are made in the analysis: (i) the phase inductance La,, remains constant during a switching period. (ii) the motor back-EMF E,, remains constant during a switching period. These assumptions are valid if the hysteresis band is small and the switching frequency is sufficiently high (> 10 kHz). Fig. 3 shows a typical current trajectory with a hysteresis controller. The current equation for the section a-b where the current is increasing can be written as, where, io = i(t) = i , ( t ) = the initial current the current at any instant during the switching time A t irer( i ) = the reference current In particular, the current at the instant 'b', i ( t ) (b can be written as 7 where, i ( t ) la = the current at 'a' on the Fig. 3. 714 The equation (11) can be expressed in terms of the reference current a, and the band width AIh as, de (e; - e j ) E - K - = K - dt (ti - t j ) a - where Oi and B j are the positions at two instants of time ti and tj respectively. In Fig. 3 when the rotor is such that the current is at point \u2018c\u2019 then 0; is the value 0 evaluated at \u2018b\u2019 and O j is the value of 0 evaluated at \u2018a\u2019 and ti = t2 and t j = t l . This approximation is sufficiently accurate at high switching frequencies. This is clear from the simulation results presented in the following sections. (Va - E a ) ( l - e - F) + i ( t ) l i 1 e - w (12) i(t)ltl+Ai = ~ Ra iv(t)ltl+At + AIh = -2-L?- (V - E - R a +(I,(t)ltl - A 1 h ) e - W (13) Direct Method of Inductance Calculation (Va - Ea) (1 - e - - ) ir(t)(tl+At + AIh = ____ Instead of calculating the phase inductance indirectly, the phase inductance can be evaluated directly in the following manner" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003545_j.abb.2004.02.016-Figure4-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003545_j.abb.2004.02.016-Figure4-1.png", + "caption": "Fig. 4. (A\u2013C) Role of the enzyme positions on the kinetic profiles of SP2 . The D1 \u00bc 0:07V//Dwm \u00bc \u00fe0:14V/Dw2 \u00bc 0:07V; (c2) Dw1 \u00bc 0:09V/Dwm \u00bc Dw2 \u00bc 0:1V. The numerical values of the other parameters are reported in", + "texts": [ + " To observe more precisely this phe- nomenon, we have analysed the influence of different numerical values of the potential (c) on CR (i.e., (c1) Dw1 \u00bc 0:07V/Dwm \u00bc 0:14V/Dw2 \u00bc 0:07V; (c2) Dw1 \u00bc 0:09V/Dwm \u00bc 0:18V/Dw2 \u00bc 0:09V; and (c3) Dw1 \u00bc 0:1V/Dwm \u00bc 0:2V/Dw2 \u00bc 0:1V) whatever the enzyme positions inside the ionic double layers (i.e., for a values ranging from 0 to 1). These simulations are d (ATP/Mg)2 , respectively, for (1) a1 \u00bc a2 \u00bc 0, (2) a1 \u00bc a2 \u00bc 1=3, (3) 0:09V/Dwm \u00bc \u00fe0:18V/Dw2 \u00bc 0:09V. The numerical values of the reported in Fig. 4A\u2013C. Clearly, we observe an optimal enzyme position permitting the higher concentration ratio. Furthermore, the effect appears more pronounced for high absolute values of the surface potential. This effect can be explained by two antagonist roles played by the membrane surface charges. For this potential (c) and on the one hand, the positive charges covering the outer membrane surface attract the negatively charged sub- strate. Therefore, the closer the membrane-bound phosphatase to the outer membrane surface (i" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000759_004051759106101210-Figure6-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000759_004051759106101210-Figure6-1.png", + "caption": "FIGURE 6. A warp goes through two filling yams.", + "texts": [ + " Considering the situation where the warp goes from the bottom of ellipse ol to the top of ellipse o2, the expression P - E becomes where yo, = the vertical coordinate of the center of the ellipse over which the warp passes, and Y02 = the vertical ,. FIGURE 4. Fabric geometry with filling crimp. at UCSF LIBRARY & CKM on March 8, 2015trj.sagepub.comDownloaded from 763 FIGURE 5. Beat-up process. coordinate of the center of the ellipse underneath which the warp passes. During beat-up with an unbalanced shed (Figure 5 ), the first filling yarn is forced downwards and the second filling yarn upwards when the back rail is above the cloth line. The situation where the warp goes through the two filling yarns is shown by Figure 6, from which we get Then Combining Equations 13 and 14, we get P - E as follows : When the positive is suitable, Y02 - Yo, < 2b + d, , and when the negative is suitable, y~ - yo, > 2b + d, . Knowing the crimp ( yo, - y~ ) of the filling yam, the diameter ( d, ) of the warp, the long, short radius (a, b) of the cross section of the filling yam, and the angle E between the warp and the cloth line, we can calculate the distance P between adjacent filling yams using Equation 15. Knowing the values of (y, - y~ ), d, , a, b and P, we can also calculate the angle E with Equation 15, using a numerical method" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002762_s0094-114x(00)00039-2-Figure2-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002762_s0094-114x(00)00039-2-Figure2-1.png", + "caption": "Fig. 2. 10-link mechanism derived from chain No. 118.", + "texts": [ + " Thus this 10- link chain yields only three ``true'' 10-link mechanisms with the displacement problems for remaining mechanisms no more complex than that of an 8-link mechanism. These three mechanisms are: (i) \u00aex link 1, input link is 6; (ii) \u00aex link 1, input link is 2; and (iii) \u00aex link 2 and input link is 3. Therefore for all chains containing a dyad which is not a part of a 4- link loop, only three true 10-link mechanisms are possible and the analysis of these three mechanisms is considered next. Consider the 10-link mechanism shown in Fig. 2 obtained by \u00aexing link 1 in chain 118. The global coordinate is oriented such that OD is coincident with the x-axis h1 0 . For this mechanism with four independent loops, the loop-closure equations are: Loop 1 (OABCD). r2 cos h2 r1 r6a cos h6 c r10a cos h10 d r3a cos h3 a1 ; r2 sin h2 r6a sin h6 c r10a sin h10 d r3a sin h3 a1 : 4 Loop 2 (CBEFG). r10a cos h10 d r3b cos h3 a2 r4 cos h4 r6b cos h6 c g p r5 cos h5 ; r10a sin h10 d r3b sin h3 a2 r4 sin h4 r6b sin h6 c g p r5 sin h5 : 5 Loop 3 (CJLKG)", + " R X26 n 1 an cosn h3 bn cosn\u00ff1 h3 sin h3 a0 0: 13 Finally a use of tangent half-angle identities for cos h3 and sin h3 leads to R X52 n 0 cntn 3 0; 14 where t3 tan h3=2 . By using the leading-coe cients condition of Theorem 3 it was veri\u00aeed that R is devoid of any tangent half-angle substitution related extraneous solutions. Therefore, given a value of h6; the I/O polynomial is of 52nd degree in t3: For the numerical data given below and h6 69:5\u00b0; the coe cients of this 52nd degree polynomial including the 52 possible solutions for t3 are given in Table 1. Consider again the 10-link mechanism of Fig. 2 but with link 2 as the input link. The loopclosure as well as the reduced loop-closure equations can be written in condensed form as Loop 1 OABCD Loop1 2; 3; 10; 6 ! f12 h3; h6; h10 ; Loop 2 OAEFGD Loop2 2; 3; 4; 5; 6 ! f3 h3; ~h4; h5; h6 ! F2 h3; h5; h6 ; 15 Loop 3 CJLKG Loop3 10; 8; 7; 5; 6 ! f4 h5; h6; h7; ~h8; h10 ! F3 h5; h6; h7; h10 ; Loop 4 OABHIKGD Loop4 2; 3; 9; 7; 5; 6 ! f5 h3; h5; h6; h7; ~h9 ! F4 h3; h5; h6; h7 : In the equations for Loop 1, no unknown rotation is eliminated. Next solving the two equations for Loop 1 (f1 and f2 in f12) for cos h10 and sin h10 in terms of h3 and h6 we get cos h10 gc h3; h6 ; sin h10 gs h3; h6 : 16 Squaring and adding both sides of Eq", + " The two solutions t3 0:6017 0:9584i from n1 yield h3 103:1 70:06i)\u00b0. Since n1 and n2 have a common factor n leading to n1 n2 0; the polynomial R0 2 k t3 (Eq. (22)) contains an extraneous factor leading to t6 i in F1 and F2;34: This factor is given as gcd n1; n2 ! n n1 t2 3 \u00ff 1:2034633t3 1:2805253 and the true I/O polynomial devoid of any extraneous roots is expressed as R R0 n X52 n 0 cntn 3 0: 25 Hence, with link 1 as the ground link and link 2 as the input link, link 3 (as well as the mechanism of Fig. 2) has 52 possible assembly con\u00aegurations. The coe cients as well as the solutions for the 52nd degree I/O polynomial of Eq. (25) are given in Table 3. Consider next the 10-link mechanism that is derived from chain 118 by \u00aexing link 2 with link 3 as the input link. In condensed form, the loop closure equations can be written as and the reduced loop-closure equations are F1 h6; h10 0; F2 h5; h6; h10 0; F3 h5; h6; h7; h10 0; F4 h5; h6; h7; h10 0: 27 Eq. (27) is similar to Eq. (19) discussed earlier" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000727_s0167-8922(08)70486-3-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000727_s0167-8922(08)70486-3-Figure1-1.png", + "caption": "Figure 1 Geometry of half toroidal CVT [7]", + "texts": [ + " Apart from a simple rolling contact, a spin motion at the contact gives a kinematic variation to the EHL problems. In some machine elements, the spin motion plays a significant role in lubricated contacts. The spin motion in a n angular contact ball bearing, a t each contact between a ball and the outer or inner race, was well described in [6]. A continuously variable transmission (CVT) with rolling elements is another application in which the spin motion has often been discussed. For example, Figure 1 shows the schematic geometry of a halftoroidal CVT. The power is transmitted from the input disk to the output disk via a pair of power rollers. At the contacts (0 and 09 between the power roller and disks, a n EHL film is produced by the rolling motion, and in fact, the power is transmitted by the shear action of the lubricant film, so called traction drive. A few percent of sliding is essential to yield a sufficient shear stress at the contact, as suggested in [7]. The geometry of the half toroidal CVT also gives a spin motion to the contacts" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002064_bf02448936-Figure4-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002064_bf02448936-Figure4-1.png", + "caption": "Fig. 4 One-mass dynamics", + "texts": [], + "surrounding_texts": [ + "where k~ is a calibration constant and w is the output of the force-sensitive element.\nTo obtain a signal proportional to the body movements, a marker was attached to the body and recorded by a video camera. As a body marker we used a white spot on a dark background, which could be detected in the video signal as a high voltage level. By detecting the high level with an electronic comparator and measuring the interval between the start of the video line and the high level in that line, a signal v is obtained which is proportional to the anterior/posterior movements x:\nx = ks v (2)\nwhere k z is a calibration constant. The movements were detected each half frame. Based on a half-frame frequency of 50Hz, this resulted in a maximum detectable frequency for the movements of 25 Hz. Both signals were sampled (sample interval 20ms, sampling time 81.96 s) and fed into a computer. The torque signal was compensated in time by a mean delay of 20 ms (one sample point) to correct for time skew between the movement and the torque signal.\n2.2 Verification o f the measurement system\nTo verify the measurement system, we used a mechanical system whose frequency response function from torque to movements was computed. This function was compared with the measured frequency-response function estimated from the torque signal and the movement signal. The mechanical system and the measurement system, comprising the video camera and the platform, are depicted in Fig. 2. The mechanical system comprised a rotating mass at a height h with respect to the platform. Rotation was slowed by internal friction. The theoretical frequency response function from torque T to movement x, both in the xy plane, can be denoted as:\nMechanical system to test the measurement system\nwhere k is a constant, m is the mass of the rotating system, r the distance from the centre of gravity to the vertical axis of the system andfc is the cutoff frequency equal to\nf~ = (1/27r)(g/h) 1/2 (4)\nwhere g is gravitational acceleration. For g = 9 .8ms -2 and h = 0.13 m this results in fc = 1.1 Hz.\nThe platform produced a signal proportional to the torque w in the xy plane. The video camera was placed on\n470 Medical &\nthe y axis above the groundplane to detect a white spot marked on the top of the rotating system. This gave an output signal v proportional to the movements in the xy plane.\nModels of parametric or nonparametric analysis can be identified. Parametric analysis, using a time discrete recursive algorithm, is not recommended for dynamic body models because of uncertainties in establishing the model and its unstable character. Nonparametric estimation by spectral analysis is more applicable, since it provides an indication about the frequency behaviour of the system. The nonparametric frequency behaviour between input and output of the model is described by the coherence, phase and power gain spectra. To avoid calibration deviations, all computed power gain spectra are shifted to 0 dB for the maximum gain. This, however, does not affect model identification.\nThe frequency spectra for the measurements of the mechanical system are depicted in Fig. 3. We note that the coherence is greater than 0-90 in the frequency range 0.05- 6.00 Hz. The upper limit is caused by the maximum excitation frequency of the mechanical system. The decreasing coherence in the lower frequency band is only caused by the low signal-to-noise power ratio density (denoted as SNR density) in that band. The phase deviates from 0 ~ for higher frequencies because the delay compensation of the torque signal is not ideal. The deviation is less than 5 ~ at 5 Hz, which corresponds to a time difference between the input and output signals of 3 ms. The variations in phase and power gain spectra for high and low frequencies are related to the SNR density for these frequencies.\nThe power gain spectrum decreases with - 40 dB decade- 1 for frequencies beyond the cutoff frequency, which is in agreement with the frequency response function as given in eqn. 3.\nFrom these measurements we concluded that the measurement system comprising the video camera and the force-sensitive element resulted in reliable measurements in the frequency range 0.05-6.00 Hz.\n3 Model evaluat ion Dynamic body models in posturography describe the relationship between joint torques and body movements. These often complex equations can be simplified by making certain assumptions, the consequences of which can be evaluated by computer simulations.\nThe simulations were compared with measurements made on four subjects. Most results for the different subjects were comparable, and therefore only spectra for one subject are shown. If differences between individual subjects are significant, they will be discussed.\nBiological Engineering & Computing September 1985", + "inverted pendulum model to describe body\nthe ankle joint is justified, the model should be a good approximation of the dynamic behaviour of the body. For this model the following equations can be obtained ( K o o Z E K A N A N I et al., 1980):\nX = - d sin 0 (5)\ny = d cos 0 (6)\nT = JO - m2d cos 0 + (mj) - m#) sin 0 (7)\nwhere x and y are the horizontal and vertical displacements of the centre of gravity, 0 is the angle of the body with respect to the vertical plane, T is the torque in the ankle joint, m is the mass of the moving body, d is the distance from the ankle joint to the centre of mass, J is moment of inertia of the body and St, j~ and 0 indicate the acceleration components.\nIf only small movements (~c ~ 1 ~ are made, eqns. 5-7 can be modified to express a linear relationship between T and 0:\nT = J~\u2022 - - mgdO (8)\nwhere J' = J + m d z If the mass and the height of the feet are not taken into account, the torque in the ankle joints is equal to the torque resulting on the platform, and the frequency response function from torque to movement can be denoted as\nB(T) - O(f) _ ( j , f 2 _ rood)-' (9) T ( f )\nThis function is often used (ISHn)A and IMAI, 1980; NASHNER, 1970) in the evaluation of postural control.\nEqn. 9 results in a coherence value equal to 1 and a phase value equal to 0 ~ over the entire frequency range. The power spectrum decreases with - 40 dB decade- ~ for frequencies beyond the cutoff frequencyfc :\nfc = (1/2rO(mgd/J') ~/2 (10)\nFor m = 7 0 k g , g = 9 . 8 m s -2, d = l m and J = l k g m 2, this results infc = 0.5 Hz.\n3.2 Measurements\nTo approximate the single-mass inverted pendulum model, the first measurements were executed with fixation of parts of the body above the ankle joints. The resulting frequency spectra from these measurements are shown in\nMedical & Biological Engineering & Computing\nSmall differences between individuals in the coherence function for high frequencies could be observed (c) the variance of the estimated phase spectrum and the power gain spectrum increased for frequencies where coherence decreased (d) the power gain spectrum decreased at less than - 40 dB decade- 1 for frequencies larger than 0.4 Hz.\nMeasured spectra from a subject with body fixation\nWe also measured the frequency spectra based on the torque and either upper body movements (marker on the shoulder) or lower body movements (marker on the hip), without fixation of the body. The subjects were instructed to keep the body rigid.\nThe features of these spectra deviated from spectra measured with body fixation in respect of (see above):\n(b) in general, the coherence function was less for high frequencies. Differences between individuals were more marked (d) power gain spectra differed for measurements of the lower part of the body (power gain decreased by less than -30dBdecade -1) and the upper part (power gain decreased by less than - 40 dB decade- t).\n3.3 Evaluation of measurements\nThe spectra shown in Figs. 5 and 6 deviated significantly from those calculated from eqn. 9. There may be several\nSeptember 1985 471", + "contributory factors causing these deviations.\n3.3.1 Low SN R density. The signal power spectra of the recorded movements and torque indicated that signal power decreases according to frequency, that the greatest signal power was in low frequencies ( f < 1 Hz) and higher frequencies have less signal power. In particular, higher frequencies of the movement signal had a low SNR density because the power gain decreased by - 4 0 d B d e c a d e - with respect to the torque power gain. In our measurements, the SNR density of the movement signal was generally less than 10 dB for frequencies larger than 5 Hz. A decreasing SNR density resulted in decreasing values for the estimated coherence and increasing variance for the estimated values of phase and power gain. We simulated the theoretical curve of the spectra for measured decreasing SNR density. These spectra, denoted with a solid line in Fig. 7, indicated that coherence decreased for f > 5 Hz.\nComparing these spectra with those depicted in Figs. 5 and 6, it may be concluded that the lower SNR density does not result in a low coherence for the measurements in the frequency range 0.05-5 Hz.\nSimulated spectra for the measured spectral SNR density (solid lines) and the nonlinearities (broken lines)\n3.3.2 Nonlinearities. Eqns. 5-8 describe the nonlinear relationship between the torque T and the movements 0. As already indicated, the linear eqn. 9 is an approximation for small movements. To quantify the contribution of nonlinearities to the frequency response function, we simulated eqns. 5-8 on a computer and calculated the frequency spectra.\nTo simulate a measurement, the distribution of movements as measured was used as 'reference for signal quantification. This rectangular distribution had a mean angle of 1 ~ and a standard deviation of 1.5 ~ The resulting frequency spectra, shown in Fig. 7, indicated perceptible nonlinearity at low frequencies. A comparison of these spectra with the measured spectra (Figs. 5 and 6) demonstrated similarities for low frequencies ( f < 1 Hz). From this analysis it may be concluded that for these measurements nonlinearities only affect frequency behaviour in the low frequency range.\n3.3.3 Control of body dynamics; two-mass inverted pendulum model. The simple feedback model including control and body dynamics (Fig. 1) indicates that control acts between the movements of the body and the resulting torque. Actually human control of body posture also coordinates the various parts of the body; postural control is therefore also active between parts of the body.\nTo determine the influence of co-ordination, a two-mass inverted pendulum model, as illustrated in Fig. 8a, was examined. For this model the following equations in the\nfrequency domain, which apply only to small movements, can be derived:\n0. = Bu T. (11)\n01 = B 1 T l (12)\nT = T . + T~ (13)\nwhere the indices u and I indicate the upper and lower part of the body, respectively, and B represents the frequency response function of that particular part of the body, which has the same structure as the frequency response function denoted in eqn. 9. The system model describing the dynamics is depicted in Fig. 8b. From eqns. 11-13 we computed the cross-correlation functions between the resulting torque T and the movements of the upper and lower part of the body:\nSro,, = Bi, Srr + B, Sr, r~ (14)\nSTO l = BiSrr + BtSTtT= (15)\nThese equations indicate that the cross-correlation functions and, consequently, the transfer functions, depend on the cross-correlation functions between the joint torques of the upper and lower parts of the body. These joint torques are co-ordinated by biological control acting between these two parts of the body. It can therefore be concluded that measuring the frequency response function between the resulting torque and the movements of the body will result in a function which is a combination of body dynamics and body co-ordination.\nThe contribution of body co-ordination to the frequency response function between torque and movements was estimated by comparing the measured spectra with the spectra without body co-ordination.\nSeveral characteristics of biological co-ordination emerged from this:\n(a) Differences between simulated and measured spectra were mainly noted for high frequencies ( f > 1 Hz). From this observed and measured lack of coherence, it appears that biological co-ordination is active in this frequency range. This lack of coherence may be caused by nonlinear and nonstationary factors in the crosscorrelation function between torque and upper and lower body movements. The similarity of the measurements of individual subjects would tend to indicate that the nonstationary factor in a given measurement is less than the nonlinear factor. (b) The differences between individuals in the coherence function for high frequencies can be explained by individual biological co-ordination. (c) The differences in decreasing power gain of spectra measured between torque and the movements of the lower or upper part of the body indicate that control mechanism varies between these parts of the body.\nIt can be concluded that, even for a rigid body (0, = 0c),\n472 Medical & Biological Engineering & Computing September 1985" + ] + }, + { + "image_filename": "designv11_65_0003063_iros.1994.407373-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003063_iros.1994.407373-Figure1-1.png", + "caption": "Figure 1: Denavit-Hart,enberg coordinates and tip vector", + "texts": [], + "surrounding_texts": [ + "1 Introduction In t,he past., we proposed t,he closed-loop kineniat,ic calibration met8hod [3], which permit,s a. nia.nipolat,or t,o be calibrated without, endpointf sensing. I n Qliis niet,liod. a manipulator forms a mobile closed kinematic chain by a,t.t,a.chment. of t,he end effectlor to the environment,. This at,tachment, ma.y be rigid or may have lip t,o 5 unsensed degrees of freedom. By placing such a const,ra.ined manipulator into a numher of poses, t,he kinemat,ic parainet,ers may be calibrated using joint, angle sensing a.lone and t,he loop closure equations.\nThis paper presenh a.n experimental implenientat,ion of closed-loop kinematic calibration on t,lie Snrros Dextrous Arm. a 7-DOF rPdundant manipulator [ll]. With rigid at,tachment. of t8he endpoint, t,o the environment,, t.he Sarcos Dextrous Arm will form a mobile closed kinemaQic chain with 1 DOF. IInfort.iinat,ely, t.liis anthropomorphic a.rni will have an immobile elbow joint. during t.liis motion. In order t,o make this joint, mobile, and hence t,o identify t,lie parameters associat,ed with this joint. it is necessary t,o free at. least one of t81ie end effector const,raint.s. One way is t,o add a a external unsensed hinge joint.: t,liis sit,uat.ion was analyzed in [3] and a. procedure was given to e l in i i i~a t~~ the hinge joint, angle from the loop closure equa.t,ions.\nIn addition. it, is necessary to define a force control stxat.egy. which permits manipulat,or pose clianges in conipliance with a constrained end effect.or and init,ially incorrect, kinenmtic pa.ra.met,ers. In a previoiis iiiiplemc.tit,a.t,ioii on the fingers of the IIt,ah/MIT Dextrous Hand [ 2 ] . t,he fingers were manually placed int,o different, poses and advant,age was h k e n of t,he hackdriva.hilit,y of t,lie fingrr joints. For t,he Sarcos Dextmiis Arm. which is a liytlraiilic ma-\nnipulator, a.n active force cont.rol st,rat.egy is required t,o move t,he joint,s.\nIn t,he present, paper we are det8ermining only t,he joint. angle offsets and t1he hinge related paramet,ers. because our main purpose is recalilwa.t,ion of t.he joint. angle offset,s. That, is to say. we presiinie t.hat, the geoniet.ric parameters are already well enough known from tlir manufacturer's specificatmioils or from previous calibrations. Joint, angle offset, recalihration is required hecaiise of t,he use of a.nalog sensing at. the wrist, joint,s a.nd increment,a.l encoders in t,he arm join t.s.\n1.1 Related Research\nOther researchers have inipleinent,ed closed-loop calibration, iintler diverse endpoint constraint,s. In [14]: a line const,raint was defined by a laser, which was tracked using an endpoint. retroreflect,or on a PITMA 5GO and a 4- quadrant, detector. In [5 ] , a fiducial point, 011 the end effector is touched t,o a fiducial point, on t,he environment in several different poses; t.his corresponds t.o the point cont,a.ct caw in i.71. I n [12]. a teleoperatett excavator witJi unsensed joints was calibra.ted by adding an addit,ional linkage (called a calilwator by the authors) with some sensed joints t,o forin a closed loop. In ['i]. a ball bar wit,li fixed length and unsensed spherical ,joint,s a.t, each end was employed. (hsed-loop calibrat.ion of a. maiiipulat,or with a camera niount,ed on an end effect,or was present,ed in [ l i ] .\nIu t,he following. we first. review open-loop kineina.t,ic cali bratmion. then derive closed-loop kineiriat,ic calihration using a hinge joint.. Experimenbs on the Sa.rcos Dextrous Arm are t,fien presented.\n2 Methodology The Deiia\\rit,-Hart,enl,erg (D-H) convent,ion is employed for t.he geomebric paraniet,ers (Figlire I ) . In t,he present, case. Haya.t,i coortlinat,es are not. rrqiiired because the Sarcos Dext.roiis A r m has a,ll neighlmring joint,s ort,liogonal. The siibseqrient, t l ~ v ~ l o p i i i e n t is t.a.kcw from [3].", + "bi\n2.1 Open-Loop Kinematic Calibration For a manipulator with 71 DOFs. the end-effector is located by the position vect,or p: and t,he orientmatlion matrix R:,:\nn\nR: = ~ , ( o j ) ~ , ( ~ j ) ('2) j = l\nwhere Rz (4) and R,(4) a.re 3-hy-3 roi.at.ion mat,rices about, the z and x axes by t,lie angle I$. Tlir subscript, c indicates t,liat, t81ie p o s i h n and orieiitat,ion are computed from t,lie model. The superscript i indicates t,he con figuration of the m a.n i pu 1 a t.or. In kineinat. ic c ali b r at,ion, the manipulator must. he placed int,o 771 poses, with 8 = [si.. . . ,19:,]*. i = 1 , . . . , m, for ri links.\nThe required geometric pa.rameters are s j , u, j . and nj for links j = I , . . . . n. We will model only t.he joint angle offset. . relat,ed t,o t.lw actual 8; and measured B j DH joint angles by the relation 8; = Oj + 19;jf, For the calibration, we use a single vect,or which holds all the unknown kinematic paramet,ers cp = a. s . a]*. where s = [ s 1 , . . . .ss,]? 4 c .\nInstead of t.he orientat.ion niat,rix R:. i t is convenient. to represent, t,he orientation by the vector or pt = [4:. 4;, 4;lT. reprtwiiting t lie roII-pit.cIi-yaw (ZYX) ~ u ~ e r angles: Rf = Rz(+~)Ry(4~)R,(~~). The computed endpoint, location can t,lian be writ,t,en as x: = [pi . p?lT and given by :\nwliere the functerat.ion continues until AA' - 0. The basis for linearization is the assumption that x: is\nclose t,o xf . Then\nwhere Alii = [d.?. dy', dzz]?' is t8he incremental position error, and Apz = [a4:. is the incremental orient8a.tion error i n t,erm of t,he Euler angles.\n{]sing differeiit,ial rotaation about orthogonal axis rather than non-orthogonal Euler angles, we have Ar' = [ar i , B y i . 02']*. The rrlat3ion between bot,h are given in [3]. The Jacobia.ns are t,lieu found by screw axis analysis as in [3].\n2.2 Closed-Loop Calibration with Fixed Endpoint\nWe next. consider a redunda.nt, manipulator (2 7 DOFs) with fixed end-point . ( - h \" l y . t,he result,iiig closed-loop chain will be mobile, since t,hr fixed endpoint. const,raint,s only 6 of the 7 DOFs of t,lie ma.nipula.tor. We ca.n set, t.lw reference fra.nie to he at. t,he fixed endpoint and t,o ha.ve zero orieirta.t,ion a.nd posit,ion. Hence x:, = 0. and no nieasiirenient,s are reyirirrd I)ecaiise t.he a.ctiia.1 posit,ion is known and is zero by drfiiiition.\n330\nI", + "Using the previous mathematical development, with one modification, we can write:\nwhere -hi = (art , ayt,, azf) is the computed orientation, and -Api = ( d r f , d y t , d : : ) is the computed position. Because the \"measured angles are defined as zero, differential rotation around non-orthogonal Euler axis is equivalent t o differential rotation around orthogonal axis. Thus, for the closed-loop case, we have Ap' = hi.\nFor the positional component of the loop closure equations, we have\nn\np', ,= sjzj-l + ajxj = 0. (9) j=l\nOrdinarily in the closed-loop procedure this equation is a problem, because the length parameters are linearly dependent. To proceed. it is necessary 6o specify one length parameter to scale the system size. This is not a problem for the present case. since t8he manipiilator link lengths are presumed known.\n2.3 Closed-Loop Calibration with a Hinge Joint\nTo make the elbow joint mobile, we add a pasqive, 1111- sensed 1-DOF rotary hinge joint.. Tlie hinge is defined to be the -1 joint,, with positioned along the hinge axis (Figure 2). The endpoint, coordinat,es being a r b i h r y , it is convenient, to make the last joint- zf coincident with t8he hinge axis ;II. The hinge coordinates origin can then he positioned a t the endpoint, coordinates origin.\nThe geometric parameters needed to define this added unsensed DOF are so, ao, and Q O , defined from hinge coordinates -1 to the manipulator base 0. and RI. We also need to calibrate the parameters between coordinate 6 and 7: ai, S i , and a7. The hinge angle 0: = -0; measured about &, is unknown and must be eliminated from the 6 kinematic loop closure equations. All other parameters related to the arm are known from the manufact,urer's specific at ions.\nTo apply the method from hinge to endpoint, we have the position vector. p: = 0 from (9) and Api = [hi, dy', &'IT from (T), and the endpoint, orientation matrix Rf. and Rio,:\nwhere IJ is the 3-by-3 identity matrix, Riot is the total rotation, including the uiisensed hinge joint angle: Rt is the total rotat,ion, excluding the hinge joint, Ozi and ayi are infinitesimal rotation along axis r and y and 0: is finite rotation along axis z of the coordinates frame -1.\nFrom equations (10) and (1 1). and using the fact that infinitesimal rotations are commutative, we have that,\nThen, from equations (10) and (11). t o solve for the ~ ~ ( 0 : ) = R T ( o ~ ) , or ( 0 6 ) = -(d:). hinge joint.. we have :\nor. 0: = afan2(R:( 'L.1) . R:cl., )). where the indices' denoke t>he elements of the rotation mat,rix Rf..\nThe desired variation ax' and Ay' are extracted from Rf.R, (6;) = R.r(axi)R,(ayi). The computed endpoint location is then given, along the hinge base frame (frame -1) by :\nAxZ = [az\" ay'. dx ' . dy' dz']T (13)\nThe equations (10) and (13) imply that everything is calciilated with respect to the hinge -1 frame. Therefore. to carry on t,he calibratS(6)!>S(1))/(X(6)!X(1)). These values along with the positions of the centres of mass are used to calculate the position of the ends of the rigid spars and thus C\u00b8(i), etc. The components of the lashing forces in the x, y and z directions are proportional to the variables C\u00b8(i), S\u00b8(i) and \u00b9\u00b8(i) for the left-hand lashings and for the right-hand lashings CR(i), SR(i) and \u00b9R(i). The length of the left-hand lashing is given by \u00b8\u00b8(i)2\"S\u00b8(i)2#C\u00b8(i)2#\u00b9\u00b8(i)2 and the force is given by F\u00b8(i)\"(\u00b8\u00b8(i)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000601_920406-Figure2-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000601_920406-Figure2-1.png", + "caption": "FIGURE 2", + "texts": [ + " The performance is expressed in terms of decay rate. Impedance Test: In this method, the damping material is bonded to a 300 mm x 30 mm x 0.8 mm thick steel panel. The panel is excited usin6 a shaker, and the response of the test panel is measured using an impedance head (Figure 1). The damping performance is expressed in terms of loss factor. Special Test Method 1 (5): A test panel consisting of sheet metal of body panel thickness and visco-elastic material is excited mechanically by a stiff frame and shaker configuration (Figure 2). The performance is expressed in terms of loss factor. Special Test Method 2: The damping material is bonded to a 300 mm x 30 rnm x 0.8 mm thick steel test panel and supported along the nodal line. The test panel is then excited at the resonance frequency. The damping performance is computed using the decay rate technique (Figure 3). Special Test Method 3: This is a slightly different version of the Special Test Method 2. The damping material bonded to a 178 mm x 80 mrn x 0.8 rnm thick panel is freely sup orted on the nodal line and the damping P per orrnance is expressed in terms of decay rate at its resonance frequency (Figure 4)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002726_iros.1994.407422-Figure3-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002726_iros.1994.407422-Figure3-1.png", + "caption": "Figure 3: Modeling a 3D revolute mechanism.", + "texts": [ + " The terrain 7 is initially described using an elevation map defined from a regular grid ( E , y) expressed in the 58 1 set of connectors c, and a set of specific points selected on the rigid bodies R, corresponding to the joint 6k. The connectors are defined in terms of visco-elastic laws (i.e. combination of springs and dampers). For instance, a 3D rotoid mechanism is represented by two rigid bodies connected through two pairs of points respectively selected on them and belonging to the rotation axis. This construction is illustrated by figure 3 in the case of the articulation between two axles of the robot. 3.3 The Physical Model of the Terrain Dealing with robot/terrain interactions requires to build a representation @(7) of the terrain which is able to capture both the geometric and the physical properties of 7 and which allows the formulation of such interactions. For that purpose, we have implemented a model based upon the concept of \u201cphysical models\u201d, initmially proposed for computer graphics applications (see [14]). According to this concept, the terrain is represented by a set of interconnected particles @(Pi) having the following properties [lo]: (1) each particle is seen as a point mass m which obeys Newtonian dynamics and which is surrounded by a spherical non-penetration \u201celastic\u201d area; (2) the set of particles corresponds to the mass, the inertia, and the spatial occupancy characteristics of the modeled object; (3) the particles are interconnected using interaction components (refered to as the \u201cconnectors\u201d)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000329_jsl.3000100103-Figure3-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000329_jsl.3000100103-Figure3-1.png", + "caption": "Figure 3 Schematic diagram of high-temperature optical rig", + "texts": [ + " ( 5 ) and (6) gives an expression for the oil film thickness alone: where A,, and Nsp are the measured wavelength of maximum constructive interference and corresponding fringe order for the spacer layer alone, without JSL IO-I 32 0265-6582 $7.00 + $2.5ff The Film-Forming Properties of Polyalkylene Glycols 33 oil. hsp+oil and Nsp+oil are similar values when oil is present. nail is the refractive index of oil in the contact. All these values can be measured. Film thickness measurement test details Figure 3 shows a diagram of the test apparatus used in the current study to measure film thickness. The flat surface of a glass disc was loaded and rotated in nominally pure rolling contact against a 190 mm diameter AISI 52100 steel ball. The disc surface was coated with a 20 nm thick chromium layer and, for the ultra-thin film interferometric tests, a silica spacer layer of thickness approximately 500 nm was deposited on top of this. The load applied was 53 N, corresponding to a maximum Hertz contact pressure of 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002529_00037-1-Figure15-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002529_00037-1-Figure15-1.png", + "caption": "Figure 15 Interlaminar fracture toughness test methods for composite materials.", + "texts": [], + "surrounding_texts": [ + "generated to date. As for biaxial loading, either a tube or a cruciform specimen can be used.\nFor a tube, the internal and/or external pressurization provides a compressive normal stress in the third (interlaminar) direction. If these pressures are not equal, the interlaminar normal stress will not be uniform through the thickness. Also, these pressures must be relatively high if they are to induce an interlaminar normal compressive stress of any significance. It is not practical to achieve significant tensile stresses (e.g., applying a vacuum would induce insignificant interlaminar normal stress).\nThe internal and external pressures in turn induce proportional circumferential stresses, just as for biaxial testing. By varying the ratio of internal to external pressure, the desired level of circumferential stress can be achieved. However, the interlaminar normal stress will vary from a stress equal to the internal pressure on the inside surface of the tube to a stress equal to the external pressure on the outside surface. If the internal and external pressures are equal, the induced circumferential stress will be zero. Thus, the versatility of the tube test for triaxial loading is somewhat limited.\nAs an alternative, a three-dimensional cruciform specimen can be used, i.e., the specimen has pairs of arms protruding in three perpendicular directions. Any combination of tension and compression loading can be applied to the three pairs of arms, thus permitting the generation of a totally general stress state.\nThe usual data presentation format is a three-dimensional plot of axial normal (s1) stress vs. transverse normal (s2) stress, vs. interlaminar normal stress s3. If data for all eight octants of the plot are generated, the resulting failure envelop is a three-dimensional surface, every point on this surface representing a combination of s1, s2, and s3 that will cause failure.\n5.06.9.3 Strain Measurement Instrumentation\nIt may not be necessary to measure strains since only a stress failure envelope is to be plotted and the stresses can be calculated from the known magnitudes of the applied forces and/or pressures. However, if strains are to be measured, for biaxial testing a two-element biaxial strain gage rosette can be used, on either the tube or the cruciform specimen surface.\nStrain instrumentation becomes more difficult for triaxial testing. When testing a pressurized tube it is difficult to measure the interlaminar (through the wall thickness) strain. It is the change in tube wall thickness that must be determined. Capacitance and proximity gages are two possibilities.\nThe triaxial cruciform specimen has all six sides of its cubical gage region covered by loading arms. Thus, there are no exposed surfaces for the attachment of strain gages. Thus, capacitance or proximity gages, or similar devices, must be used in all three directions. There is much development work remaining in this area.\n5.06.10 FRACTURE MECHANICS TEST METHODS\nThe combined stress-failure criteria discussed in Section 5.06.9 predict failure based upon the local two- or three-dimensional stress state at a point. In contrast, fracture mechanics predicts failure based upon the stress intensity or strain energy density at the tip of a preexisting crack in the material being sufficient to cause that crack to grow.\n5.06.10.1 Historical Introduction\nFracture mechanics was initially developed for metals and similar isotropic materials in the 1940s, assuming linear material stress\u00b1strain response. Thus, this early work was termed \u00aaLinearly Elastic Fracture Mechanics,\u00ba or more commonly, simply LEFM. Later these works were extended to incorporate inelastic material response.\nWhen composite materials became prominent two to three decades later, initial attempts were made to use the same LEFM concepts to predict their failure as well. Major efforts in this regard were made during the early 1970s. Perhaps what is the obvious was eventually realized, viz., composite materials do not exhibit the same failure modes as homogeneous materials such as metals. For example, a preexisting crack in an isotropic material will often propagate in a \u00aaself-consistent\u00ba mode, i.e., it will", + "continue along the plane of the initial crack, in the same direction. Cracks in a composite material, because of the nonhomogeneous (fibers in a matrix), orthotropic, and often layered nature of the composite, propagate in much more complex and varied modes. As a result, fracture mechanics as applied to composites gained a very poor reputation and tended to\nbe ignored by the composites community for some time.\nWith the surge in the introduction of new types of composite materials in the early 1980s, and the corresponding desire to identify the tougher of these composites, as discussed in Section 5.06.1.4 there was renewed interest in applying the basic principles of fracture mechanics (O'Brien et al., 1987). Fortunately, the prior mistakes were not repeated, and current fracture mechanics efforts have a much sounder basis.\n5.06.10.2 Fracture Mechanics Principles\nMaterials in general are assumed to fracture in one of three modes, viz., Mode I\u00d0the opening mode, Mode II\u00d0the shearing mode, or Mode III\u00d0the tearing mode, or some combination of two or all three of these modes. The basic fracture modes are indicated in Figure 14. They are always designated by Roman numerals, as indicated above.\nThe fracture mechanics specimen configurations in most common use for composite materials at the present time include the Double Cantilever Beam (DCB), the End-Notched Flexure (ENF), and the Mixed Mode Bending (MMB) delamination test. These specimen configurations and the corresponding methods of loading are indicated in Figures 15 and 16. The DCB is a pure Mode I test, the ENF a Mode II test, and the MMB a mixed Mode I and Mode II test. In all cases the dominant crack propagation is parallel to the reinforcement, or the plies if it is a laminated composite.\nThe mixed-mode bending apparatus (Crews and Reeder, 1988) is capable of inducing any", + "prescribed ratio of Modes I and II loading, including pure Mode I or Mode II.\nThe various fracture mechanics test methods each require their own specific data collection and reduction procedures, which need not be presented here. The interested reader is referred to an excellent and concise presentation by Carlsson and Pipes (1987).\nOf primary interest for composite materials is the Mode I fracture toughness (ASTM D 5528, 1994). Much less work has been done with Mode II testing, and very little with Mode III, although an Edge Crack Torsion (ECT) specimen is currently under examination by ASTM (Li et al., 1997).\n5.06.11 NONAMBIENT TESTING CONDITIONS\nA brief caution about the low rate of moisture absorption and desorption of moisture in polymers and polymer\u00b1matrix composites, and the problems this can cause, was given in Section 5.06.3.1.1. This is one major problem. However, there are also other potential problems, and corresponding precautions to be taken when conditioning and testing composite materials.\n5.06.11.1 Common Problems/Precautions\nAs previously noted in Section 5.06.3.1.1, once moisture enters the surface of a \u00aadry\u00ba\npolymer (or a polymer\u00b1matrix composite), then even if the environment is soon returned to a \u00aadry\u00ba state, e.g., zero percent relative humidity (0% RH), half of the moisture in the surface layer will continue to propagate inward. Assuming that moisture is entering the material from all surfaces of the composite, then only when two moisture propagation fronts meet will the moisture reverse direction and propagate back out of the material. In even a relatively thin composite (e.g., 2mm thick) at room temperature, the total dry-out process can thus take weeks or months, depending upon the type of polymer, i.e, its moisture diffusivity (Springer, 1981).\nOn the other hand, since moisture absorption in most polymers is slow, brief exposures to high humidity air or liquid water result in relatively little, perhaps a negligible amount of, moisture weight gain. Thus, e.g., there is usually no cause for concern with the brief exposures to water during specimen cutting operations.\nMoisture diffusivity is a strong function of temperature, and therefore moisture absorption can be minimized by keeping the exposure temperature low. Correspondingly, absorption and dry out can be hastened by elevating the exposure temperature. The same general principles apply to other absorbing fluids also, e.g., gasoline, hydraulic fluid, and solvents in general, assuming of course that the fluid does not chemically attack the polymer.\nBoth moisture and temperature reduce the stiffness of the polymer, and usually the strength as well. The possible exception is when the polymer is very brittle in the room temperature, dry condition. Then moisture and/or an increase in temperature can soften the polymer, making it less sensitive to local stress concentrations, and thus making it appear to be stronger. Correspondingly, the matrix-dominated properties of the composite, e.g., transverse tensile strength and shear strength, can also increase.\n5.06.11.2 Elevated Temperature Testing\nThe two principal problems encountered in elevated temperature testing are gripping of the specimen if necessary and strain instrumentation.\nAdhesives tend to soften with increasing temperature. Thus, a tab adhesive that performs well at room temperature may not hold at elevated test temperatures. It is then necessary to switch to a higher use temperature polymer. Most high performance epoxy adhesives are serviceable up to temperatures of 120\u00b1140 8C." + ] + }, + { + "image_filename": "designv11_65_0002634_icnn.1994.374601-Figure7-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002634_icnn.1994.374601-Figure7-1.png", + "caption": "Figure 7 : BIT-CLOS guidance", + "texts": [ + " In this manner the missile will be gathered onto the line of sight and then kept there until impacting with the target A guidance controller at the ground station uses tracker i n f o d o n about the angular velocity and acceleration of the line of sight to provide demanded accelerations for the missile. These can then be transmitted to the missile via a radio l i . Here it is assumed that bank-to-rum, (BIT), control is employed to manoeuvre the missile. Thus the ailerons arc employed first, to roll through to the desired angle, with the elevators then used to accelerate towards the line of sight, as shown in figure 7. In practice both ailerons and elevators arc employed simultaneously, leading to nonlinear cross-coupling, (Roddy, 1985). Likewise B'IT control complicates the calculation of missile position as a highly nonlinear transformation to inertial axes is now required. By the application of the standard Euler equations, and following the assumptions adopted by Roddy, (1989, such as constant mass, small attack angle a and small sideslip angle thm the following sixth order linear model can be developed. a = Z,a + q - pJ3 + Z,6 q = Maa + Mqq + Mas P = Y,P - r + p n a i = NBp + N,r + = P" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001713_1.1317233-Figure4-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001713_1.1317233-Figure4-1.png", + "caption": "Fig. 4 The universal joint", + "texts": [ + " 2European Commission Directorate General XII MAST III Contract: MAS3CT95-0024. \u00a9 2000 by ASME Transactions of the ASME 7 Terms of Use: http://www.asme.org/about-asme/terms-of-use Downloaded F means of two plates which constitute the top and bottom of the finger. The plate at the basis is considered fixed with respect to the finger support while the plate at the top can freely move and it constitutes the site for installing either \u2018\u2018nails\u2019\u2019 or other kind of tips. The central points of the two plates are connected through a rigid link articulated by a universal joint ~Fig. 4!, also called cardan joint. The fingertip is allowed to span a portion of a spherical surface and it moves driven by the length of the three bellows. Whenever the internal pressure of a single bellow is different from the external one, a force proportional to this difference acts on the bellow itself, thus a proportional length variation can be observed. In this sense the movements are driven by a hydraulic system which pumps oil within the bellows: a set of linear motors operates on a relevant set of control bellows which are connected one rom: http://dynamicsystems" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001951_s0389-4304(01)00102-3-Figure3-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001951_s0389-4304(01)00102-3-Figure3-1.png", + "caption": "Fig. 3. Power transmission principle.", + "texts": [ + " Input and output discs, which receive the thrust generated by the loading cam, and the two power rollers of each variator sandwiched 0389-4304/01/$ 20.00 # 2001 Society of Automotive Engineers of Japan, Inc. and Elsevier Science B.V. All rights reserved. PII: S 0 3 8 9 - 4 3 0 4 ( 0 1 ) 0 0 1 0 2 - 3 JSAE20014345 between the discs transmit driving force by means of a special traction oil. This traction oil develops exceptionally high shear resistance under a condition of high contact pressure to accomplish power transmission without any metal-to-metal contact (Fig. 3). This CVT has a dual-cavity design that combines two sets of input/output discs and power rollers. Since it has double the number of contact points between the discs and power rollers that transmit driving force, it is capable of handling higher levels of torque. In addition, the two variators are positioned symmetrically to the center of the output gearset, making it possible to cancel out the opposing thrust forces that act on the discs. Coupling the two input discs via the shaft forms a single thrust generation mechanism that allows equal thrust force to be applied to both variators [2]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000095_0957-4158(94)e0025-l-Figure14-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000095_0957-4158(94)e0025-l-Figure14-1.png", + "caption": "Fig. 14. Perspective view of a piezoelectric single stage amplifier device producing 200/an displacement.", + "texts": [], + "surrounding_texts": [ + "microprocessor is invariably seen as a marketable attribute and for it to fulfil its marketing potential it must show its presence by providing a panoply of \"extra functions\" or customer definable choices. So far it has been rare for manufacturers to consult customers and users as to what features they actually want. Some products have reached levels of complexity where the users choose to ignore almost all of their enhanced features--most telephones are now capable of remembering ten or more telephone numbers; how many ever get taught even one? We still have a long way to go in developing user friendly systems, partly, I suspect, because of the point of sale appeal of \"more features\". Purchasers have long had difficulties in assessing the inherent quality of mechanical and electro-mechanical products. Industrial designers have sought to convey the qualities of products by styling, often using forms and motifs borrowed from elsewhere; their aim being to imply the underlying quality of a product which the purchaser can only assess from its exterior appearance. This problem of conveying product quality can be even more extreme with a microprocessor based product. Much of the quality inherent in a microprocessor-embedded product may exist only in the quality of its programming-directly visible to no-one but the most determined practitioners of \"reverse-engineering\" and appreciated indirectly only once the user is familiar with the operation of the product--usually long after the point of purchase. Obvious external manifestation of the processor's presence--the \"feature glut\" -- therefore takes on a formidable marketing importance. Persuading the customer that \"less-is-more\" is difficult, but there are signs that simplicity may be becoming a sales asset as customers for microprocessor-embedded products become more seasoned. A different and more serious problem lies in the relationship between mechatronic devices and humans. As our machines become more capable they appear more \"intelligent\". Some of the operations they can perform would stretch the mental capabilities of most humans, their response times are often quicker than ours and their sensory capabilities can, admittedly in limited ways at present, be superior. It is, therefore, sometimes very difficult for us to decide whether such machines are in complete control and performing correctly, and conversely when they should not be trusted! Even when functioning correctly from a programming point of view the most impressive mechatronic systems currently only have a very limited sensory perspective. They can, therefore, be quite unaware of problems or fault conditions which a human being could hardly fail to notice. How many mechatronic cameras, for example, are quite happy to auto-focus on the inside of the aircraft window rather than the interesting scene you wanted to record, 10,000 metres below? When the \"smart\" photocopier reports that it has detected a paper feed problem part way through a long double-sided, collated print run should you believe its confident assertion that it can sort out the problem? Clearly, major safety implications are possible in some other applications and these are not necessarily related to \"failure\" modes, only, perhaps, to estimates of appropriate behaviour based on insufficient sensory data. Humans make these kinds of mistakes too; we may need to incorporate some aspects of human factors and the psychology of perception into their programming. The \"cognitive\" potentialities of our machines have been dramatically increased in the past decade by rapid developments in the fields of microprocessor architectures Millwrights to mechatronics 113 and programming techniques. Almost as startling have been the improvements in sensory systems, many of them based on visual imaging, which are now enabling machines to experience our world more fully. The area which now appears to be lagging behind is the one which 20 years ago appeared the most advanced--the provision of suitable actuator technologies. For many years almost all electro-mechanical actuation techniques have been electromagnetic; electric motors, electromagnets and solenoids have usually been the only options. We need new actuator technologies if our mechatronic devices are to be as dextrous and as energy efficient as their potential applications demand. Some progress has been made in this area. One possibility lies in the application of piezoelectric materials [18]. These offer the attractive possibilities of generating large forces very rapidly, and can be highly energy efficient--requiring virtually no power to maintain a holding force, for example. Unfortunately currently available piezoelectric actuators require relatively high operating voltages (typically 100 V or more) and produce minute displacements. Research and development will soon produce a new generation of devices with significantly lower operating voltages. The author and co-workers are investigating efficient ways of mechanically amplifying the output displacements of piezoelectric actuators to enable their use in the faster, more 114 T. G. KING dextrous mechatronic systems of the future. Specific designs and appropriate design techniques have been produced [19-22] and work is ongoing. Figures 14 and 15 illustrate two of these mechanically amplified actuators, based on flexure hinged structures, whilst Fig. 16 shows a piezoelectrically actuated motor concept, currently being researched, capable of very fine increments of rotation, again employing flexure hinged displacement amplifying techniques." + ] + }, + { + "image_filename": "designv11_65_0001031_jsvi.1997.1141-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001031_jsvi.1997.1141-Figure1-1.png", + "caption": "Figure 1. Front view of suitcase with non-dimensional excitation moment and restoring moment.", + "texts": [ + " (2) The loss of energy when one of the wheels impacts the ground is modelled with the use of a coefficient of restitution e (0Q eQ 1), so that the angular velocities ( du/ dt)\u2212 and ( du/ dt)+ just before and just after impact, respectively, are related by the equation ( du/ dt)+ =e(du/dt)\u2212. (3) For the particular suitcase under consideration, the fixed parameters are chosen as I=3\u00b784 kgm2, Mb =20\u00b72 kgm2/sec2, Mh =81\u00b73 kgm2/sec2, and e=0\u00b7913 [1]. The following non-dimensional quantities are introduced: t= t(Mh /I)1/2, g=Mb /Mh =0\u00b7248, A= q0 /Mh , V=v(I/Mh )1/2, b= k0 /Mh , d=D(Mh /I)1/2. (4) Then equation (1) becomes d2u(t)/ dt2 + sign (u(t))g cos u(t)\u2212 sin u(t)+ bu(t\u2212 d)=A sin (Vt+ h). (5) The non-dimensional excitation moment and restoring moment are depicted in Figure 1. It is assumed that the suitcase is vertical (u=0) before t=0. In non-dimensional terms, the weight of the suitcase furnishes an initial restoring moment g about each wheel of the suitcase (i.e., when u=0). Therefore rocking will only occur if the periodic excitation moment exceeds this value at some times, i.e., if Aq g, and only this range needs to be considered. The phase h will be chosen such that the initial applied moment A sin h is equal to g, so that motion will begin at t=0. For this value of h, for the values of g and e listed earlier, and for specified values of b, d, A, and V, equation (5) is integrated numerically for 20 cycles of the excitation moment, i" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003454_1.1757489-Figure2-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003454_1.1757489-Figure2-1.png", + "caption": "Fig. 2 A sketch of slider-disk assembly and Cartesian coordinates", + "texts": [ + " Here the first order slip conditions @7# are chosen because an analytical solution for the air velocities can be obtained for an air bearing. In this model we have @4,5# uguz505U1l ]ug ]z U z50 (1) uguz5h52l ]ug ]z U z5h (2) vguz505V1l ]vg ]z U z50 (3) vguz5h52l ]vg ]z U z5h (4) where h is the local height of the air-bearing; U, V are the speeds of the disk in the x and y-directions respectively, and l is the mean free path of the air. The Cartesian coordinate system used in this air bearing analysis is illustrated in Fig. 2. Using these boundary conditions, we obtain for the non-dimensional form of velocity components of the air 746 \u00d5 Vol. 126, OCTOBER 2004 rom: http://tribology.asmedigitalcollection.asme.org/ on 08/30/2017 Term Ug5 P0 2rgU2 hm l Reh ]P ]X ~Z22ZH2KnhH !1S 12 Knh1Z 2Knh1H D (5) Vg5 P0 2rgU2 hm l Reh ]P ]Y ~Z22ZH2KnhH !1 V U S 12 Knh1Z 2Knh1H D (6) where P is the dimensionless pressure, or pressure divided by the ambient pressure P0 ; Z5z/hm is a nondimensional coordinate; H5h/hm is a nondimensional spacing of the air-bearing; Reh 5rgUhm /m is the Reynolds number; Knh5l/hm is the Knudsen number defined in terms of the nominal spacing height hm ", + " The particles are first assumed to be uniformly distributed above the disk surface with velocities close to the air-bearing\u2019s velocity where the particles are located, as determined by Eqs. ~5! and ~6!. The particle\u2019s initial vertical velocity is assumed to be zero. In this section the three-dimensional airflow effects on particle trajectories and contamination are studied. A representative mod- rom: http://tribology.asmedigitalcollection.asme.org/ on 08/30/2017 Term ern negative pressure slider is chosen for this study ~Fig. 1!. The flying height of the slider is 26 nm. A sketch of the slider-disk assembly and coordinates is given in Fig. 2. The pressure profile shown in Fig. 3 is obtained without three-dimensional effects by solving the Reynolds equation for the air-bearing of the slider using the CML code Quick 4. To study particle flow in the air-bearing, we first calculated the spacing function between the slider and the disk. The slider-disk spacing map is shown in Fig. 4, where it can be observed that particles may enter the recessed region of the air-bearing through the leading edge of the slider. Next, the air streamlines at different levels of the slider/disk spacing were calculated using Eqs" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001136_s0020-7403(98)00106-4-Figure3-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001136_s0020-7403(98)00106-4-Figure3-1.png", + "caption": "Fig. 3. Simpli\"ed motion of the ship.", + "texts": [ + " That is Zin\"XG sin(b)sin(u p t) . (b) A horizontal linear motion, Xin, (i.e. surge) in the X direction, equivalent to the horizontal movement of the deck due to the pitching of the ship. That is Xin\"!ZD sin(b)sin(u p t). (c) A horizontal linear motion, >in, due to the swaying of the ship. (d) A rotation of the axes ox@y@z@ (\"xed in the body of the ship) about an axis parallel to the OX axis (where OX>Z are the \"xed global axes) and equal in magnitude to the roll angle of the ship. This simpli\"ed ship's motion is shown in Fig. 3, where OX>Z are the \"xed axes and ox@y@z@ are the moving axes which are \"xed with respect to the ship. The axes oxyz and ox@y@z@ are \"xed relative to each other. Also, so that the relationship between the pitch and roll motions of the ship could be investigated, the expression for the pitch angle, used in the mathematical model, is given by h p \"b sin(u p t#'), where ' is the phase angle between pitch and roll. The relationship between ox@y@z@ and OX>Z is given by x@ y@ z@ \" 1 0 0 0 cos a sin a 0 " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001830_2013.8860-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001830_2013.8860-Figure1-1.png", + "caption": "Figure 1. Structure of the free play of the three\u2013point hitch.", + "texts": [ + " In order to achieve the objective, we have attempted both numerical and experimental approaches. Numerical simulation was conducted with the tractor vibration model developed for a small\u2013sized paddy field type tractor by Sakai et al. (1988). An experimental investigation was conducted with an 11 kW paddy field type tractor. An ISEKI (14 kW) tractor, both with and without a rear\u2013mounted implement, was run over an artificial profile. When the implement was hitched to the tractor\u2019s three\u2013point hitch, free play was allowed in the three\u2013point hitch linkage system (fig.1). Vertical acceleration and link force data were measured and recorded using a 7\u2013channel data recorder. The experimental apparatus was set up as shown in figure 2. In the without\u2013implement condition, the implement was disconnected from the tractor\u2019s three\u2013point hitch, so free play did not affect the tractor ride vibration. In the with\u2013implement condition, the implement was hitched to the tractor, and there was free play in the three\u2013point hitch. As is typical during transport, the three\u2013point hitch was in the fully raised position, which still allows free play" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002866_095440603322769938-Figure3-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002866_095440603322769938-Figure3-1.png", + "caption": "Fig. 3 Design of a rack cutter with ring-involute teeth for generating the gear and pinion", + "texts": [ + " To obtain the rack cutter surface, the normal section of the ring-involute rack cutter is attached to the plane yn \u00a1 zn, rotated along the zn axis and translated along the y1 axis with respect to the coordinate system S 1\u2026O1, x 1, y1, z1\u2020, as shown in F ig. 3. Thus, the normal section of the rack cutter surface can be represented by the coordinate system S 1\u2026O1, x 1, y1, z1\u2020 by applying the homogeneous coordinate transformation matrix method. The variables b and `h are parameters of the rack cutter surface, where `h represents the tool cutting depth design parameter of the normal section in Fig. 3 and the range of b is \u20300, 2p\u0160 in order to create the rack cutter ring-involute teeth surfaces. Region ab of the rack cutter surface can be considered to generate the bottom land of the gear with ringinvolute teeth (convex teeth). In terms of the gear and region ef of the rack cutter surface, the addendum surface may be taken as the shape of the gear blank. Proc. Instn Mech. Engrs Vol. 217 Part C: J. Mechanical Engineering Science C02203 # IMechE 2003 at Purdue University on July 22, 2015pic.sagepub" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003092_cdc.2001.980683-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003092_cdc.2001.980683-Figure1-1.png", + "caption": "Fig. 1. Flatness control actuators of the 6-high UC cold rolling mill.", + "texts": [ + " Mill Model for Flatness Control There are many mills that are designed to perform the flatness control. In this chapter, the Universal Crown mill is introduced for modeling the flatness actuators. 2.1. Flatness Control of Universal Crown Control Mill The Universal Crown Control (UC) mill is designed to control not only the gage of the strip, but also the flatness. To perform the flatness control, it is equipped with workroll bender, intermediate-roll bender, intermediate-roll shift and skew as shown in Fig. 1. There are two types of flatness defects, known as the global flatness defects and the local flatness defects. A roll bender controls the symmetrical components in the global flatness defects. When the bending force in the roll bender increases, the rolls push together in center so that the strip flatness becomes loose in center and tight in edges. The strip flatness due to the difference in the roll gaps between roll edges manifests loose edge at the closer-gap side and tight edge at the wider-gap side" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001966_ijmtm.2002.001439-Figure27-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001966_ijmtm.2002.001439-Figure27-1.png", + "caption": "Figure 27 RRT, a NHTSA component test device (see online version for colours)", + "texts": [ + " Most of the methods that have been reported in the literature are quasi-static tests performed by rotating the occupant by 180\u00b0 in a test fixture around a stationary axis, thus addressing only the rotational phase of vehicle motion without taking the lateral translational motion into consideration. These methods have been developed for different objectives, including studying occupant excursion, restraint system effectiveness evaluation and development, and repeatability. These methods are briefly reviewed in Sections 3.1\u20133.3. Rollover restraint tester (RRT) is a component test device developed by NHTSA (Rains et al., 1998), as shown in Figure 27(a) and (b). Figure 27(a) shows a schematic of the fixture, while Figure 27(b) depicts the actual hardware. RRT is a \u2018spit test\u2019 type of device that is capable of generating a free flight motion in the airborne phase of a rollover event. A drop tower and free-weight system are used to provide the driving force for the RRT. The angular velocities of RRT, ranging from 180 to 290\u00b0 s 1, are attained using various combinations of drop weight and drop height. The main features of the RRT are composed of a supporting framework, a counter-balanced test platform with a pivot axle, a free-weight and drop tower assembly, and a shock tower, as shown in Figure 27(a). Dummy head movement can be digitised by tracing tape markers on the head with respect to a reference grid behind the dummy\u2019s head. In 1997, Moffatt et al. (1997) also developed a very similar device called head excursion test device for studying head excursion of seated cadavers, volunteers and Hybrid III dummies. Source: Rains et al. (1998). This fixture, as shown in Figure 28, was developed by Key Safety Inc. (formerly Breed) and includes only a portion of the occupant compartment to simulate a quarter turn roll with no free flight motion" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000406_105971239900700201-Figure3-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000406_105971239900700201-Figure3-1.png", + "caption": "Figure 3. The time profile of the error function E2 (0) = sin 20 - Kid . The error function became zero within 10", + "texts": [], + "surrounding_texts": [ + "142\nthe phase difference, 0 = i - 0. o (a) Learning mode (first 10 seconds). Two oscillators were synchronized, and\nthe desired phase relation Od = 0.8 was acquired within a few cycles by learning. (b) Recalling mode (10-20 seconds). At t=10 [s], the learning was stopped, and the learned phase difference was immediately regenerated from a random initial phases.\nis obtained when the dynamics of the CPG can be transformed to the Poincare\u2019s normal form for Hopf bifurcation and the attraction to the limit cycle is strong (Nishii et al.,1994). The dynamics of the robot is shown in the appendix. .3..~ ~ Learning a desire d h opping ~e~~~~,\nFirst, the simulation results for learning a desired hopping height are shown. The error function is given by:\nwhere x is the desired height and x;, is the time-aver-\naged height of the trunk of the robot .x~; that is:\nwhere z = 5.0/ OJ is the time constant. The learning rule\n(13) is applied, and the time-averaged function Ri is\nobtained by:\nFigure 5 shows the sumulation results. One period of hopping is about one second in a steady state. No learning was carried out in the first ten seconds so that a steady relation between the CPG and the robot could be achieved. The desired heights, Xd = O.G, 0.7, 0.8, 0.9 ~m), were obtained within 100 seconds. Figure 6 shows the time course of the hopping height after the acquisition of the desired parameters for x j = 0.9 [m], which indicates that the learned height was immediately acquired. In this simulation, learning failed for larger and smaller target heights (x~ <_ 0.5 and x~ d 2:: 1.0) and for ini-\nat UNIV OF CALIFORNIA SANTA CRUZ on April 2, 2015adb.sagepub.comDownloaded from", + "143\nseconds by learning for each parameter, Kd = -0.4, 0.3,1.0. The parameters and initial condition are the same\nas those shown in Fig.l.\ntial parameters which desynchronized the CPG and the robot.\n3.2.2 Learning a maximum hopping height.\nAlthough it was shown that the desired parameter set can be learned by an error function, it is difficult to obtain the maximum value allowed by the physical system, such as maximum hopping height, because the desired value is unknown in such a case. Even if the\ndesired value is known, the appropriate parameter set might not be obtained by the error function given by the error between the actual value and the desired value\nbecause such an error function does not satisfy Condition 3 in the theorem.\nIn these cases the learning can be done by setting the error function as:\nwhere c is a positive constant and (0) is the variable which we want to maximize. With this error Fanction, g becomes maximum if jZ\u2019(0) changes its sign twice in\n0 r= S like the function sin 2TC~ . Based on this consideration, we set the follow-ing error function instead of Equation 21 to obtain the maximum hopping height in the simulation experiment of the adaptive control of the hopping robot:\nat UNIV OF CALIFORNIA SANTA CRUZ on April 2, 2015adb.sagepub.comDownloaded from", + "144\n~r~ = 0.6, 0.7, 0.8 0,9 were adaptively achieved. Parameters are given as (~, ~) = (0, 0.8); ~ = 0.3, and~ =1/3. As an initial condition, we put ~(o)=0.5,(~(o),~~(o))=(l.O,-l.o),<~(o)=1.0[Hz],~(o)=0.1[m], x 0 (0) = Xi (0) + I = 0.8 [m]. The lines look thick because of the small high-frequency changes in the height X-0 . o\nwhere L1\u00f8max is the change in the maximum height Xmax between each new hop from the previous hop, and L1\u00f8max is the change in phase difference between the robot and the oscillator since the previous hop. By defining the phase of the robot as zero when the robot reaches maximum height, \u00d8max is given by \u00d8max = - 0,,,~,,. , where emax is the phase of the oscillator at the moment. Small random values are added to the coupling weights during learning to avoid a stagnation in learning by the eventually obtained value 4x~~~ z 0 - Figure 7 shows the timeaveraged height of the robot X-0 , which indicates that ~y - 0.95 [m] was obtained by the learning. The obtained value was almost same as the maximum value obtained in the previous simulation (Figure 5). This value would be the maximum height allowed by this physical system.\nhopping height xd = 0.9 [m] was achieved.\nat UNIV OF CALIFORNIA SANTA CRUZ on April 2, 2015adb.sagepub.comDownloaded from" + ] + }, + { + "image_filename": "designv11_65_0003038_s0094-5765(01)00002-9-Figure3-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003038_s0094-5765(01)00002-9-Figure3-1.png", + "caption": "Fig. 3. SimpliAed system showing a space platform supporting one-module manipulator.", + "texts": [ + " With this as background, Part I of the paper develops a rather general, three dimensional, order N Lagrangian formulation for a class of novel variable geometry manipulators with modules of slewing and deployable links (Fig. 2). The Oexible manipulator is supported by an elastic platform in an arbitrary orbit. The snakelike manipulator has several advantages: It reduces coupling eIects resulting in relatively simpler equations of motion and inverse kinematics; decreases the number of singularities; and facilitates obstacle avoidance (Fig. 3). Dynamics and control of such Multi-module Deployable Manipulator System (MDMS), free to traverse an orbiting elastic platform and carrying a payload, represent a challenging task. A relatively general character of the model makes it applicable to a large class of systems. A number of existing space- and ground-based manipulators become particular cases of the general model developed here. Part II of the paper illustrates versatility of the formulation, through application to a variety of multi-module manipulators, revealing complex interactions between orbital, librational and vibrational degrees of freedom during manipulator maneuvers and other disturbances" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002623_a:1015265514820-Figure9-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002623_a:1015265514820-Figure9-1.png", + "caption": "Figure 9. Placement and detection regions for the robot sonars.", + "texts": [ + " These checked for obstacles such as stairs going downwards that may have posed threats to a wheelchair operating indoors. Another pair of sonars (s1 and s2 in Fig. 8) checked for obstacles in indoor environments directly in front of the robot and above ground level such as the flat surface of tables. The ultrasonics were clustered into eight groups corresponding to eight proximity sensors with different resolutions, located in a ring around the robot, with a sensor every 45\u25e6 (see Fig. 8). In the front there were a total of eight sonars (p1\u20132, n2\u20133, n4\u20135 in Fig. 9 along with s1\u20132 from Fig. 8) offering increased resolution in the three directions the chair most frequently moved; straight ahead, and up to 45\u25e6 to the left or to the right. There was one sonar for the remaining five directions. An aluminum frame that surrounded the chair was used for mounting all the proximity sensors. Five of the ultrasonics (p1\u20133, s1\u20132) worked in a fail-safe mode, thus increasing the reliability of the sensor readings in these directions. The absolute orientation of the robot was extracted using a pair of encoders attached to the inner side of each of the two back wheels of the chair", + " The width of doors and corridors in this case was approximately 1.2 times the width of the wheelchair in order for the robot to fit. A 60% success in a total of 15 trials was achieved in the scenario shown in Fig. 12. Most of the time, failure was caused in area A in Fig. 12. The main reason for the failure in this case was the limited spatial resolution of the chair in the front direction during closed turns, due to the limited accuracy of the sonars that were scanning this direction. In particular, sensors n1, n2, n3, n4, n5 and n6 in Fig. 9 had a minimum sensing distance of 32 cm. The reason for this relatively large lower bound was that each one of these sensors used a single membrane both as an emitter and as a receiver. Whenever the sensor emitted a pulse using this membrane, the membrane had to come to rest before it could start measuring the reflections and estimating obstacle distances. In all of the failures in area A the robot followed a turning trajectory that led to a collision with one of the side walls without attaining a minimum distance of 32 cm from these walls" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001579_0022-4898(87)90009-7-Figure8-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001579_0022-4898(87)90009-7-Figure8-1.png", + "caption": "FIG. 8. Relationship between the driving force of the sprocket and the distance between the tracks.", + "texts": [ + " Referring this figure, the thrust force for the outer crawler, Zo can be obtained as Mo : latPob(LoV2 + 2x'2)/2. Zo : latPob(Lo2/2 + 2x'2)/2B + Fo can be obtained. Driving force of the sprocket of the track. (6) (7) REDUCTION OF TURNING MOTION RESISTANCE 257 The above equations show that the thrust force Zo can be varied depending on the distance x' and the distance between tracks B. Also the minimum value of Zo, Min(Zo), can be obtained when x=0, where the point O 8 is located, and it is reduced with the increase of the distance between crawlers (tracks) B. Figure 8 shows the relationships between the thrust force Zo, the shifted displacement of the pivoting point x' and the distance between tracks B. In Fig. 7, Mo and M, can be expressed as follows. Mo = p~oLo2/4 = #t WoLo/4 (8) = Fo + #t (WoLo + W~LO/4B = F~-#t (WoLo + W~L i)/4B. It can be found that the thrust force Zi for inner crawler is related to the braking force and the turning resistance moment, and Zo is related to the rolling resistance in straight travel and the turning resistance moment. Let Fo be Fo' when the contact lengths Lo and L, can be shortened to Lo' and Li', respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000330_elan.1140081110-Figure2-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000330_elan.1140081110-Figure2-1.png", + "caption": "Fig, 2. Thin-layer cell and electrodes: 1) working electrode; 2) auxiliary electrode; 3) reference electrode; 4) circulation of the inert gas (Ar) or CO; 5) thermistdnce; 6) fritted glass disks; 7) junction of the thermistance.", + "texts": [ + " 1), which leads to a stable radical anion at room temperature [6]. This species was found to be an efficient oxygen probe. Data are presented on the stabilization and characterization at low temperature of reduced a-alkyl iron porphyrins which are unstable at room temperature. The compounds studied here are the o-C2HS-Fe\u201d compounds of Fe tetrapentafluorophenylporphyrin (FeTFSPP, Fig. 1) and those of Fe tetraphenyl-Poctochloroporphyrin (Fe/3C18TPP, Fig. 1). 2. Experimental 2.1. Description of the Thin-Layer Cell Figure 2 gives a diagram of the spectroelectrochemical design of the cell. Electrounulysis 1996, 8, No. 11 0 VCH Verlugsgesellschuft mbH, 0-69469 Weinheim, 1996 1040-0397/96/1111-1029 $ 10.00+.25/0 2.2. Instrumentation The auxiliary electrode was a platinum wire and the reference electrode was a Ag wire covered with AgCl in DMF containing NEt4CI M and NEt,CI04 0.2M ( E = -0.35V VS. SCE). Cyclic voltammograms were obtained using a home-built potentiostat and current measurer, a function generator (Parr 175) and a X-Y chart recorder (IFELEC 2502)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002335_20.877741-Figure7-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002335_20.877741-Figure7-1.png", + "caption": "Fig. 7. Flux distribution (discontinuous method, coarse mesh) D = 4:75.", + "texts": [ + " The meshes on the magnet and pole piece sides are generated separately and the meshes are put together at the arbitrary position . The meshes are divided uniformly in the -direction and the mesh sizes in this direction for the coarse and fine meshes are 1 mm and 0.5 mm respectively. Therefore, the coarse and fine meshes become the ordinary conforming meshes when is a multiple of 1 mm and 0.5 mm. The periodic boundary conditions are imposed on boundaries \u2013 \u2013 \u2013 and \u2013 \u2013 \u2013 and the Dirichlet boundary conditions are imposed on the boundaries \u2013 and \u2013 . Fig. 7 shows the flux distribution at mm obtained from the discontinuous method with the coarse mesh. Fig. 8 shows the change of the -component of flux density, , at the point s, shown in Fig. 5, with when the coarse mesh is used. The symbols and and in Figs. 8 and 9 denote that the results are obtained from the ordinary conforming meshes. The results obtained from the discontinuous and continuous methods in Fig. 8 are both reasonable, because they change smoothly even if the meshes become of the nonconforming type" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003864_50002-2-Figure2.9-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003864_50002-2-Figure2.9-1.png", + "caption": "Fig. 2.9. Top view of tyre contact area showing its position with respect to the system of axes (0 o, x o, yO, z o) fixed to the road and its deformations (u, v) with respect to the moving triad (C, x, y, z).", + "texts": [ + " In the following, a set of partial differential equations will be derived that governs the horizontal tyre deflections in the contact area in connection with possibly occurring velocities of sliding of the tyre particles. For a given physical structure of the tyre, these equations can be used to develop the complete mathematical description of tyre model behaviour as will be demonstrated in subsequent chapters. Consider a rotationally symmetric elastic body representing a wheel and tyre rolling over a smooth horizontal rigid surface representing the road. As indicated in Fig.2.9 a system of axes (O ~ x ~ y~ z ~ is assumed to be fixed to the road. The x ~ and y ~ axes lie in the road surface and the z ~ axis points downwards. Another coordinate system (C, x, y, z) is introduced of which the axes x and y lie in the (x~ O o, y o) plane and z points downwards. The x axis is defined to lie in the wheel centre plane and the y axis forms the vertical projection of the wheel spindle axis. The origin C which is the so-called contact centre or, perhaps better: the point of intersection, travels with an assumedly constant speed Vc over the (x~ O o, y o) plane", + " The angular deviation of the x axis with respect to the x~ axis (that is the yaw angle) becomes ~t - fl +~ (2.41) For the angle fl the following relation with y o, the lateral displacement of s i n f - d y ~ (2.42) ds occur when rolling on a frictionless surface), with coordinates (Xo, Yo), are indicated by u and v in x and y direction respectively. These displacements are functions of coordinates x and y and of the travelled distance s or the time t. The position in space of a material point of the rolling and slipping body in contact with the road (cf. Fig.2.9) is indicated by the vector p - c + q (2.43) where c indicates the position of the contact centre C in space and q the position of the material point with respect to the contact centre. We have for the latter vector: q - (x o + u) e~ + (Yo \u00a7 v) ey (2.44) with ex (= l) and ey (=t) representing the unit vectors in x and y directions. The vector of the sliding velocity of the material point relative to the road obviously becomes: - - x + ~ { ( X o + U ) e - (Yo+V)e~} (2.45) Vg =lJ d + (1 Vc + (~o+a) e + (feo+O)ey y where Vc = t~ denotes the vector of the speed of propagation of contact centre C" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000028_bfb0013964-Figure8-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000028_bfb0013964-Figure8-1.png", + "caption": "Fig. 8. Checking edge visibility", + "texts": [ + " When the propagation reaches a point p = (q, e) such that the position uncertainty shape grown by the robot radius R (circular approximation) is included in a landmark region, the uncertainty is reset to the value eL of the landmark and it remains constant for all the neighbors such that the disc of radius/~ + eL lies inside the landmark. Obstacle relocalization is checked each time a neighbor node is not collision free because of a too large uncertainty (ie. the current node (Ptl) lies on the boundary of the Free nodes). A visibility test is performed to verify that an edge can be reliably reached from the current position by a Move to Wall primitive (Fig. 8). When this test succeeds, the contact position is computed and labelled Edge, its parent node being the last developed. If a docking primitive starting from this Edge node is collision free, a child node is created corresponding to a robot configuration parallel to the reached edge. Edge nodes propagation is limited to sample directions given by obstacle edges. Finally when this propagation reaches a vertex or another obstacle, the uncertainty is reset to zero (or sensor accuracy) for the corresponding neighbor" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002946_abb-2003-9693532-Figure13-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002946_abb-2003-9693532-Figure13-1.png", + "caption": "Figure 13 Measuring points.", + "texts": [ + " During brachiation, it is almost impossible to measure the body position, eg the tip of the free arm or the centre of gravity of the robot, because the slip angle at the catching grip is not directly measurable using a potentiometer or rotary encoder. We therefore use the real-time tracking system Quick MAG System IV, which measures the three- dimensional locations of the eight points at 60 Hz sampling frequency, using two CCD cameras and coloured markers. The seven measuring positions shown in Figure 13 are chosen to approximately calculate the centre of gravity of the robot based on the following assumptions: (1) elbow of the grasping arm is kept straight; (2) both legs are controlled synchronously; and (3) two joints on the shoulder are adjoining and attached at almost the same position. The adaptation algorithm was applied to adjust locomotion, which indicates six activation coefficients ( )2 2 2 2 2 2 3 4 5 6 3 4, , , , ,r r r r t t to the corresponding four local behaviour controllers: leg stretch, body rotation 2, body lift and arm reaching" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001966_ijmtm.2002.001439-Figure13-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001966_ijmtm.2002.001439-Figure13-1.png", + "caption": "Figure 13 Deceleration rollover test sled: (a) roll side and (b) non-roll side", + "texts": [ + " (1989) presented this test procedure using passenger cars and utility vehicles, analysed vehicular kinematics results (including time histories of displacement, velocity and energy) and compared results of occupant responses with those from dolly tests at similar speeds with a similar vehicle. The deceleration rollover sled (DRS) test is a versatile methodology developed by Autoliv North America for rollover testing of a vehicle using a lateral deceleration sled (Rossey, 2001). The DRS consists of a sled and a set of deceleration pulse generating brakes, as shown in Figure 13. A test vehicle can be positioned laterally on the platform of the sled. On the leading side of the platform, a set of \u2018curbs\u2019 are bolted to the surface of the platform. The vehicular wheels are placed against the vertical edge of the curbs. The sled accelerates in the direction of the sled track until it reaches a predetermined speed. Brakes are then applied to decelerate the sled to a desired \u2018g\u2019 level. Combined brake position and brake pressure can be adjusted to generate different test conditions as shown in Table 2 given by Rossey (2001)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000835_ip-cta:19971028-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000835_ip-cta:19971028-Figure1-1.png", + "caption": "Fig. 1 Note: + ~k indicates the ship is 'below' the track. Using rudder to minimise tracking error", + "texts": [ + " Consequently, human knowledge and understanding is required to set appropriate training targets for the SISO NN controller (see Section 3) , so that overshoot and other undesirable characteristics can be lessened. The advantage of the SIMO method is that both track and heading can be controlled. Both approaches require specification of the track, but in the SIMO approach the desired ship heading on the desired track is also required. The more technical aspects of the differences between SISO and SIMO approaches to track keeping are now discussed. 2 Configuration of SISO and SIMO neural network controllers As shown, the SISO approach is based on using the rudder to maintain track, see Fig. 1. The schematic representation of the associated controller is provided in Fig. 2. rudder I - - - \\ \\ \\ - - - I I \\ \\ \\ 0 \\ \\ - _ I Fig.2 Schematic diagram of SISO controller The inclusion of the rudder limiter in Fig. 2 is based on the understanding that the command rudder signal generated by the neural network may not be physically realisable, because (a) the difference ISkc ~ Sk-ll is too large to be achieved in one time interval, (b) the actual rate of rudder operation is limited. To represent these physical limitations of rudder operation, we introduce a ramp threshold function, see Fig", + " Hence, &k defined by is very large. Here, At is simply the time interval between observations at time k and observations at time k-1. Furthermore, on appealing to the linearised Taylor series for 6, it follows that i k + 1 + (nt)&k = i k + ( e h - &&I) (15) Hence, because \\&k - Bk-11 >> IEk - &k-lI then 191 >> l&kl, and hence, in this case, there is a potentially large change in &k+1. The tendency of the ship movement, as 157 0, + 0, depends on whether the ship is \u2018above\u2019 or \u2018below\u2019 the planned route, see Fig. 1, and this is identifiable from the following arguments. Initially, we assume the ship is \u2018above\u2019 the planned route, that is sign(ykG - y k Q ) = + 1. As 8, and are both negative, then 6, + 0 requires that is more negative than &, i.e. 6, - is large and positive. Hence, from eqn. 15, we know that tends to become less negative or even positive. Here, we have identified a potential overshoot situation. Hence, to leave ckd at its nonzero value, when we have identified a possible overshoot situation, is not logical, and we should, therefore, use Ekd 4 0 for 0, - 0, with the ship in the \u2018above\u2019 track position", + " To avoid ambiguity in the learning process, we have found [ 181 that the correct sign of a gradient is more important than knowledge of the gradient magnitude. Hence, we advise replacement of d&k/ddk by sign(d&k/a6k) and aqkia6, by sign(dqkld&). Furthermore, see [14, 18, 191, we find that use of sign (2) = -1 and sign (2) - 3 is sufficient. This is consistent with the work of Saerens et al. [16]. An explanation for the existence of eqns. 30 and 31 is as follows. In the ship track-keeping problem, the track error is defined as &k, see Fig. 1. It is obvious that an increase of rudder 6, will reduce this error Ek, given the assumed positive sign convention of positive rudder changes being to port and positive track error being associated with a ship \u2018below\u2019 the desired track. A positive increase in rudder 6, will reduce the heading +k, with a positive q k defined in Fig. 4. Using eqns. 29-31, eqns. 25 and 28 are replaced by L 7 _ _ _ - _ - - _I t Fig. 10 160 SIMO neural control scheme IEE Proc.-Control Theory Appl., Vol 144, No. 2, March 1997 for the SISO NN controller, and for the SIMO NN controller", + " Throughout the SISO simula- IEE Proc-Control Theory Appl., Vol. 144, No. 2, March 1997 tions, the range of numerical values used to define the implicit range (NB, PB) is selected in accordance with Table 2, i.e. we have considered using different definitions of the ranges PB to NB for the three parameters E, i. and Ed. In the SIMO simulations qkd = d 2 . E s: -1 00 L - J u t , , , , , , , , , I I 0 1000 2000 3000 GOO0 5000 6000 x,m SISO NN controller in straight line course (no wind, no noise) Fig. 1 1 -.- reference track -Y- with loose PB and NB values -X- with medium PB and NB values -A- with tight PB and NB values -3001 , , , , , , , , , , , I . F E 50 5 , 30 I \u2019 D -30 z g 0 0 1000 2000 3000 LOO0 5000 6000 xtm Fig.12 noise) - .- reference track -Y- with loose PB and NB values -X- with medium PB and NB values -A- with tight PB and NB values SISO NN controller in straight line course (under wind, no I 0 1000 2000 3000 LOO0 5000 6000 x,m Fig.13 noise) -a- reference track -Y- with loose PB and NB values -X- with medium PB and NB values -A- with tight PB and NB values SISO NN controller in straight line course (under wind and In Figs" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002335_20.877741-Figure11-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002335_20.877741-Figure11-1.png", + "caption": "Fig. 11. Mesh of induction motor.", + "texts": [ + " A nonconforming technique is applied to a 2D magnetic field analysis in an induction motor in order to investigate its effectiveness . Fig. 10 shows the model of an induction motor without skew. The motor is one of the verification models proposed by the IEEJ. The detailed data of this motor is shown in Table I. The rotor and stator cores are made of the silicon steel 50A1300 and the nonlineality is taken into account. The conductivity of the aluminum rotor bar is assumed to be S/m by revising the conductivity to 2-D analysis. Only 1/2 of the whole region is analyzed due to symmetry. Fig. 11 shows the mesh with the first order quadrilateral elements. The nonconforming technique is applied in the gap between the rotor and the stator. The mesh is made fine enough to neglect the effect of the discontinuity of the potential. The total number of elements and nodes are 23 778 and 24 790, respectively. The discontinuous method is applied. The steady state is obtained using the step-by-step method. The time interval is chosen so that the rotor rotates at 0.88 deg. during . The periodic boundary condition is used" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002479_13640461.2002.11819467-Figure2-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002479_13640461.2002.11819467-Figure2-1.png", + "caption": "Fig. 2 Sampling for metallographic examination and for testing of frictional wear intensity", + "texts": [ + " The fusion process was run with such a water flow as to ensure disappearance of steam bubbles at the sample-water boundary, as observed through the transparent walls of the calorimeter. Super-fine dendrite Surface layer refinement of castings Orlowicz and Mr6z rate v=531\u00b0C/s; d) cooling rate v=125\u00b0C/s. Etch with Dix and Keith reagent precipitates of the a-phase and the a+{j eutectic phase were obtained in the re-melted area. An example of fusion area structure is shown in Figs. lb-d. Samples of material from the fusion area were taken for metallographic examination and frictional wear intensity testing. The sampling method is presented in Fig. 2. Fusion geometry measurements The measurements of fusion width and depth were per formed on metallographic micro-sections made in a plane perpendicular to the longitudinal axis of fusions. The applied method enables reading of measured values to O.Olmm. Micro-structural parameters Structural testing of the fusion area was aimed at deter mining the distances between the main axes of a-phase dendrites- the value of Am, (the distance between axes of second-row branches) - the value of }.20, as well as distances between silicon phase in the eutectic - the value of AE, e" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001171_ac9708194-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001171_ac9708194-Figure1-1.png", + "caption": "Figure 1. Schematic representation of components in the doublebioreactor flow cell. A, assembled reactor; B, upper cell body; C, lower cell body; D, top view of lower cell body; a, fixed bioreactor film (with immobilized GDH); b, rotating bioreactor (with film of immobilized GOx); c, Pt ring working electrode; d, O-ring; e, electrical connection; f, auxiliary electrode (stainless steel tubing); g, reference electrode (Ag/AgCl, 3.0 M NaCl); h, electrical connection. All measurements are given in millimeters.", + "texts": [ + "50 mL of 0.10 M phosphate buffer, pH 7.00) and GDH (10.00 mg of enzyme preparation in 0.25 mL of 0.10 M phosphate buffer, pH 7.00) were immobilized as described earlier.8 After being washed with purified water and 0.10 M phosphate buffer of pH 7.00, the enzyme preparations were stored between uses in the same buffer at 5 \u00b0C. The immobilized GOx and GDH preparations were totally stable for at least 1 month and 2 weeks of daily use, respectively. Flow-Through Reactor/Detector Unit (Double Bioreactor). Figure 1 illustrates the design of the double-enzyme reactor/ detector system. In this case, a modification of the system described in ref 6 was introduced. The body for the reference electrode was located at the top of the closed cell. With this modification, the problem that may be created by air bubbles is minimized. Glucose oxidase is immobilized on the top of the rotating reactor. The reactor at the upper part of the cell consists of a film of double-sided tape with CPG-immobilized GDH. The distance between the two reactors with immobilized enzymes is about 1 mm", + " A more complete reagent homogenization is achieved, because the cell works as a mixing chamber by facilitating the arrival of substrate at the active sites and the release of products from the same sites. The net result is high initial rate values (see Table 4). The main advantages of this system are its simplicity and the ease with which it can be utilized for the determination of glucose at very low levels. Moreover, this strategy is easily adapted to continuous-flow processing. The implementation of continuous-flow/stopped-flow programming and the location of two facing independent reactors (Figure 1), each containing one of the immobilized enzymes involved in the sequence illustrated earlier, permits (a) reduction of K\u2032M, (b) instantaneous operation under high initial rate conditions, (c) easy detection of accumulated products, and (d) relatively low enzyme loading conditions. According to the theoretical considerations offered by Kulys et al.,11 the amplification factor G is given, in this case, by the following equation: where K1 ) (IR)max/K\u2032M(GOx), K2 ) (IR)max/K\u2032M(GDH), and (IR)max is the maximun initial rate" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002671_iros.1997.649067-Figure3-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002671_iros.1997.649067-Figure3-1.png", + "caption": "Figure 3: Calculation of B for, (a) ak; a rotation axis (b) ak; a pivoting axis", + "texts": [ + " For searching more states, the planner simulates to rotate the object alternatively about all feasible pivoting axes available at SI, and mark an orientation state as a new one whenever an axis a k ( a k E { R j ) , { P j } ) is raised to feasibility (represented by a',). The question arises that why only the feasible pivoting axes at an state s k are considered for searching its offspring states and why not the feasible rotation axes. This will be explained later in Proposition 2. The above process is represented by the following equation of rotation matrix, [@(a,, e)]ak = ;ak # a p (1) a k E (Rj 1 ; then for feasibility ai must be parallel If Uk E (6) ; then for feasibility a& must lie in ?q- Axes ap, ak and u i are shown in Fig.3. to the *z-axis of {H)zy2 (assumption (4 ) ) . plane of {H}zv2 (assumption (5)). When ab can be raised to feasibility, then the angle 6 is derived from the above equation. Fig.3(a) & (b) is given to illustrate that how angle 6 is calculated for a k being a rotation and pivoting axis respectively. After calculating 6, if available, the object is rotated at 0 about axis up to reach some other orientation state s k . If SI, does not pre-exist in (Sz}, then it is added in IS,}. The following proposition is useful to check an available rotation axis if it could be raised to feasibility by rotating it about currently feasible pivoting axis. Proposition 1: A necessary and sufscient condition in order to make a vector a k E ( R j } parallel to the fz-axis by rotating it about a pivoting axis a,, is that it is orthogonal with ap" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003675_j.triboint.2003.11.004-Figure5-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003675_j.triboint.2003.11.004-Figure5-1.png", + "caption": "Fig. 5. Schematic illustration of experimental cam and roller follower.", + "texts": [ + " The total number of tests was: three types of industrial oils two viscosity grades twenty machine tools = 120. Fig. 4. illustrates experimental machine tool with cams in the machine assembly. During the test procedure on the experimental machine tool (Fig. 4), the cutting conditions were: Machine tool: Capstan lathe Cutting tool: Turning tool (cutter) Speed: 100\u2013300 (m/min) Feed: 0.1\u20130.4 (mm/revolution) Depth: 0.2\u20131.5 (mm) Coolant: Microemulsion (4%) Workpiece material: Medium carbon steel (YU) Contact conditions in a cam and roller follower contact for machine tools have been studied. Fig. 5 shows experimental cam and roller follower. The cam profile details and roller follower geometry are shown in Fig. 6. The contact load between cam and roller follower was set 930 N in maximum value. Working conditions of these cams (Fig. 4) and other influential factors are shown in Table 2. The results of the experimental investigation are pre- sented for two viscosity grades and three types of lubri- cating oils: . Hydraulic oils, category symbol HM, ISO VG 46 and 68 (first period of time) . Hydraulic oils, category symbol HG, ISO VG 46 and 68 (second period of time) " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003698_978-1-4020-2249-4_35-Figure4-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003698_978-1-4020-2249-4_35-Figure4-1.png", + "caption": "Figure 4. (a) An (IIM, Rl, RO)-singularity with 7rfs == 7ro; ,D is through 0 on 7ro . (b) An (RPM, IIM, 10, II)-singularity with two C4 legs. (c) An (RPM, Rl, IIM, IO)-singularity with e E 5, qA E ct, qC E ef, qL E err, L = B, D.", + "texts": [ + " Crl' although leg L is singular, the platform is constrained only by Wo and dim T = 4 (no 10). Either qL E Cr - Crl and we have an RPM singularity, or qL E cf and there is RI. In either case, since there is no 10, there must be an IIM-type singularity. (This follows from the interdependence rules of the singularity types, Zlatanov et al., 1995.) Moreover, if qL E cf and there is no RPM, an RO-type singularity must be present as well. Thus we have at least (RPM, IIM) or (RI, RO, lIM) and each can be augmented with (RO, II) if the other three legs cooperate. Figure 4(a) shows an example of a singularity of class (IIM, RI, RO). Thus, ll\"r5 ==1l\"0 allows the six singularity classes with lIM and without 10, impossible for \"usual\" PMs (Zlatanov et al. , 1994). Below, we generally assume that no ll\"r5 == 1l\" o. of the passive joints screws, dim PL < 4, i.e., dim VL > 2. According to Section 2, this is equivalent to qL E cf for some L. Each C4 leg has one passive dof, so the redundant passive freedoms of the mechanism are as many as the C4 legs. Figure 4b represents a configuration with two RPM freedoms. Provided that a1l7l\"r5\u00a27I\"0, a necessary condition for OE7I\"r5 is e E 13. Therefore, under this hypothesis, all RPM-type singular configurations have an end-effector in 13 and the workspace interior int (X) is free of them. (As we saw, if the geometry of some leg is such that 7I\"k5 == 71\"0, it is possible to have an RPM-type singularity and e E int (X) .) mechanisms is given in Zlatanov et al., 1994. The approach can be applied to the instantaneously equivalent PM" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003623_acc.2003.1238934-Figure3-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003623_acc.2003.1238934-Figure3-1.png", + "caption": "Fig. 3 Model of a two-link robotic manipulator.", + "texts": [ + " - i ) 1\u20191 t y p \u2019 -dj ) -s iLh, To ensure (28) being less than zero, the hitting control should be selected as (28) \u2018hi =sgn(s,)[Ifj l m x +C;=II~, l m x \u2018 I c e # l t l Y 2 ) I t 1 ~ ~ ~ l l n p ~ r ~ ~ \u201d 1 +Dj] This means that the inequality V! = siSi < 0 is obtained and the hitting control actually achieves a stable WSMC system. I 6. Simulation Results We demonstrate the proposed AWNNC by the tracking control of a two-link robotic manipulator with 2 degrees of freedom in the rotational angles described by angles q, and q2 , as shown in Fig. 3. The dynamic equations describing the motion of the robotic system are of the following form [14] m,r: +m,r,r,c, ( m , + m , ) r ~ + r n , r : + Z m , r , r , c , + ~ , m,r:+m - m , r , r d , (4, t i2 11 +!ccm, + c nli rI s d where m,, m2, J,, J2, r, = 0.51,, and r, = 0.51, are the mass, the moment of inertia, the half-length of link 1 and 2, g = 9.8mls\u2019 , and shorthand notations c2 = cos(q,) , s, =sin(q,) , cl2 =cos(ql + q 2 ) , etc. In the control experiments described below, the kinematics and inertial parameters of the arm are chosen as I , =2" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003864_50002-2-Figure2.10-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003864_50002-2-Figure2.10-1.png", + "caption": "Fig. 2.10. A motorcycle tyre (left) and a car tyre (right) in cambered position touching the assumedly frictionless road surface. The former without and the latter with torsion and bending of the carcass and belt. Top: rear view; bottom: top view.", + "texts": [ + " The terms associated with Oy/Ot have been neglected in the above equation and related neglections will be performed in subsequent formulae for Yo The radius of the outer surface in the undeformed state ro may depend on the lateral coordinate y. If the rolling body shows a touching surface that is already parallel to the road before it is deformed, we would have in the neighbourhood of the centre of contact: ro(y)= ro(O)- y sin~,. In case of a cambered car tyre a distortion of belt and carcass is neexted to establish contact over a finite area with the ground. The shape of the tyre cross section in the undeformed state governs the dependency of the free radius with the distance to the wheel centre plane. In Fig.2.10 an example is given of two different cases. The upper part corresponds to a rear view and the lower one to a plan view of a motorcycle tyre and of a car tyre pressed against an assumedly frictionless surface (p =0). It may be noted that in case of a large camber angle like with the motorcycle tyre, one might decide to redefine the position of the x and z axes. The contact line which is the part of the peripheral line that touches the road surface has been indicated in the figure. When the tyre is loaded against an assumedly frictionless rigid surface, deformations of the tyre will occur", + " At camber, due to the structure of a car tyre with a belt that is relatively stiff in lateral bending, the compression/extension factor Ox will not be able to compensate for the fact that a car tyre shows only a relatively small variation of the free radius ro(y) across the width of the tread, while for the same reason the lateral distorsion factor Oy appears to be capable of considerably counteracting the effect of the term with sin), in the second equation (2.48). For a motorcycle tyre with a cross section approximately forming a sector of a circle, touching at camber is accomplished practically without torsion about a longitudinal axis and the associated lateral bending of the tyre near the contact zone. Figure 2.10 illustrates the expected deformations and the resulting much smaller curvature of the contact line on a frictionless surface for the car tyre relative to the curvature exhibited by the motorcycle tyre. Since on a surface with friction the rolling tyre will be deformed to acquire a straight contact line, this observation may explain the relatively low camber stiffness of the car and truck tyre. As mentioned above, circumferential compression of belt and tread resulting from the normal loading process (somewhat counteracted by the presence of hysteresis also represented by the factor Ox ) gives rise to a decrease of the effective rolling radius" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002432_bf01574849-Figure8-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002432_bf01574849-Figure8-1.png", + "caption": "Fig. 8. The test arrangement", + "texts": [ + " In order to formulate the interaction analytically we introduce the symbol M in Eqs. (12) and (13). i The Eqs. (12) and (13) are then differentiated to have dTs = M . di e + i ~- a_MM di (15) ~M d ~ l q ~ M . d i q -~ i q 9 - - di (16) I~ewriting Eq. (15) and Eq. (16) gives dTe = [ M @ (is)2 ~M] ie . iq ~M diq (17) 7 dis +---7. d ~ q iq . i s 8M . d i e + M -I- (iq)2 7 .diq 5 An experiment to verify the equations The experiment was made on the machine previously described in Sect. 3. The test arrangement is shown in Fig. 8. The stator d-winding was supplied from an electronically regulated d- -c current source with a very high a- -c impedance while the q-winding was supplied with a d- -c current with an a- -c component superimposed. The d--c components I e and Iq were varied to have different values of the angle cr while at the same time the resultant airgap d--c current Io was kept constant. The amplitude ~ac of the a- -c component in the q-winding was kept constant and at a small value compared to I 0. For different values of the angle c~, the voltage induced in the d-winding was recorded and analysed. Figure 8 also gives a qualitative explanation of the magnetic interaction between the d- and q-windings. The fundamental in the flux density distribution is represented by the arrows B~ and it is obvious that due to the non linear relation between BI and the resultant current or mm/, the change of flux linked with the d-winding caused by the moving of B~ is not compensated by a proportional change of the amplitude of B1. The current in the d-winding is constant and with di ~ = 0, the Eqs. (17) and (18) will take a very simple form" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001492_iros.1997.649095-Figure3-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001492_iros.1997.649095-Figure3-1.png", + "caption": "Fig. 3 mechanism of a leg", + "texts": [ + " In section 5 , the transition walking fiom the walking on a straight line to the walking on a curved line only use the step width control is mentioned. In section 4 and 5, the experiments was done to certify the methods. And in section 6, we conclude this paper. 2. Centipede Type Walking Robot Fig. 1 shows the centipede type walking robot. In this figure, the entipede walking robot is consisted with six units. The length of this robot is about 120 cm and the Fig. 1 photo of centipede walking robot ROC. IROS 97 0-7803-4119-8/97/$1001997 IEEE 403 Each leg hias the mechanism which has two degree of fieedom. Fig. 3 shows the mechanism of a leg. The motion of the foot end is made by the rotational motion of the base joint and the telescopic motion of the leg. Using this mechanism, the foot end is moved on a vertical plane. On this robot, each leg has no degree of fieedom to move to the lateral direction. The units are linked by a rlod. Both units at ends of a connecting rod rotate oppsite direction each other at the rod ends. And to equalizing the rotate angle of neighbor units against the rod, gears which are fixed on the both units is equipped (show Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003114_robot.1998.680618-Figure2-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003114_robot.1998.680618-Figure2-1.png", + "caption": "Fig. 2. Zoom of the manipulator\u2019s endeffector with 6 DOF force/torque sensor (ATI, FT Mini 100/5), to be controlled with the simulated control laws in the near future", + "texts": [], + "surrounding_texts": [ + "1 Introduction\nMany robotic applications involve intentional interaction between the manipulator and the environment where contact stability to moving objects and surface tracking with defined contact force typify demanding tasks [6, 11, la] . Establishing stable contact to rigid objects as well as tracking surfaces with unknown shape and elasticity are the prerequisites of a manipulator handling its environment naturally. In several approaches of force/position control, the contact between the endeffector and the environment is assumed to be soft, realized by aflexible tool [a, 4,7]. However, an artless tool and the environmental surface are often stiff, e.g. a screw driver and a metal surface like a screw head. At the contact point, the endeffector and the environmental surface will maintain their shape regardless of the forces exerted. Current techni-\ncal approaches for hard contact are limited to special cases with a reduced number of degrees of freedom [5] To make the endeffector of a manipulator follow in a stable way the edge or the surface of a workpiece while applying prescribed forces and torques we developed a novel control concept: Neural Force Control (NFC)\u2019 .\nThis work was supported in part by Federal Ministry for Education, Science, Research, and Technology (RMBF) under grant \u201cDEMON\u201d\n0-7803-4300-~-5/98 $10.00 0 1998 IEEE 2048", + "NKN\n2 Results\n2.1 Control concept\nNFC is a neural hybrid force/position controller [3, 91 based on the \u2018hard contact\u2019 approach. Conlaining the two modules NKN (fig. 3) and NDN (fig. 4), NFC (fig. 4) computes the control signals for the manipulator from the joint angle errors and the force errors at the endeffector. NKN delivers a fast computation of the singularity robust, differential kinematics (Jacobian matrix and its inverse), necessary for the NFC calculations. The inverse dynamics of the manipulator are performed by the NDN module. To establish &able contact to unknown surfaces or objects, CVC (fig. 6) is applied to the control concept to reach a tender impact. Fig. 1 shows the 6 DOF robot equipped with a 6 D-wrist force/torque sensor (Assurance Technologies Incorporation, F T Mini 100/5).\nNKN: The NKN module (fig. 3) realizes the differential inverse kinematics (DIK) by calculating a modified Jacobian matrix J N ~ ~ (e) and its inverse J i i t ( e ) . N K N\nmeets two major requirements in robust robot control:\n1. Singularities, appearing in the entire workspace, must not lead to undefined high joint angle velocities or tremendously increasing force errors in case of contact with surfaces. A strategy to maintain singularity robustness by using a modified inverse of the Jacobian matrix (SR-Inverse) lhas already been described [lo]. This kind of inverse kinematics delivers continuous and feasible solutions at or in the neighborhood of singular positions. The SR-Inverse can be computed using singular value decomposition (SVD). To avoid the cumbersome SVD in the control loop, the SR-Inverse has been trained offline into a radial basis function type neu-", + "ral network (NN). 2. Another feature of NKN is the handling of robot\nspecific constraints as avoidance of self-collisions,\nstatic obstacles in workspace and joint angle margins. The implementation of the DIK as a NN includes the capability of modifying the inverse Jacobian such that critical positions cannot be reached.\nThe RBF net used in NKN consists of about 20,000 neurons. I ts inner structure optimizes the total number of neurons by increasing the neuron density in critical regions (e.g. node allocation). The NN can be trained offline with the SR-Inverse of the Jacobian matrix as teacher input or online by using the variation of joint angle Q and Cartesian vector 2 while tracking test movements. In the second approach the NN learns the DIK mappings only by using sensory input without any previous knowledge concerning the manipulator.\nNDN: The NDN performs a nonlinear decoupling and feedback linearization of the manipulator's dynamics using the plants inverse dynamics model, trained into a neural network. NDN takes into account inertias, coriolis- and zentripetal effects as well as gravitation and various kinds of joint friction (coulomb- , stribeckand viscose friction). NDN is embedded into a neural hybrid force/position controlsystem NFC, achieving an orthogonal separation of force- and position constraints.\nThe most significant difference to common approaches is the realization of the inverse dynamics by neural networks [1, 8, 141. The matrix elements of the inverse dynamics equatio? (eq. 2) like mass matrix M(8), coriolis coupling e(@, e),gravitational influence G(B) and different types of angle velocity dependent frictions R(8) (coulomb friction, stribeck friction and viscose friction) are represented by RBF nets respectively. The diagonal entries of the matrix S (ey. 1, 3, 4) associated with the components of the force to be controlled are chosen equal to one, and the remaining elements are zero. In the adaptation phase, the matrix S is completely set, to zero, to turn off the force controller which allows to learn the invers model of the plant from the error AT" + ] + }, + { + "image_filename": "designv11_65_0002771_s100510170199-Figure2-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002771_s100510170199-Figure2-1.png", + "caption": "Fig. 2. Sketch of the amplification mechanism of disturbance proposed for sidebranching. Perturbations are exponentially amplified and linearly stretched as they propagate down the side of growth cell. The WKB approximation asks that disturbance wavelengths are small compared to the tip curvature radius, a condition which is merely satisfied in practice.", + "texts": [ + " A conclusion about the relative role of tip undercooling and tip form in the amplification of cell tip disturbances is then drawn. Especially, a criterion on cell tip curvature radius is derived for the noise amplification mechanism to be compatible with thermal gradient-induced sidebranching in directional solidification. We address the effect of tip undercooling on the growth factor of tip disturbances in the framework of the noise amplification theory. Within the noise-induced mechanism, sidebranches result from the amplification of tip disturbances during their propagation down the side of growth interfaces (Fig. 2) [5\u20138]. Amplification is given by the same mechanism as that invoked for the primary instability of growth interface. However, its consequences are modified by a wavelength stretching generated by advection by a non-uniform flow: the tangential flow appearing in the reference frame of a growing curved shape. This stretch makes a given wavelength drift within the instability band and thus experience a varying growth rate. Following this, the amplification factor at a given distance from the tip depends on the initial wavelength of disturbance", + " In addition, compatibility \u2013 not only qualitative but also quantitative \u2013 of the growth factor expression (47) with the critical surface for sidebranching [13], will provide an even more severe criterion for the relevance of a local linear convective instability for sidebranching. From these confrontations, interesting improvements of the understanding of sidebranching may be reasonably expected. Denoting s the curvilinear abscissa on a growth interface with origin taken at the tip, q the wavenumber of disturbances, \u03c3(s, q) their local growth rate, and v\u03c4 (s) their local tangential velocity along the interface, the growth factor reads, within a WKB approximation [5\u20138] (Fig. 2): \u0393 (l) = \u222b s=l s=0 \u03c3(s, q) v\u03c4 (s) ds. (A.1) Tangential velocity v\u03c4 is expressed by taking into account both the advection velocity (V \u00b7 \u03c4) \u03c4 of disturbances induced by growth velocity V=Vz and the migration velocity \u2212Vcz of phase heterogeneities in thermal gradient Gz: v\u03c4 = (V \u2212 Vc)z \u00b7 \u03c4. (A.2) Change of disturbance wavenumber q by stretch during advection towards the grooves obeys: \u2202(qv\u03c4 ) \u2202s = 0. (A.3) Meanwhile, normal directions n to the front turn following front curvature. Both evolutions modify the local growth rate \u03c3(s, q) in agreement with the dispersion relation of the primary instability of growth fronts [12]: \u03c3(s, q) = q(V \u2212 Vc)z \u00b7 n\u2212Ddq \u2212 V |z \u00b7 n| D \u00b7 (A" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001124_02783649922066574-Figure2-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001124_02783649922066574-Figure2-1.png", + "caption": "Fig. 2. Projection of a point according to the pinhole camera model.", + "texts": [ + " It is important to notice that in the determination of the view parameters included in the orthographic camera model, a precision-enhancing measure known as flattening (Skaar and Gonz\u00e1lez-Galv\u00e1n 1994; Gonz\u00e1lez-Galv\u00e1n and Skaar 1996; Gonz\u00e1lez-Galv\u00e1n et al. 1997) has been used. This measure consists of modifying the raw camera-space samples acquired while the maneuver is executed, so that they become more consistent with the orthographic projection-estimation model of eqs. (2), (3), and (4). The flattening procedure is based on a presumption of a pinhole projection of physical points into the two-dimensional image plane. Consider the diagram depicted in Figure 2. If (xci , yci ) represents the ith raw camera-space sample, the flattened sample is determined by ( xci Zi Zr , yci Zi Zr ), where Zi represents the location of the sample along the optical axis of the camera, and Zr is the location of a reference point, usually chosen to be near the place where the maneuver ends. Gonz\u00e1lez-Galv\u00e1n and colleagues (1997) explained the way in which Zi and Zr are evaluated in our experiments. However, to give an intuitive idea of the flattening procedure, consider the case in which a photograph of two persons of the same height is taken, as depicted in Figure 3" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003626_1350650041323368-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003626_1350650041323368-Figure1-1.png", + "caption": "Fig. 1 Typical automotive gasoline engine piston and piston ring pack", + "texts": [ + "eywords: piston ring, piston groove, wear, impact, sliding, microwelding NOTATION F total axial friction force per unit circumference owing to hydrodynamic action and asperity contact FP radial friction force per unit circumference at the pivot g acceleration due to gravity m piston ring mass per unit circumference p1 gas pressure above the ring p2 gas pressure below the ring PA axial applied gas pressure force per unit circumference PL radial gas pressure relief force per unit circumference at the unwetted lower edge of the ring PR radial gas pressure force due to gas pressure acting radially on the ring PU radial applied gas pressure relief force per unit circumference at the unwetted upper edge of the ring RH axial component of hydrodynamic force per unit circumference RP axial reaction force per unit circumference at the pivot t time T radial force per unit circumference owing to inherent ring elastic tension U velocity of cylinder liner relative to the piston WC radial force per unit circumference owing to asperity contact WH radial component of hydrodynamic force per unit circumference x axial coordinate relative to the centre of mass z radial coordinate relative to the centre of mass Reciprocating internal combustion engines use one or more pistons to convert the energy obtained from combustion into useful kinetic energy. A typical design for a modern automotive, four-stroke, gasoline engine is shown in Fig. 1. The piston has a series of three piston rings, the piston ring pack, for sealing the combustion chamber to maximize combustion gas pressure and The MS was received on 31 July 2003 and was accepted after revision for publication on 14 October 2003. * Corresponding author: School of Mechanical Engineering, University of Leeds, Woodhouse Lane, Leeds LS2 9JT, UK. J04503 # IMechE 2004 Proc. Instn Mech. Engrs Vol. 218 Part J: J. Engineering Tribology at Univ of Newcastle upon Tyne on October 16, 2014pij" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000949_1.2802323-Figure3-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000949_1.2802323-Figure3-1.png", + "caption": "Fig. 3 Experimental system", + "texts": [ + " Fortu nately, this is solvable by modern theory of estimation as well as by methods for numerical differentiation and data smoothing. As an example, even the third derivative of mechanical position (that is, \u00ab -I- 1 derivative) is successfully utilized in (Ostojic et al , 1993), by combining numerical differentiation and lowpass filtering. Although experimentally evaluated, the use of estimated instead of measured error-derivatives in recursive control is not yet investigated analytically. This is a topic of current research. Proposed approach was utilized in designing motion control of a rigid pendulum shown in Fig. 3. The pendulum was driven by a direct-drive electrical motor. Basic parameters of the sys tem are listed in Table 1. Motion controller consisted of an 1386- Table 1 Basic parameters of the plant Parameter m h I z mi kr 'max\u00bb *miii Encoder res. Value 7.0 0.15 0.3 0.06 1.0 4.5 \u00b110 655 360 Unit kg , kg m m m kg Nm/A A ppr based personal computer with DSP board (NEC /xPD77230, 32- bit, floating point) and a common I/O board. Sampling period in the system was 1 ms. The plant is modeled by V mz + mj . , \u201e, , 9 = -\u2014- ^ ^ s i n ( 6 ' ) + md (26) where 9 is the output, m is the total mass of the rod and rotor, /o is the moment of inertia, m,^ is the load, / is the length of the rod, z shows location of the mass center, g is the gravity con stant, and k-r is the motor constant" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002733_pesc.1999.785604-Figure6-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002733_pesc.1999.785604-Figure6-1.png", + "caption": "Fig. 6. Encoder.", + "texts": [ + " The reference signal corresponds to the reference speed signal which is used for switch-on signal, and the rotor encoder signal to the rotor position signal which is used for a switch-off signal. From reference frequency signal and encoder signal, phase detector generates advance angle to meet instantaneous torque variation. The output of phase detector is used as the signals for advance angle of dwell angle as well as the input of loop filter. A flat torque response could be made by the proper gain of loop filter. The output of loop filter is used in the buck-converter to regulate voltage. B. Rotor encoder Fig. 6 shows the 2-bit slot-disc type encoder. The switch-off signals are made using this encoder. The encoder is mounted C. Switching angle control The switching angle has been regulated to remain within the maximum value. It is selected not to develop negative torque. If it exceeds the maximum value by an abrupt increase of load or overload, then the frequency of the reference speed signal will be lowered. On the contrary, when the difference between the reference and feedback signal is very small, which may occur during an abrupt decrease of load, the reference signal will be advanced and be moved to the dynamic dwell angle mode" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003324_icsmc.2003.1244217-Figure6-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003324_icsmc.2003.1244217-Figure6-1.png", + "caption": "Figure 6: Two bicycle states evaluated as equal", + "texts": [ + " 5 is used by the planning algorithm to verify if the current state has already been explored (it is in closed-states) or is already in the list of states pending exploration (it is in open-states). If the state has not been explored but it is part of the list of states pending exploration, then the algorithm checks whether its cost needs to be updated based on the current path. If the current state is part of neither the list of closed-states nor the list of open-slates, then the state is added to the list of open-states and set for pending exploration. In Fig. 6 and Fig. 7 we show a graphical representation of how the function in(state, list-of-states) evaluates whether a state is present in a list of states or not. In Fig. 6 we see two states considered as equal. The consideration is based on the fact that the Euclidian distance between the two center of gravities is less than a given position threshold and the difference of the two heading angles is less than a given angular threshold (heading threshold). Fig. 7 instead shows the situation when two states are not considered equal because either one (or both) of the two differences is greater then the corresponding threshold. In the next section, we proceed by showing some simulation results for the planning algorithm" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002318_cdc.1985.268463-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002318_cdc.1985.268463-Figure1-1.png", + "caption": "Figure 1", + "texts": [ + " = R i= 1 Yi Y where Fxi, Fyi are the horizontal components of the force on foot i and yi are the coordinat of its position proJected into the horizontal plane. xi : Location of the Force Center Equations (1 to 3 ) can be written in the form n F = R - -ti -xy i= 1 n x E . = T i=l Exvi -tl -2 where UR is the horizontal component of the resultantxYforce, T is the vertical component of the resultant torsfie, E is the projection in the horizontal plane of H$ position of foot point i, and E . is the horizontal component of the force on f?ot i, as shown on Figure 1. The central force field constraint requires that 13 where Q is the position of the force center and L is a constant for the force field. k is a unit vector in the positive z (vertical) direction. Substitution from equation (6) in equation L 4 x 1 Exyi - nL4xE = R i= 1 -Y It is convenient to designate che position of the centroid of the foot point pattern by p. Then: ( 4 ) gives n . n E = ; z Exyi i=l Thus - k x g = k x ? - - R 1 nL -xy ( 9 ) Also, since 2 lies in the xy plane k x 0, let us and u~, be the velocities in the regions -x~ - ~ ~< x ~< -x~ + ee and x~ - e3 ~< x ~ x2 respectively, 240 August 91 Vol 24 No 4 then according to the velocity profile, shown in Fig 2, one may observe the following: 0u~ >0, a0, 0~ = M(t) (6) \u00b72 mr - ~ r tp = F(t) Here ml , ~ are masses of the spindle and of the arm respectively, 11 is a mo ment of inertia of the spindle with res pect to the axis of rotation, 12 is a moment of inertia of the arm with respect to its centre of mass, Z is a vertical displacement of the arm, r is a dis tance of the centre of inertia of the arm from the axis of rotation, ~ is an angle of rotation of the arm, g is the gravity acceleration" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003864_50002-2-Figure2.8-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003864_50002-2-Figure2.8-1.png", + "caption": "Fig. 2.8. Side view of front part of vehicle (motorcycle) with wheel turned over angle ~ about steer axis (a) (Exercise 2.2).", + "texts": [ + " The front frame can be turned with a rate 3 (=dO/d t) with respect to the rear frame. At the instant considered the front frame is steered over an angle c~. It is assumed that the effective rolling radius is equal to the loaded radius (re = r, C* = C). The longitudinal slip at the front wheels is assumed to be equal to zero (Vsx =0). Derive expressions for the lateral slip speed Vcy, the linear speed of rolling Vr and the lateral slip tana for the right front wheel. Exercise 2.2. Slip and rolling speed of a wheel steered about an inclined axis (motorcycle) The wheel shown in Fig.2.8 runs over a flat level road surface. Its centre A moves along a horizontal straight line at a height H with a speed u. The rake angle e is 45 ~ The steer axis BA (vector a) translates with the same speed u. There is no wheel slip in longitudinal direction (Vsx = 0). Again we assume re = r. For the sake of simplifying the complex problem, it is assumed that the wheel centre height H is a given constant. Derive expressions for the lateral slip speed Vcy, the linear speed of rolling Vr and the turn slip speed ~b in terms of H, u and ~ for 6 = 0 ~ 30 ~ and 90 ~ Also show the expressions for the slip angle a, the camber angle y and the spin slip (p with contributions both from turning and camber" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001226_20.717826-Figure5-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001226_20.717826-Figure5-1.png", + "caption": "Fig. 5. Distribution of flux density, ZA = 0.33 A, I B = 0 A.", + "texts": [ + " 4a shows the flux density distributions in the case that rotor position 81,L=150, that is, the magnets of the rotor are at a position corresponding to the teeth of phase A. Large flux density appears in the teeth of phase A. Fig. 4b shows the flux density distributions in the case that 8,,,=22.5\", that is, the magnets of the rotor are at a position corresponding to the teeth of phase B. Then, large flux density appears in the teeth of phase B. In the case that 8,=Oo and 7.5\", the flux densities have the same amplitude as and the opposite direction to those in the case that 8,,=15\" and 22.5\", respectively. Fig. 5 shows magnetic flux distributions with exciting current I ~ = 0 . 3 3 A. Excitation of phase A produces flux density in the lower section of the motor. In the case that 8,=15\", the flux density produced by the excitation of phase A is added to the flux density produced by the rotor magnet. This situation corresponds to the stable po- sition of the motor. In the case that 0,=7.5\" and 22.5\", the flux density exists in four teeth. These situations correspond to the position developing the maximum torque" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003375_icmech.2004.1364455-Figure4-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003375_icmech.2004.1364455-Figure4-1.png", + "caption": "Fig. 4. Tbr eseigg optimal motion closely follows the obstacle contours.", + "texts": [ + " Also the necessary run.time increases nonlinearly with the accuracy. Typically with a sliort duration (T = 20) it iz energy optimal to fold the manipulator before turning and afterwards unfolding it. The obstacles mainly interfere within the unfolding motion. Taking into account joint limitations the preceding motions are likely to be impossible. i n figure 3 the revolution of the last four joints is restricted to 450 degree. Thus (e') contains [ / q - q ~ . , J + = 0, a = 2,. . . , 5 . Another scenario is shown in figure 4. The energy optimal control with a safety margin A- = 0 very closely follows the obstacle boundaries. Due to model uncertainties and the finite number of sensors it is nossible that is the result of a numerical solution with a required accuracy of in the objective value. The solution is found within 10 a control scheme using the determined U,, will cause collisions. For safety the collision detection should bc more conservative as shown in fignre 5 . Therein the same task is accomplished with A, = 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000007_piae_proc_1922_017_075_02-Figure4-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000007_piae_proc_1922_017_075_02-Figure4-1.png", + "caption": "FIG. 4. I", + "texts": [ + " y = thc pressure-angle. In Fig 3, which is a summary of tests for combined radial and thrust load for spherical ball bearings, it will be seen that formu12 (7) givecu sufficient aclouracy. The use of the formule here, however, should involve greater care owing to the difficulty in determining the actual thrust load in many applications. SINGLE-ROW RADIAL BEARINGS. The corresponding data for single-row journal ball bearings would appear to be covered in the following manner:___ Nl a = vt (Ref. Fig. 4) for stationary outer race ... (8) I + - 2 2 . . . . . . . . . . . . . . a = r, for stationary inner race (9) 1 + - TY where Nl = the number of balls. For single-row bearings without filling-slot (10) ........................ 264 K; - 8.8 264 4a.n + 5 Ki = ___ + 8.8 lb. per (& in.)2 . . . . . . . . . . . . ( 1 1 ) Ki = K (1 + 0.0001 U) (1 + 0.007 D') (12) (13) . . . . . . . . . For single-row bearings with filling-slot 19s ..................................... a.n = (--- Ki - 6.6 ) 3 - 5 198 4a" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002302_iros.1997.655123-Figure5-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002302_iros.1997.655123-Figure5-1.png", + "caption": "Figure 5: Definition of motion directions over a type-B C-face.", + "texts": [ + " the commanded velocity t$ must allow the motion of the mobile object while maintaining contact with the fixed objects: where e v\"b: is the component that al%ows the motion towards the goal, moving the mobile object reference point tangentially over the C-face; it is on the desired motion direction over the tangent plane. e i?f is the component that allows the contact maintenance, producing a reaction force; it is in the opposite direction of the edge of the generalized friction cone which is determined by the motion direction. The magnitudes ut and vf depend on the desired velocity and on the desire< reaction force, respectively. Let us define (figure 5 ) : e p1 the start configuration ( 2 1 , y1, ~1). e p2 the desired stop configuration ( x 2 , y 2 , q 2 > . e paux the configuration obtained by rotating the object an angle a, = (q2 - q l ) / p around the contact point at configuration pl : x,,,, = 2 1 + r , COS a,. - ry sin a,. (12) yanX = y~ + r, sin a,. + ry cos a,. Qaux = Qz 0 dp the vector from paux to pa: Proposition 5: For one basic contact : where as = (.it, &) and a , = (42 - q I ) / p . Proof: The direction of rotation about the contact point is t\", and the amount to be rotated is a," + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003963_bf02844861-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003963_bf02844861-Figure1-1.png", + "caption": "Figure 1b) Inertial coordinate system.", + "texts": [ + " This can be solved by integrating the full nonlinear six-degrees-of-freedom equations of motion numerically on the basis of an aerodynamic database constructed from wind tunnel test data. The flight trajectory has already been simulated (Seo et al., 2006). Control parameters are initial conditions. All the initial conditions can be controlled by the kicker as the ball is kicked. There are six initial conditions: these are the magnitude of V \u2192 0, the flight path angle, \u03b30, the azimuth angle, \u03c70, the magnitude of \u03c9\u21920, the pitch angle, \u03980, and the yaw angle, \u03a80, as shown in Fig. 1a and Table 1. In Fig. 1a, (XE, YE, ZE) is the inertial coordinate system, while (xb, yb, zb) is the body-fixed coordinate system. The origin of the body-fixed coordinate system is defined as the centre of mass of the ball. Its xb axis is aligned with the longitudinal axis, and yb and zb are aligned with the transverse axes, with the zb axis passing through the valve. The ranges of the parameters in Table 1 are defined such that they cover the practical ranges. Since the longitudinal axis, xb, is aligned with the spinning axis, the direction of the angular velocity vector \u03c9\u21920 is defined by Euler angles \u03980 and \u03a80. The initial roll angle, \u03a60, is assumed to be 0\u00b0. The inertial right-handed coordinate system is shown in Fig. 1b. The origin is defined as the point of intersection of the goal line and the left-hand 88 Sports Engineering (2006) 9, 87\u201396 \u00a9 2006 isea touchline from the kicker\u2019s view on the ground, where the XE axis is in the horizontal forward direction, the YE axis is in the horizontal right direction and the ZE axis is in the vertical downward direction. In this study, an elitist non-dominated sorting genetic algorithm was applied for the optimisation (Deb, 2002). An outline of the procedure is given in the following", + " However, very few players utilise this skill. It is difficult to gain longer distances by kicking with the outside of the foot. Therefore, for convenience, P > 0 is assumed to be from a left-footed kicker, and vice versa. With a left-footed kicker, the ball tends to hook to the right (Seo et al., 2006). The kick into the right-hand touch from the kicker\u2019s point of view is optimised. The initial position is assumed to be (XE0, YE0, ZE0) = (20, 50, \u20130.5). An image of the flight trajectory is shown as I in Fig. 1b. Two objective functions F1 and F2 are shown in Table 2. The point at which the ball makes contact with the ground is denoted by tf. F1 is the flight distance in the forward direction at tf multiplied by \u20131, and F2 is the absolute value of the difference YE between the ball and the right-hand touchline at tf. Both objective functions must be minimised for this optimisation. The constraint g1 is defined as g1 = YE(tf) \u2013 70 > 0 (1) This means that the ball should go out into the righthand touch at tf . The kick into the left-hand touch is also optimised. The initial position is assumed to be (XE0, YE0, ZE0) = (20, 20, \u20130.5). An image of the flight trajectory is shown by the broken line II in Fig. 1b. Two objective functions F1 and F2 are shown in Table 3. F2 is the absolute value of the difference YE between the ball and the left-hand touchline at tf. The constraint g1 is defined as g1 = \u2013YE(tf) > 0 (2) This means that the ball should go out into the lefthand touch at tf. The results are shown in Fig. 3. The data for I are shown by the open circles, and those for II are shown by the open triangles. Two objective functions are shown in Fig. 3a. XE increases with increasing \u2206YE. Although the largest value of XE and the smallest value of \u2206YE are the ideal situation, it is impossible for both objective functions to be satisfied simultaneously" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000412_s0263574700017768-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000412_s0263574700017768-Figure1-1.png", + "caption": "Fig. 1. Schematic of the seven-joint manipulator.", + "texts": [ + " And the favorable formulation made possible the derivation of analytical solutions for joint rates. In addition, it may be noted that the simplified description for this velocity Jacobian is really useful to detect the presence of singularities of the mechanism. This paper starts with the geometric identity method to find the feasible arm solutions for one case of the seven-joint manipulator and is followed by an algorithmic approach to the velocity Jacobian as the main part. 2. FUNDAMENTAL SCHEMES OF MODELS As shown in Figure 1, the manipulator to be studied is composed of a seven jointed structure with a shoulder S, an elbow E, a wrist W and a hand H, where the individual Cartesian positions from the origin S are expressed in terms of (xe, yc, ze), (xw,yr,zw) and (xh, )>h, zh). And the desired position and orientation of the hand are described by the homogeneous matrix n, o, ay Yd *z %d 0 0 0 1 (1) The two operators x and \u2022 denote a vector and a scalar product, respectively, and J, = sin 0, and ci = cos 0, are used as the conventional notation throughout the derivation", + " it-iSZg (36) (37) Hence, we may write an expression for 'v, as 'V iff _ '7 'i- ' Y 'lr \u2014 ' V 'Ir (for a revolute joint /) (38-a) (with 0y- = dy), (for a prismatic joint;') (38-b) Now, we can derive a compact relation between ('Vr, i 0, and tx = tx0(t) when K < 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002629_iros.2001.973411-Figure5-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002629_iros.2001.973411-Figure5-1.png", + "caption": "Figure 5: Ball shooting behavior", + "texts": [ + " The PTP motion is then realized by combining joint rotations for each joint: PTP,(A0) = { J R d ~ ~ ( f i ~ ) } where fii is given by VmiY the knowledge array for joint rotation of joint i. Combining several simple PTP motions gives a complex and dexterous motion: JPTP7(a) = {pTp~~(Af? i ) } , which is, in the ICMC, referred to as an intelligent PTP motion. The control parameter is a = {qi ) . On the other hand there is no formulary form of motion parameters 7 , because intelligent PTP motions are used as element motions for a wide variety of behaviors. Next a ball shooting behavior by Lauron I1 [8] is A considered (Figure 5) . Lauron I1 is a six-legged walking robot that can walk on rough terrain with force sensor feedback as well as walk forward/backward and sideways based on the modular controller architecture, MCA [9]. When a goal position p~ and a ball position & are given, the robot approaches to the ball and kicks it, that is, Shoot is a sequential composition of Appna ch and Kick in which Apppeca to PA = (ZA,YA) is executed first and without waiting its completion Kick begins at time tK: where O+ represents that Appoaca starts simultaneously with Shoot", + " The resultant o p timal parameters for the case po = (3000,0), OK = [180,300] x [-400, -2801, J1 = Ja = 4 , are stored as a knowledge array VK = {wgai1} E R(313, Ja, J I ) as shown in Figure 9. With this empirical knowledge, VK, the robot can adaptively realbe sub-optimal Kick for arbitrary bK E OK even for inexperienced ball positions. Secondly Applloeca from various initial robot positions & to a speciiic approach target pA = (xA,YA) near the ball position & is optimized. pA or & is relative to &, thereby the above is restated as \"Approarh fiom a specific initial robot position = 0 to various approach target pAn (See Figure 5). Here, for simplicity, Apputwh is formulated as just one Step in a standard posture with control parameter aA = ( v , t ) , and then aA is optimized with criterion llpA - &(uA)(( where is the actually reached robot position. The resultant knowledge array VA = ( u t a j , } E B(312, J2, J 1 ) for the case OA = [1300,1700] x [-400,0], Jr = Ja = 3 , is shown in Figure 10. Giving an arbitrary target pA to VA, the knowledge array outputs the sub-optimal control parameters for the requested Appoach. Finally Shoot for ball positions & = (zB,~B) E QB is optimized", + " The control parameter us = (pA,tK) is optimized for the criterion Js(aS) = llpo - Pg(uS)JJ +Ts. As a result sub-optimal Shoot is successfully obtained by optimal composition of Appmerll and Kick. Two examples of sub-optimal Shoot for the case po = (4350, -650) are given in Table 1. (Note that the onginal necessary time for Appnaca is 2.O[sec], and tK in the results shows that the necessary time is shortened 4.5% and 3.5%, respectively.) When & is given, then approach target pa and estimated ball position for Kick, psK = L(&,pA), with tK are transfered to VA and VK, respectively (See Figure 5 again). In other words, only by giving a ball position sub-optimal Appoeea and sub-optimal Kick are successively executed. 6 Concluding remarks In this article, it is shown that a dexterous shooting behavior is achieved through a knowledge array network. It is noted that the desirable shooting is realized not by solving the equations of motions of all the robot legs and the ball, but by learning the component element motions at fist and adequately combining them for the second. The point is that there is no need to consider the details about lower level component element motions to learn a higher level composite behavior" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002271_acc.2002.1023214-Figure3-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002271_acc.2002.1023214-Figure3-1.png", + "caption": "Figure 3: Differentiation of e.", + "texts": [ + " 4 Application t o an Optical Storage System In this section, we introduce a typical example of the servoing technique widely used in the optical storage systems such as the CD-ROM/ReWritable Disc drives as well as the DVD-ROM/EtAM drive systems. For an eccentric disc with the excessive eccentricity, we examine the conventional feedback control and add the proposed feedforward compensation loop. Also, some issues for implementation are discussed. Time-derivative of the tmcking error: To obtain the timederivative of the tracking error, we designed the band-pass filter as shown in Fig.3. The first cutoff frequency was chosen considering the disc operating speed (e.g., x4 constant linear velocity (CLV) control in this experiment). The xl-CLV speed in CD-ROM drives means that the relative velocity of the pickup lens and the disc surface is 1.3 m/sec by the CD-ROM specification. Since the radius of data t racb ranges from 2.5 cm to 6 cm in 12cm discs, the rotational frequency of the discs in the lx-CLV mode should be around 3 Hz (outermost) to 8 Hz (innermost). Therefore, in the 4x-CLV mode, the maximum rotational frequency becomes 32 Hz" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002623_a:1015265514820-Figure7-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002623_a:1015265514820-Figure7-1.png", + "caption": "Figure 7. The trial arena and the robot trajectory for the noise-free case.", + "texts": [ + " The second one was impulse noise and was defined as a certain probability (INp) for the value of a sensor to become zero when it was actually positive (false negative), or become positive when it was actually zero (false positive). The simulator modeled orientation noise as an equal probability for the final sensor value to lie in the interval of plus or minus a maximum deviation (Oe) from the original sensor value. This deviation was expressed in degrees. Table 1 describes the simulation results in an environment with above average complexity (see Fig. 7) and with different noise parameters. Ten trials were Table 1. Simulation results. Parameter settings WN = \u00b110% WN = \u00b120% WN = \u00b110% WN = \u00b110% Inp = 5% Inp = 5% Inp = 7.5% Inp = 5% Oe = \u00b110\u25e6 Oe = \u00b110\u25e6 Oe = \u00b110\u25e6 Oe = \u00b115\u25e6 Success 80% 70% 60% 60% Near-Miss 10% 0% 0% 0% Failure 10% 30% 40% 40% performed for each parameter setting. The same starting and goal positions were used in all trials. Assuming no noise in the sensors, the robot was able to conclude the task in 130 simulation cycles. With noise in the sensors we let the simulation run for a little more (170 cycles)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003666_1350650042128003-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003666_1350650042128003-Figure1-1.png", + "caption": "Fig. 1 Schematic diagram of the four-pad tilting-pad journal bearing", + "texts": [ + " As the three-dimensional turbulent THL model by Taniguchi et al. (model B) and the two-dimensional THL model by Mikami et al. (model C) are explained in references [6] and [3] respectively, this paper omits repetition. It is noted that numerical calculations for the two models in this paper were conducted with computer programs coded by the present authors\u2019 team. Proc. Instn Mech. Engrs Vol. 218 Part J: J. Engineering Tribology J05203 # IMechE 2004 at RMIT UNIVERSITY on July 7, 2015pij.sagepub.comDownloaded from Figure 1 shows a four-pad tilting-pad journal bearing, experiments on which have been reported in reference [6]. The diameter is 479mm, and the width is 300mm. The pad clearance is 0.500mm, and the preload factor is zero. The pad arc angle is 808. The bearing load is 169.5 kN, and the load direction is LBP. The four pads are numbered counterclockwise in the direction of shaft rotation, and the top left pad is pad 1. The static characteristics of one of the two loaded pads, pad 2, are of interest. Lubricant oil VG 32 was supplied to the bearing at a temperature of 40 8C and at a pressure of 98 kPa" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000869_s0957-4174(97)00057-2-Figure4-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000869_s0957-4174(97)00057-2-Figure4-1.png", + "caption": "FIGURE 4. Vectors Pj and P/ corresponding to the desired goal configuration and an arbitrary configuration of the end-", + "texts": [ + " We now describe how g(O,i) is implemented (for a revolute joint, the case for prismatic joint is similar). It is analytically derived by symbolically differentiating the cost function c(-) in equation (4) w.r.t, qi. It is computationally advantageous to represent the goal frame w.r.t, frame ~i. Let Pj=(xj, Yi, zj), j= 1...3 denote the vectors that represent the tips of unit axes vectors ( i j and k) of ~ w.r.t. ~;i in configuration 2g, i.e. T(2g). Similarly, Pj '=( x j ' , y j ' , z / ) , i= l . . . 3 denote the vectors that represent the tips of unit axes vectors ~;~ w.r.t. ~ (see Fig. 4) in an arbitrary configuration 0 of the robot. With this formulation, the z coordinate (for a revolute joint, hence we explicitly use 0i instead of qi) remains constant and therefore simplifies the symbolic differentiation of the metric d. For example, in the expression for dx in equation (3), effector. Both are defined w.r.t. ~ d~ = (x~ - x,'* cos(0i) - y , '* sin(0i)) ~ + (y, - x~' * sin( Oi) + y~' * cos(0i))2+ (z~ -z~') 2 the term ( z ~ - z / ) : is a constant and therefore has no influence in the optimization" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000172_bf00582834-Figure2-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000172_bf00582834-Figure2-1.png", + "caption": "Fig. 2a, b. Structure of the SW interchange. Kinetochores are shown in black, a Pachytene pairing configuration. Positions of chiasmata leading to a type II chain quadrivalent are indicated by \"X\". Interstitial regions (between the kinetochores and the translocation break points) are shaded, b Type II chain as it appears at the start of prometaphase; e central kinetochores; p peripheral kinetochores", + "texts": [ + " However, the two kinetochores are so close together that they appear as one unit in living cells. Hence we will refer to a pair of sister kinetochores as one kinetochore. The interchange, designated \" S W \" (to celebrate our colleague, Suzanne Ward), involves the longest chromosome pair of the Melanoplus complement and another large- or medium-size pair. All Melanoplus chromosomes appear to have only one arm because the kinetochore is very near the end. Hence the interchange has only four regions (Fig. 2a): two short interstitial regions in 401 the center and two long translocated segments, one on either side. The quadrivalent is nearly symmetrical, but one translocated segment is slightly longer than the other. The translocated segments always form one chiasma each. The interstitital regions are indistinguishable from each other. One interstitial segment or the other commonly lacks a chiasma, which produces a chain of four chromosomes in prometaphase (Fig. 2 b). The result is a so-called \"Type I I \" chain (reviewed in Rickards 1983), with two homologous kinetochores at opposite ends (the \"peripheral kinetochores\") and the other two close together near the middle of the structure (the \"central kinetochores\") (Fig. 2b). Occasionally, both interstitial regions have chiasmata, and a ring is formed. However, these are rare and hard to track in living cells, so attention was focused on chain configurations. Four segregation types were observed in chain quadrivalents (Fig. 3). They originate from four relatively stable orientations of the quadrivalent. Two of them belong to the group of 2:2 segregations: alternate/adjacent 1 (\"AA1 \"; Fig. 3a) and adjacent 2 ( \" A 2 \" ; Fig. 3b). The third orientation type produces a 3:1 segregation (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001595_bi00297a011-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001595_bi00297a011-Figure1-1.png", + "caption": "FIGURE 1: Plot of kohd vs. [HEF] for the reduction of T2D, laccase type 1 Cu(I1): 25.0 OC, p H 7.0, and I = 0.5 M.", + "texts": [ + " The reversibility of type 2 copper removal is verified by the restoration, to 87% of the original value, of turnover activity in reconstituted T2Dr laccase. Kinetics of the Reduction of T2D Laccase by Substituted Hydroquiones and HEF. Excellent first-order analytical plots were derived from anaerobic, 614-nm stopped-flow studies of electron transfer to T2Dr-type 1 Cu(I1) from substituted hydroquinones (H2Q-X) and (hydroxyethy1)ferrocene. The substrate concentration dependences of koM (Table I, H2Q-X; Figure 1, HEF) indicate a simple rate law in each case: case. Rhus vernicifera laccase was isolated from lacquer tree acetone powder and purified by the method of Reinhammar (1 970). Type 2 copper-depleted laccase was prepared and reconstituted with cupric sulfate according to the method of Graziani et al. (1976), with minor modifications. Nonspecifically bound Cu(I1) was removed on a Chelex-100 column following the reconstitution procedure. Type 2 copper-depleted laccase prepared by the above method contains oxidized type 1 copper and a reduced type 3 site (designated T2Dr; see below)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002503_tdcllm.1993.316227-Figure8-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002503_tdcllm.1993.316227-Figure8-1.png", + "caption": "Fig. 8. Illustration of the manipulator (ground operation type)", + "texts": [], + "surrounding_texts": [ + "6 - 6", + "Also, FRP covers are installed over the external surfaces of the arms to avoid inter-phase shorts and ground faults.\nFurther, the driving force transmission system is switched from a rope drive to a shaft drive, which\n(3) Operation system As is the case with System Makeup I (On-boom Operation Type), the operator can choose between the following two operation systems to suit the individual work requirements.\n0 Twin-arm master-slave operations which are\n@ Joystick-controlled operations which permit intuitive and easy to perform. speed-regulated linear movement.\nHowever, as the operator cabin is located a t the assistant driver's seat, the cabin space is small (reduced to about half that provided in the second step). Therefore, if the master arm is of the same polar coordinate type a s the manipulator (hereinafter referred to as the slave arm), it is difficult to provide an adequate operation range. To avoid such a problem, a variously structured master-slave system (the master and slave arms differ in structure) is employed so that\nMEW27 4 5 3", + "6-6\nthe origin offset and magnification change functions can be executed as needed.\nAs a result, a rectangular coordinate type master arm, which has a simple mechanism and is easy to operate, is used so that switching between the masterslave and joystick-controlled operation methods can be effected with one controller.\nAlso, a frequently used tool drivehlide base transport switch is mounted in the gripper section to offer increased operational ease. Further, force feedback bilateral control, which is considered to be effective in work requiring fine positioning control, is exercised over the force applied in the directions of 6 axes at the end of the manipulator during master-slave operations. Therefore, the operator can feel the force applied to the manipulator and this results in enhanced operability and safety.\n(4) Monitoring equipment The monitoring function is one important point for the ground operation type manipulator system. The data derived from the first and second step verifications indicates that the image inputs from a 3- D camera and several supervision cameras are effective in performing actual operations. Therefore, efforts have been made in the third step development to provide the operator with easy-to-observe images.\n0 3-D camera The binocular parallax, whose effectiveness was verified when the second step system was designed, is employed. T w o cameras corresponding to the rightand left-hand eyes of human beings are used to shoot the object to be worked on, and the 3-D camera image monitor presents the image picked up in that manner to the right- and left-hand eyes of the operator who wears polarizing glasses (3-D glasses).\nFurther, the 8-power zoom and left-right camera translation functions are added to offer increased operational ease.\n8 Supervision cameras The top camera, as in the second step, is added to the system to offer increased visual field directions. A 6.5-inch liquid crystal display monitor is positioned on both the right- and left-hand sides of the 3-D camera image monitor. The operator can select a desired camera image t o perform proper monitoring operations.\n(5) Specialized tools The tool drive system employed is the same as indicated under System Makeup I (On-boom Operation Type).\nIn the case of the ground operation type, however, tool replacement is made on top of distribution line poles. This means that work efficiency is increased by reducing the frequency of tool replacement. To offer a host of functions (combination of wire stripping and brushing), about 15 types of tools have been developed.\nAs described above, the operator replaces the tool at the slave end with a tool in the tool box attached to the crane end by remotely controlling the manipulator. When the tool is to be replaced, entering the number of the next tool to be mounted causes the crane and tool box swing axis to automatically operate so that the tool box is moved to the front of the slave arm.\n(6) Truck The truck employed is the same as indicated under System Makeup I (On-boom Operation Type). The noise reduction power unit could not be mounted, both because it was too heavy and because vehicle space was limited. Therefore, a 3.5-ton vehicle chassis which can accommodate the system without the noise reduction power unit was employed.\n6. OPERABILITY EVALUATION (1) On-boom operation type The on-boom operation type manipulator system was used to conduct operability testing and on-site field testing in actual-scale assembling t o check four types of target operations (switch work, bypass wire cuttinghonnection, construction switch work, and transformer work), which involve many elements of high-voltage live-line work.\nAs a result, the conceptual development targets such as system operational ease enhancement, vehicle size and weight reduction, and productivity increase have been attained. Moreover, the practicability of line installation work by the manipulator has been demonstrated under complicated work conditions at actual working sites, indicating the possibility that the manipulator system can be put to practical use.\nThe working time required for each step in switch work is shown in Table 6.\n(2) Ground operation type As with the on-boom operation type, the ground operation type manipulator system was used to conduct operability testing in actual scale assembling. As shown in the following table, the ground operation type has slightly lower operability than the on-boom operation type. However, all the target operations could be performed with the ground operation type, and it was fully confirmed that the ground operation type can be remotely controlled from the ground.\nM86327\n4 5 4" + ] + }, + { + "image_filename": "designv11_65_0003017_1.1481369-Figure2-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003017_1.1481369-Figure2-1.png", + "caption": "Fig. 2 Pendulum and its bond graph representation", + "texts": [ + " For the sake of clarity the signal bonds have not been displayed for the modulations of MTF elements. Each modulus is a function of one of the two variables u1 and u2 . The application of LaCAP, lLCAP, HaCAP, and lHCAP is recapitulated in Table 1 where numbers on the different causal bond graphs indicate the step order in which causality has been propagated ~encircled numbers refer to multiple strokes and squared numbers refer to added elements!. The resulting dynamics equations are given in Table 1. Pendulum \u202012\u2021. The system and its bond graph representation are shown in Fig. 2. For the sake of clarity the signal bonds have not been displayed for the modulations of the MTF and MGY elements. The MTF moduli are functions of u. The MGY ratios are functions of the angular velocity components v1 , v2 , and v3 . The application of BHCAP and lBHCAP is recapitulated in Table 2 with the resulting dynamics equations. This paper proposes a set of alternative causality assignment procedures. It provides an algorithmic frame to derive different mechanical oriented formulations from the bond graph representation" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002020_0301-679x(86)90038-1-Figure2-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002020_0301-679x(86)90038-1-Figure2-1.png", + "caption": "Fig 2 Effect of viscosity grade on torque efficiency (manual transmission): (o) SAE 75W; (A) SAE 90; (D) SAE 140", + "texts": [ + "2 kg without changing the test piece at each load step. The test pieces used were the Timken blocks and cups specified by ASTM D 2782. The viscosities of the test oils and the composition of the EP and friction reducer additives are presented in Tables 2 and 3. Results and discussion Effect of gear oil viscosity on transmission torque efficiency Single v iscosi ty grades The transmission torque efficiency at speeds between 20 and 80 km/h with loads corresponding to 0 and 10% slope for oils of varying viscosity are shown in Fig 2 for the manual transmission and Fig 3 for the rear axle. The same viscosity grade of oil was used simultaneously in each gear box, the oil used in the manual transmission was of API G L ~ classification and that used in the rear axle TRIBOLOGY international 313 Kubo et al - transmission efficiency was of API GL-5 classification. The same additive package, A, was used but the treatment rate was different. To ensure that differences in efficiency measured were due to the change of viscosity grade and not due to some running-in effect, the efficiencies were measured for a standard SAE 90 oil before and after the above test" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000012_piae_proc_1922_017_029_02-FigureI-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000012_piae_proc_1922_017_029_02-FigureI-1.png", + "caption": "FIG. I8 . -ha r hub for 25/50 h.p. car.", + "texts": [], + "surrounding_texts": [ + "374 THE INSTITUTION OF AUTOMOBILE ENGINEERS.\nThe point arises here as to how the hand of the gear should be selected for the bevel or crown-wheel so as to impose the heavy thrust on one or the other. The conclusion just reached regarding the double-purpose unit at the crown-wheel is on the assumption that the light thrust load oome's in this position, and this is in accordance with general practice. In order to make full use of the possibilities of the single-row bearing, however, the hand of the gear might be reversed, and certainly there will, in many cases, be more opportunity for obtaining a large capacity in the crownwheel position than behind the bevel-pinion\nPilot Bemirig. Although the pilot bearing, Fig. 20, does not appear on any of\nthe graphs, it will be remarked that it is in many casea an& example of a bearing applied to relieve a heavily loaded one without kaking into account the possibility of its own failure, which frequently occurs.\nClutch. The bearing selection for the clutch is not so ainenable to the same treatment, but it will usually be found that a light type single-row bearing w i t b u t fillibg-slot has su5cient capacity .to deal with a load which is only in occasiond operation.\nFigs. 17 to 20 show the aotual arrangements of the front-hub, rear-hub. gear-box and rear-axle bearings of the car referred to as No. 12.\nSUMMARY. The difficulties in the way of fixing definite factors of safety which will meet all cases are fully appreciated, but it is hoped that the preceding recommendatipns will form a useful basis for investigating existing designs. For convenienoe, the factors have been collected in Table I., the letters in Fig. 15 qorres o d ' with those used thmughout the drawings of the Sizaire-brm? chassis.", + "THE ENDURANCE ,OF BALL HEARINGS. 375", + "376 THE INSTITUTION OF A UTOMORILE ENGINEERS.\nFIG. lg.-Gear-bux for 25/50 h.p. car." + ] + }, + { + "image_filename": "designv11_65_0003621_jsen.2003.820358-Figure2-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003621_jsen.2003.820358-Figure2-1.png", + "caption": "Fig. 2. Piezoelectric ring loaded and polarized along the axis of symmetry.", + "texts": [ + " The principal of conservation of energy and the assumption that energy is not lost due to friction or electrical resistance leads to the following equation: (6) where is the work due to displacement or the total energy imparted to the piezoelectric element, is the total mechanical energy stored in the element, and is the total electrical energy stored in the element. Further assuming that the element behaves within its elastic range, then can be expressed as [6] (7) where is the volume of the piezoelectric element. The term can be expressed as (8) Substituting (7) and (8) into (6) yields (9) which is the total amount of electrical energy that can be stored in a piezoelectric element. 3) Geometrical Consideration: When a piezoelectric ring (Fig. 2) is loaded parallel to the axis of symmetry, it is in a state of approximately constant stress through the thickness. If this ring was also polarized in the axial direction, the above described constitutive model can be applied. When installed as a mold cavity pressure sensor, the piezoelectric ring needs to be protected against thermal stress from the high-temperature molten plastic. A cap design was employed in the presented study, which decouples the ring from the mechanical force acting on the ring. The force placed on the ring can be expressed as (10) where is the pressure acting on the cap surface. Stress in the piezoelectric ring can then be calculated as (11) where is the cross-sectional area of the piezoelectric element (12) From (9)\u2013(12) the energy associated with applying a given pressure to the ring structure in Fig. 2 can be calculated as (13) The voltage across the electrodes in an open circuit is readily calculated using the definition of the voltage constant (14) Similarly, the charge can be calculated using the definition of charge constant : (15) The open and closed circuit displacements can be calculated from the definition of compliance (16) (17) where and are related to each other through the piezoelectric coupling coefficient [6]: (18) The utility of the above derived linear model is subject to several constraints", + " The complexity of the preceding analytical model suggested the use of a numerical model to assist in the analysis of energy extraction mechanism. A finite element model was thus developed, which took into account the neglected effects noted above. Specifically, it incorporated the body inertial effects and material damping, and provided a detailed description of the transient process experienced by the energy extraction device when subjected to high strains from the molten polymer. Given the geometrical structure of the piezoelectric ring (Fig. 2), a two-dimensional (2-D) axisymmetric static coupled-field analysis was performed using the finite element software package ANSYS [12]. A convergence study was conducted to determine the appropriate element size that allows for a good balance between accuracy and computational time. It was determined that an element edge length of 0.2 mm corresponding to four elements through the thickness direction of the ring and a total solution time of below 1 s satisfied the criteria for the 2-D model. To verify the validity of the 2-D model assumption, a three-dimensional (3-D) coupled-field analysis was also developed and a convergence check performed" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003198_robot.1997.606773-Figure3-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003198_robot.1997.606773-Figure3-1.png", + "caption": "Fig. 3. Equivalent kinematic scheme for the S-10 robot.", + "texts": [ + " The second case illustrates the inverse situation, when the robot control system plans the straight movement between the points & and X,, and, due to non-exactness in the kinematic model, this causes the end effector to move along the curve X,X,. The third and the fourth cases are worth including in the simulation to test the real and standard kinematic problems. For these two cases, the line at the input corresponds to the line at the output. Example. Robot GMFanuc S-10 with the standard parameter values (Table l), [8] has been examined. Figure 3 demonstrates the kinematic scheme of this robot. The parameters of the \u201creal\u201d model were assigned the random values in such a way that the linear parameters differ by no more than 1 mm, and angular parameter no more than 1 degree from the standard values. The robot control system plans the movement between the points &=[537.753 269.433 292.828IT and &=[237.753 569.433 792.828]* mm. The correspondent positions of the end effector are XA=[575.516 230.434 325.806IT and XB=[283.956 527.791 823.271IT mm" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003946_aict.2005.53-Figure7-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003946_aict.2005.53-Figure7-1.png", + "caption": "Figure 7. Traffic balance obtained by DRMAP", + "texts": [], + "surrounding_texts": [ + "[1] S. P. M. Choi, D. Y. Yeung. Predictive Q-Routing: A memory-based reinforcement learning approach to adaptive traffic control, Advances in Neural Information Processing Systems, 8:945-951, Editors: Touretzky D. S., Mozer M. C., Hasselmo M. E., MIT Press, Cambridge, 1996. [2] W. J. Dally. Virtual channel flow control. IEEE Transactions on Parallel and Distributed Systems, 3:194\u2013205, 1992. [3] W. J. Dally, C. L. Seitz. The torus routing chip. Journal of Distributed Computing, 1:187\u2013196, 1986. [4] A. Kasprzak. Packet switching wide area networks. Oficyna Wydawnicza Politechniki Wroc\u0142awskiej, Wroc\u0142aw, 1997 (In Polish). [5] L. P. Kaebling, M. L. Littman, A. W. Moore. Reinforcement learning: a survey. Journal of Artificial Intelligence Research, 4:237\u2013285, 1996. [6] P. Kermani, L. Kleinrock. Virtual Cut-Through: A new computer communication switching technique. Computer Networks, 3:267\u2013286, 1979. Proceedings of the Advanced Industrial Conference on Telecommunications/Service Assurance with Partial and Intermittent Resources Conference/ELearning on Telecommunications Workshop 0-7695-2388-9/05 $20.00 \u00a9 2005 IEEE [7] S. Koenig, R. Simmons. Complexity analysis of real-time reinforcement learning applied to finding shortest path in deterministic domains. Machine Learning: A Special Issue on Reinforcement Learning, 12: 234\u2013345, 1997. [8] T. Mitchell. Machine learning, McGraw-Hill International Editions, 1997. [9] A. W. Najjar, A. Lagman, S. Sur, P. K. Srimani. Analytical model of adaptive routing strategies. Technical Report CS94-105, Colorado State University, 1994. [10] R. Rudek. Q-Learning based algorithms for solving unknown structure multi-criteria graphs. Proc. of the IIIth Polish Conference on Computer Pattern Recognition Systems KOSYR, Mi\u0142ko\u0301w, 2003, pp. 190\u2013195 (In Polish). [11] R. Rudek. Experimentation system for evaluating and analysis of adaptive routing algorithms. M.Sc. Thesis, Wroc\u0142aw University of Technology, Wroc\u0142aw, 2004 (In Polish). [12] P. Stone, M. Veloso. Multiagent systems: a survey from a machine learning perspective. Autonomous Robots, 8:345\u2013 383, 2000. [13] R. Sutton, A. Barto. Reinforcement learning, MIT Press, Cambridge, 1998. [14] M. Tan. Multi-agent reinforcement learning: Independent vs. cooperative learning, Morgan Kaufmann, 1997. [15] A. S. Tanenbaum. Computer network, Prentice-Hall, 1988. [16] J. Upadhyay, V. Varavithya, P. Mohapatra. A traffic- balanced adaptive wormhole-routing scheme for twodimensional meshes. IEEE Transactions on Computers, 2:190\u2013197, 1997. [17] C. Watkins, P. Dayan. Q-learning. Machine Learning 8:279\u2013 292, 1992. Proceedings of the Advanced Industrial Conference on Telecommunications/Service Assurance with Partial and Intermittent Resources Conference/ELearning on Telecommunications Workshop 0-7695-2388-9/05 $20.00 \u00a9 2005 IEEE" + ] + }, + { + "image_filename": "designv11_65_0000431_s0094-114x(97)83002-9-Figure8-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000431_s0094-114x(97)83002-9-Figure8-1.png", + "caption": "Fig. 8. A robot following a three-straight-line path.", + "texts": [ + " 6 where the mode shape is normalized such that the mode amplitude at station 1 (see Fig. 2) is equal to unity. Applying the solution procedures outlined in Fig. 1, the transverse deflection due to the specified motion results, as shown in Fig. 7. These curves are considered to be elastodynamically-induced errors of the manipulator under investigation, i.e. the unwanted deviation of the robot 's position (or trajectory) because of its flexibility. Example 3: Continuous multi-straight-line trajectory Figure 8 depicts a top view of the RP manipulator (dashed), with a prescribed multi-straight-line continuous path to be followed. Suppose that the manipulator is required to move from P0 to P, , passing through via points P,, P2 . . . . . P,_ ~. For brevity, only two via points, or three straight lines (n = 3), along the entire path, are shown in the example of Fig. 8. It is assumed that the robot traces out a straight line between each pair of two neighboring points (e.g. POP,, P~P2, etc.) and that the revolute joint maintains a smooth motion profile O(t), while the longitudinal motion is determined according to a segment-by-segment straight line path with the manipulator members treated as rigid. Such a determination may be carried out, one after another, for each triangular segment bounded by three lines connecting two adjacent intermediate points on the path, P,_ ,, Pk, and the pivot point for the revolute joint, O" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003115_746901-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003115_746901-Figure1-1.png", + "caption": "Fig. 1. Caricature of rigid-body motion of the sight and barrel while target tracking as the tank traverses a bump.", + "texts": [ + " Superimposed on the unique manufactured barrel centerline is the flexed barrel shape that can occur prior to firing while the vehicle is on the move. In order to understand and quantify the effects of barrel flexure on gun accuracy, it is necessary to determine what combination of fundamental mode shapes is most likely to occur. A method to accomplish this task is described in this paper. In order to maintain target tracking while on the move, the gun barrel of a main battle tank is subjected to translational and rotational motion, as illustrated in Fig. 1. In brief, the barrel is rotated by the angle, \u03b8, that is required to keep the gunner\u2019s line of sight pointed at the target while traversing uneven terrain. Rotation is produced by the action of a hydraulic actuator located on the breech side of the trunnion axis, as depicted in Fig. 2. In addition to the rigid-body motion, flexural modes of motion will be excited by vehicle motion. Thus, the dynamic state of the prefiring barrel can be described by the following: y(x, t) = rigid body \ufe37 \ufe38\ufe38 \ufe37 yt(t) + \u03b8(t)[x \u2212 xt] + Y (x, t) \ufe38 \ufe37\ufe37 \ufe38 flexing mode , (1) where y gives the vertical component of the lateral barrel displacement (relative to the static barrel centerline) at time t in an earth-fixed coordinate frame; x specifies the axial barrel coordinate relative to the breech face (xt is the location of the trunnion); yt gives the vertical displacement of the trunnion axis; \u03b8 specifies the rotation of the barrel about the trunnion axis; and Y gives the vertical component of the lateral barrel displacement (relative to the static barrel centerline) in a coordinate frame that is rotating and translating with the barrel" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002778_jpdc.2001.1820-Figure2-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002778_jpdc.2001.1820-Figure2-1.png", + "caption": "FIG. 2. Rigid multibody serial chain.", + "texts": [ + " (12) as POi \u00bc #SOi; cm\u00bdIi; cmbOi \u00fePi; cm : \u00f014\u00de Expanding and simplifying, POi \u00bc IOi \u2019#S T Oi; cmVOi \u00fe \u2019IOiVOi: \u00f015\u00de It then follows that the force at point Oi is given by FOi \u00bc IOi \u2019VOi \u00fePOi: \u00f016\u00de Our interest now shifts to the computation of the EOM for serially coupled rigid bodies considering the equations obtained for single bodies (serial chains of n rigid bodies are considered). For simplicity and clarity, it is assumed that all bodies have the same number of DOF and degrees of constraint (DOC). Considering that each element/link of the serial system in Fig. 2 is a rigid body with defined cm location and inertia tensor, then its mass distribution is fully characterized. Also, from the fact that each element is considered rigid it follows that the EOM can be developed on any point of the body, as shown in Section 3.1. For convenience, this point is set to be the connection point between adjacent bodies, Oi. The procedure consists of a forward, base to tip, propagation of kinematics parameters \u00f0V ; \u2019V \u00de and a backward, tip to base, propagation of spatial forces (F) considering d\u2019Alambert\u2019s principle" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002711_a:1014049410772-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002711_a:1014049410772-Figure1-1.png", + "caption": "Figure 1. Unicycle kinematic model.", + "texts": [ + " Results of experiments with a real robot employing the discussed three methods are described in Section 6 where the robot is approaching a static goal, following a moving target and driving through a fast changing environment. The final Section 7 gives some conclusions drawn from these experiments. The model, describing the motion of the cartesian unicycle vehicle is given by x\u0307 = u cos \u03c6 y\u0307 = u sin \u03c6 (1) \u03c6\u0307 = \u03c9 with u being the linear velocity in the direction of \u03c6 and \u03c9 the angular velocity (Fig. 1). In a point-to-point navigation task, the vehicle starts at point (xS, yS) with heading \u03c6S and should be driven with appropriate u and \u03c9 to the goal. Without loss of generality the goal can be chosen to be (xG, yG, \u03c6G) = (0, 0, 0). Since on-line, reactive, and realtime systems are considered here, u and \u03c9 cannot explicitly depend on time but only on the state variables thereby reducing the control system to a set of autonomous differential equations of the state variables. Brockett\u2019s Theorem (Brockett, 1983) proves that the system (1) cannot be stabilized using smooth timeinvariant feedback" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000768_0957-4158(95)00074-7-Figure2-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000768_0957-4158(95)00074-7-Figure2-1.png", + "caption": "Fig. 2. Tapered three-lobe (a) and elliptic (b) pistons in eccentric and tilted position.", + "texts": [ + " Piston Cylinder block ~ \\ / / ~ Efficiency of the axial piston motor 285 for the shaded zone and r = ( - x 2 q- V [ x ~ - x 2 - s ~ -~- 4r2])/2 with (2) x2 = Xl cos (0 - )') + sl sin (0 - ~,) (3) for o the r regions, whe re Xl, Xz, r l , r2 and sl are the th ree - lobe pa rame te r s . The film thickness hi is given by hi = e c o s 0 - r + ~ / [ (ecos 0 - r ) 2 + 2 e r c o s O - r : + R z - e2], (4) whe re e is eccentr ici ty of the pis ton in the cyl inder bo re and R is radius of the bore . Cons ider ing a t ape r t on the pis ton, the fi lm thickness var ia t ion a long the length x can be wri t ten as h = h I + a'x (5) 286 with and K. SADASHIVAPPA et al. = t/1 = d h / d x (6) h: = hi + t. (7) Similarly, for an elliptic profile, from Fig. 2b, the radius is given by r = Via 2 sin 2 (0 - y) + b 2 cos 2 (0 - - y ) ] (8) and the film thickness is given by hi = ~/[R 2 + e 2 + 2ercosO] - r . (9) Substituting h a from Eqns (4) and (9) in Eqn (5), the film thickness variation along x is obtained. 2.2. Viscous friction force and leakage f l o w rate Calculation of the viscous friction force and leakage flow rate in the annular clearance between the piston and cylinder of hydraulic components is an important requirement when assessing the performance of such components" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002211_0-306-46956-1_9-Figure6-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002211_0-306-46956-1_9-Figure6-1.png", + "caption": "Figure 6 Side force and aligning moment for isotropic brush model tire in pure sideslip for a particular load.", + "texts": [ + " For small slip angles, the ratio of the aligning moment to the side force, called the pneumatic trail, is 1/6 of the contact length. As the lateral slip increases and sliding extends further forward in the contact region, of the tread base structure at times zero, and are shown in plan view. Bristle distortion is proportional to distance back from the first point of contact; tan(slip) = the pneumatic trail decreases, reaching zero when sliding extends over the whole contact patch and the tire is generating its peak side force (Fig. 6). The compliance of real tires, which has been represented as belonging entirely to the bristles of the brush model, resides substantially in the tire carcass. The structure is very different longitudinally from laterally, so that the real tire is non-isotropic. The pure slip properties are qualitatively similar but quantitatively different, and the rather simple combined slip properties of the isotropic brush are substantially more complex for real tires. The simplest way of thinking about the carcass flexibilities is to imagine that the circumferential ring of the tire is rigid, taking the place of the undeformable base structure of the brush, and that it is connected to the wheel hub by an elastic restraint system allowing it all six degrees of freedom, relative to the hub" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003764_rob.20045-Figure9-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003764_rob.20045-Figure9-1.png", + "caption": "Figure 9. Spatial 6-link manipulator.", + "texts": [ + " In the optimal motions, the manipulator touches slightly the surfaces of the obstacles and this may be considered as imperfect obstacle avoidance. The minimum clearance to assure safe avoidance must be added to the actual geometric sizes of obstacles. Figure 8 shows the saturation of the actuators. In the case of no obstacle, joint 2 [dotted line in Figure 8(a)] is not fully saturated. On the other hand, all joints are almost saturated in Figure 8(b) to avoid the obstacles by the local minimum (F). Figure 9 shows a configuration of a PUMA 560 type manipulator at zero-displacement. All joints are revolute pairs around their z-axes. The base coordinates are the same as the first link coordinates at zerodisplacement. Link 4 is connected to link 3. The link parameters in Denavit\u2013Hartenberg notation and the specifications are listed in Tables 4 and 5, respectively, where gravity is acting in z0 direction and c is about twice the static actuator torques necessary to endure gravity in fully stretched configurations" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003666_1350650042128003-Figure8-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003666_1350650042128003-Figure8-1.png", + "caption": "Fig. 8 Variation in the maximum pad surface temperature with shaft speed according to Taniguchi et al. [6] (reproduced by the present authors)", + "texts": [ + " Engineering Tribology J05203 # IMechE 2004 at RMIT UNIVERSITY on July 7, 2015pij.sagepub.comDownloaded from 3000 to 3300 r/min, which reasonably corresponds to the temperature inflection with shaft speed shown in Fig. 4. Figure 7 shows the variation in the maximum pad surface temperature with bearing load. The prediction of model A is in better agreement with measurements than those of model B and model C. To sum up, model A is better than model B and model C in predicting the thermohydrodynamic performance of journal bearing operated in the super-laminar region. Figure 8 shows the theoretical and experimental variation of the maximum pad surface temperature with shaft speed obtained by Taniguchi et al. [6]. The prediction of model B is in good agreement with measurements qualitatively and quantitatively, which is different from Fig. 4. This results from the difference of cross-film slicing of the calculation domain. The present authors used 100 slicing layers, while Taniguchi et al. used only eight slicing layers. In addition to the cross-film slicing number, numerical calculations for turbulent THL analyses are sometimes found to be sensitive to various factors, such as the methods of making equations discrete, values of over- or under-relaxation coefficients and convergence criteria, and solutions significantly different from each other are sometimes obtained, depending on the factors" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000339_1350-6307(95)00030-5-Figure2-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000339_1350-6307(95)00030-5-Figure2-1.png", + "caption": "Fig. 2. Finite element model of the bogie frame.", + "texts": [ + " 1. Semi-Loof shell elements and beam element. elements. The displacements and rotations are assumed to vary cubically along the element, giving 24 independent degrees of freedom. Seven of these are statically eliminated to give 17 degrees of freedom actually connected to other elements. There are six degrees of freedom at each of the two end nodes of the element, three degrees of freedom at the central node, and two torsional rotations at two Loof nodes. The beam element is shown in Fig. 1(c). Figure 2 shows the whole finite element model of the bogie frame. The eight and six-noded shell elements, together with the three-noded beam elements, are used to idealize this box frame. The element mesh has been optimized to fit the stress analysis, and to reduce the instantaneous front size to suit the computer facilities. The model consists of six groups of elements, a total of 1712 elements and up to 21,526 degrees of freedom. The PAFEC (Program for Automatic Finite Element Calculations [4] package is used to carry out the structural analysis" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003347_978-1-4020-2249-4_32-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003347_978-1-4020-2249-4_32-Figure1-1.png", + "caption": "Figure 1. a) 6 dof. Parallel Kinematic Maschine Linapod with fixed length legs. b) Geometry of one leg of the PKM Linapod. P, U, S denote prismatic, universal and spherical joints, respectively.", + "texts": [ + " This can be done by minimizing the condition number of the error amplification matrix following the approach for the Jacobian of the TCP (Ranjbaran et al., 1996). In this section we give an examples of a 6 dof fully parallel manipu lator where no closed form solution for direct kinematics is known. A performance comparison is presented. The 6 dof parallel kinematic machine tool Linapod (Wurst, 1998) was developed at the Institute for Control Engineering of Machine Tools and Manufacturing Units at the University Stuttgart (Germany) and is depicted in Figure 1a. Six rigid links connect the mobile platform to the fixed frame with spherical/universal joints. The lengths of the upper and lower links are lu = 1. 7m and II = 1.25m, respectively. The pivot points on the frame are actuated with linear motors, moving all parallelly to the z-axis. The radii of the the upper and lower platform pivot points and the base frame are ru = 0.2m, rl = 0.22m and rb = 0.886m, respectively. The distance between the two plains of the platform is tlh = 0.2m and the triangles of the pivot points have a mutual rotational offset of 70\u00b0. We define the home position in the center of the workspace and assume small errors in the geometry of every leg. Applying the algorithm from section 3, we can calculate the error amplification matrix Jl for errors in the length of the bars. Therefore, we have to solve the direct kinematic problem to get the position Ti and the direction Ui of the six legs with respect to the TCP (Figure 1 b ). Then we calculate the internal forces in the bars from force equilibrium conditions when applying the unit forces/torques to the TCP, and we get the matrix equation where Pi = Ui x Ti, and the vectors Ii are the internal forces in the direction of each leg, respectively, that result from the unit wrenches applied to the TCP. Here we find Jr = F = A-I. With the parameters of the Linapod Jl becomes at its home position 0.058 0.558 -0.616 -0.009 0.567 -0.557 -0.678 0.389 0.288 -0.649 0.316 0.333 ff= -0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002490_s1474-6670(17)60988-1-Figure5-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002490_s1474-6670(17)60988-1-Figure5-1.png", + "caption": "Fig. 5. Non-optimal (a) and or,timal (b) motions of the robot 'Universal\".", + "texts": [ + " Optimization was perfor med numerically by means of a computer. As an example we present here some re sults for two Soviet industrial robots. For the robot \"Universal\" we considered two-dimensional motions of the load in the horizontal plane while for the other robot having three links we considered two-dimensional motions in the vertical plane. In both cases we were interested only in the position of the centre of the load but not in its orientation. We have here n = 3, m = 2, so that there is one redundant degree of freedom in both cases. Figure 5 illustrates some results of op timization for the robot \"Universal\". Its three active degrees of freedom corres pond here to rotations of the platform and of the arm about their (vertical) axes and to the horizontal translation of the arm. The transfer of the load from the position r O (xo = 1.4m, yO = 0) to the position rl (xl = 0, yl = = - 1.4m) is presented here. This trans fer may be performed by the simple rota tion of the platform by the angle of 900 , see Fig. 5a. This motion takes the time Tl ~ 3.6 sec. The optimal control regime obtained by solving Problem 5 requires T2 ~ 2.1 sec. which is 41% less than Tl \u2022 Optimal initial and terminal configura tions for this regime are shown in Fig. 5b. Figure 6 shows some optimal regime for the robot of anthropomorphic type with three links. Here a non-optimal mo tion consists in rotation of the last (smallest) link while for the optimal mo tion all three degrees of freedom are ac tive. A number of similar optimal regimes were calculated, and conSiderable gain in the time of the motion was obtained. Experi mental realization of optimal regimes shows a good agreement between calculated and real times of motions. CONCLUSIONS We considered some problems of optimal control for robots" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002335_20.877741-Figure10-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002335_20.877741-Figure10-1.png", + "caption": "Fig. 10. Induction motor (thickness: 42 mm).", + "texts": [ + " Therefore, the effect of the discontinuity of potential on the electromagnetic force cannot be neglected, because higher accuracy is required for force calculation. Fig. 9(b) shows that the spurious oscillation in the discontinuous method can be reduced if the fine mesh is used. Therefore, in order to evaluate the electromagnetic force using the nonconforming mesh, the continuous method, or the discontinuous method with fine mesh, should be used. A nonconforming technique is applied to a 2D magnetic field analysis in an induction motor in order to investigate its effectiveness . Fig. 10 shows the model of an induction motor without skew. The motor is one of the verification models proposed by the IEEJ. The detailed data of this motor is shown in Table I. The rotor and stator cores are made of the silicon steel 50A1300 and the nonlineality is taken into account. The conductivity of the aluminum rotor bar is assumed to be S/m by revising the conductivity to 2-D analysis. Only 1/2 of the whole region is analyzed due to symmetry. Fig. 11 shows the mesh with the first order quadrilateral elements" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002100_pime_proc_1986_200_139_02-Figure11-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002100_pime_proc_1986_200_139_02-Figure11-1.png", + "caption": "Fig. 11 Sliding velocities in spinning and slipping contacts", + "texts": [ + " l t ~ 5 10 15 20 Axial load kN 0 0 I By imposing fixed values of slip on the inner race contact it has been possible to determine how the forces acting within the bearing changed with slip to produce the characteristic sudden change into the slipping regime . this mechanism can no longer operate. At this stage the change in input power must be accompanied by a corresponding change in churning. Since this falls only slightly with slip very large reductions in cage speed are necessary to maintain equilibrium. This behaviour can be explained by consideration of the contact velocities (Fig. 11). When only spin is present the forces acting on either side of the contact ellipse are equal and opposite, giving rise to a moment resisting spin. However, as the sliding velocity becomes greater than the spin velocity the forces both act in the same direction but with different magnitudes, thus maintaining the moment about the centre of the ellipse, albeit at a reduced value. It can be shown that for this bearing at 20 per cent surface slip the slide-roll ratio on both sides of the ellipse will be greater than 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002218_s026357470000120x-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002218_s026357470000120x-Figure1-1.png", + "caption": "Fig. 1. Gripping frame (R) of a manipulator M.", + "texts": [ + " Several such cases cannot be solved by the use of only one arm. The second section proposes a solution to the problem of the gripping modifications; the generality of the study permits one to consider the case of any number of co-operating robots. The implementation of the method is then described in the last section. Two examples are shown to demonstrate the possibilities of the proposed solution. II. GRIPPING A PART CORRECTLY THANKS TO SEVERAL ROBOTS4 /. Definitions Consider the manipulator M of Figure 1. R = (0, X, Y, Z) is a unitary frame, linked with the terminal device of M, and it is called a \"gripping frame of M\". * Lamm-Ustl Place E. Bataillon 340060 Montpellier (France) ** Ministere de I'lndustrie et de la Recherche, direction des industries electroniques et de l'informatique, 32, Rue Guersant, 75017 Paris (France) Let us term by \"gripping frame of a part P by a manipulator M\" every unitary frame R( linked with P, which can be theoretically made to coincide with R (Figure 2). Those frames depend on the M's terminal device structure (magnet, grippers," + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002720_robot.1992.220099-Figure4-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002720_robot.1992.220099-Figure4-1.png", + "caption": "Figure 4: Probability of collision for first example, front view.", + "texts": [], + "surrounding_texts": [ + "We consider the derivation of the conditional probability of collision, conditioned on the past and present observations. The derivation and calculation is done for increasing degrees of complexity. We investigate the case of errors only on the position component X , of S,, then the more general case of errors in U, (translation errors in position, speed and acceleration), and finally the global case of errors in all components of the kinematic state S, (translation and rotation). This paper present only the simple case of errors in position. Detailed derivations for the other cases are given in [3]. In this case the errors in translation velocity and orientation can be considered negligible compared to the position error on X , . This corresponds to setting the evolution model translation errors V,\" and rotation errors v,\" to (3) In the target frame the vector DL, i.e. the relative position of the robot with respect to the target is given by the change of reference DL = R-a,(Dn - X n ) (4) where R , denotes the rotation matrix of angle a. Dn and a, being fixed, the random vector DL depends only on X , . We may simply write DL = L1(Xn) . Using the previous equation along the target m e tion model we show that the uncertainty at present time on the position of the robot in the target reference and the uncertainty on the motion model evolution are carried to the relative position of the robot at the next time instant, as where: (5) 2442 x, ~ h ( D h ) = R(a,,-an+l)DL+R-an+l ( D n + l - D n - X n - - ) ~53(Vn) = -R-an+I V,X* (7) D i and V: are independent since D i depends only on X, and as a result of the stochastic model embedded in the Kalman filter formulation, X, and V,\" are independent random'vectors. V,\" has a Gaussian d i s tribution, i.e. f\"$(VX) = +l,r:,(vX) with T: given by the probabilistic model. As discussed before in section 2.2, X, is Gaussian. Therefore, D i being an affine function of X, is also Gaussian fD: (D') = N ( E [ D : ] , C f ' ) ( D f ) with mean E[Di ] and variance E!' equal to (9) D' X E n = R-anCn Run where X,,, and E: are output of the Kalman Filter. We now express the probability of collision. We have a collision event for all pairs (DL,DL+l) for which D&+, belongs to the shadow of the target with respect to DL. By (5) it is equivalent to consider all pairs (DL, V:) such that V,\" E D ( D L ) where the domain D(DL) is given by 'D(Di) = { v: : LZ(Di) + L3(V,x) E Sh(DE)} * (10) Finally, using (10) and independence of the vectors, PC( Dn Dn + 1 ) f v2 (VX)dVx dDf (11) V V x eV(Dr) 1 = l v D f D : ( D ' ) (1 Computation is done as follows. The first step consists in the determination of the shadow with respect to a given D i , which can be done beforehand by calculation of the visibility graph of the polygonal target. The above integral is numerically approximated but some of the burden in this computation can be alleviated by direct analytical solution of the innermost integral (basically the integral of a Gaussian over a polygonal region). Alternatively a Monte Carlo method can be used on a SIMD machine to obtain a fast approximation. We give here examples of typical probability of collision profiles for different situations and error values in the case of position error, calculated for a grid of candidate destinations. The shape of the target used is described in Figure 2. The setup is as follows (all vectors are given in the global coordinates system.) The target center of inertia estimated position X,, its covariance E: are assumed to be x, = [ 0\":] ; E; = (12) The target is first assumed immobile: [ ;: ]x, = x, = The current and next target orientation angle: The target position error covariance: 2443 The robot current position is and plans to reach inside the target concavity. For conditions stated above we give the resulting probability of collision associated with a set of candidate robot destinations. This set of candidate destinations is given on a 25 samples grid centered on the position Dn+l = [-0.5,0.0] and of width equal to 1.00. The plot of the probability of collisions is given in Figures 4 a n d 3 . A second example is examined under the same conditions but the target is given now a rotational movement during the time interval of :. The corresponding results are given in Figures 6 and 5. 4 Optimal Destination The type of robotic operation to be accomplished on the target defines a subset of operational s ta t e s AO, comprising the set of states (relative position or kinematic states for instance) for which this operation can be successfully carried out. We measure in a probabilistic sense how close we are to the operational state and incorporate this with the probability of collision in a general cost function. For simplicity, AOcould consist of the positions in the target frame of reference where the robot can.act on the target. The destination Dn+l is operational if DLtl E AO; the conditional probability of being operational is The optimal solution of the problem of trajectory planning defined previously is the destination Dn+l that minimizes a composite cost function as 2444 References Di+1= Argmin(yPC(Dn,Dn+1) + (1 - y)( l - Po(Dn+l))) VDn+1 E Cn+l [l] A.Basu. A framework for motion planning in the presence of moving obstacles. Technical Report CS-TR-2378, Center for Automation Research, University of Maryland, 1989. with 7 E [0,1], where vu 1 (15) The search space Cn+l is reduced to those destinations satisfying the dynamic constraints on the robot. The search for a suboptimal solution (for example a destination meeting an acceptable level of collision probability) might be preferable to balance the cost associated with the computation of the probability of collision by using a scheme that interleaves search and calculation. 5 Conclusion For many robotic applications (robotic satellite maintenance, autonomous vehicle or industrial robot guidance, rotorcraft navigation.. ), the dynamic risk assessment is essential. While the risk is interpreted here as collision risk, we can extend this notion to other hazards that the robotic system might encounter which can be expressed in terms of the relative states of the robot with respect to a target or obstacle. This paper has studied an approach to probabilistic navigation from sensors in dynamic environments considering as optimality criteria the probability of colliding with an obstacle and the probability of accessing an operational state. We describe a computationalframework in which a probability of collision can be effectively derived. This probabilistic description is appropriate for its use with classical decision theoretic tools and can be integrated in a low level trajectory controller that operates with a higher level planner; the high level planner would consider navigational subgoals such as moving toward a position in space, locking in translation or rotation with a target object, going into orbit around a moving object,.. and provide associated levels of acceptable risk. (I4) [2] C.K.Yap. Algorithmic motion planning. In J.T. Schwartz and C.K. Yap, editors, Advances in Robotics vol. 1: Algorithmic and Geometric Aspects. Lawrence Erlbaum, Hillsdale, N.J., 1986. [3] P.Burlina D.Dementhon and L.S.Davis. Navigation with uncertainty: I. reaching a goal in a high risk region. Technical Report 565, Center for Automation Research, University of Maryland, June 1991. [4] G.S.Young and R.Chellappa. 3d motion estimation using a sequence of noisy stereo images: Models, estimation, and uniqueness results. IEEE Trans. PA MI, 12( 8) :735-759, August 1990. [5] T.H.Wu G.S.Young and R.Chellappa. A simple kinematic model based approach for 3d motion and structure estimation. Technical Report 2755, Center for Automation Research, University of Maryland, September 1991. [6] H.P.Moravec. A bayesian method for certainty grids. Technical report, Robotics Institute, Carnegie Mellon University, 1988. [7] J.Borenstein and Y.Koren. Real time obstacle avoidance for fast mobile robots in cluttered environments. In Proc. IEEE Conf. Robotics and Automation, pages 572-77, 1990. [8] J.T.Schwartz and C.K.Yap. Algorithmic and Geometric Aspects of Robotics. Laurence Erlbaum Associates Publishers, 1987. [9] N.C.Griswold and J.Eem. Control for m e bile robots in the presence of moving objects. IEEE Transactions on Robotics and Automation, 6(2):263-68,1990. [lo] 0.Khatib. Real time obstacle avoidance for manipulators and mobile robots. International Journal of Robotics Research, 1(5):90-99, 1986. [ll] T.J.Broida S.Chandrashekhar and R.Chellappa. Recursive 3d motion estimation from a monoeillar camera sequence. IEE Trans. Aero. and Elect. S ~ S . , 26(4):639-655, July 1990. 244.5" + ] + }, + { + "image_filename": "designv11_65_0003115_746901-Figure2-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003115_746901-Figure2-1.png", + "caption": "Fig. 2. Rigid-body motion of the barrel due to terrain and actuator forces.", + "texts": [ + " A method to accomplish this task is described in this paper. In order to maintain target tracking while on the move, the gun barrel of a main battle tank is subjected to translational and rotational motion, as illustrated in Fig. 1. In brief, the barrel is rotated by the angle, \u03b8, that is required to keep the gunner\u2019s line of sight pointed at the target while traversing uneven terrain. Rotation is produced by the action of a hydraulic actuator located on the breech side of the trunnion axis, as depicted in Fig. 2. In addition to the rigid-body motion, flexural modes of motion will be excited by vehicle motion. Thus, the dynamic state of the prefiring barrel can be described by the following: y(x, t) = rigid body \ufe37 \ufe38\ufe38 \ufe37 yt(t) + \u03b8(t)[x \u2212 xt] + Y (x, t) \ufe38 \ufe37\ufe37 \ufe38 flexing mode , (1) where y gives the vertical component of the lateral barrel displacement (relative to the static barrel centerline) at time t in an earth-fixed coordinate frame; x specifies the axial barrel coordinate relative to the breech face (xt is the location of the trunnion); yt gives the vertical displacement of the trunnion axis; \u03b8 specifies the rotation of the barrel about the trunnion axis; and Y gives the vertical component of the lateral barrel displacement (relative to the static barrel centerline) in a coordinate frame that is rotating and translating with the barrel" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003345_800161-Figure3-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003345_800161-Figure3-1.png", + "caption": "Fig. 3 - Articulated vehicle model", + "texts": [ + "2 m, there is no significant difference between the critical speeds associ ated with the second linear approximation and the quadratic expression. For smaller values of b, the second linear approximation provides 76 a conservative estimate of critical speed. On the other hand, the first linear approximation is entirely unsatisfactory, since it leads to the conclusion that there is no critical speed. ARTICULATED VEHICLE MODEL The equations governing perturbations of the constant speed straight line motion of the model of an articulated vehicle sketched in Fig. 3 can be written in the form r > + [C] i r }+ [K] < 0 (15) where the elements of the [M], [C] and [K] matrices are M(l,l) = -(ra 4- nvj,) M(l,2) = (h + e 4- d) - m ^ (s - 2e)/2 4- (b - d) M(l,3) - nyl 4- m T U (b - d) M(2,l) = M(l,2) M(2,2) = -m T (h 4- e 4- d ) 2 - m p u [ (s - e ) 2 + (~e)2]/2 - m ^ [2(h 4- e) 4- b 4- d] (b - d) .P - (I YAW + ill. + iH.\u201e + i.TD.,) YAW YAW YAW M(2,3) M(3,l) M(3,2) K(3,3) 0(1,1) C(l,2) C(l,3) C(2,l) -mT (h 4- e 4- d) d - m ^ (h 4- e 4- b 4- d) (b - d) - ) + L(l - yi -sin20) (178) From Fig. 3-45, we may deduce that ( n\\ 2 -) sin2 (9 L / Hence, x = R(l- cos(9) + L (l - y/l - (R/L)2 sin2 9)) (179) is the desired relation between x and 9" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000629_fst99-a98-Figure28-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000629_fst99-a98-Figure28-1.png", + "caption": "Fig. 28. COMPASS-D experiment No. 18524 showing the control response during a gradual increase to the elongation of the plasma: phase-plane diagram.", + "texts": [ + " The growth rate gp is measured at the time of the disruption, and the corresponding stability region Vmax is calculated. The growth rate is raised by increasing the plasma elongation, which is proportional to the ratio between the shaping current Ish and the plasma current Ip. In experiment No. 18524 the elongation is increased by ramping-up the shaping current Ish from 7000 to 8500 A, as shown in Fig. 27. The plasma current is held constant at Ip . 170 3103 A from 0.1 s to the end of the test. The growth rate gp rises from 1000 to 3000 s21 before the plasma disrupts due to the occurrence of a large ELM. In Fig. 28 the phase-plane diagram of the state trajectory is shown. The circles correspond to two different times also marked in Fig. 27. The predicted stability regions Vmax for the values of gp 51000 and 3000 s21 are shown. As the shaping current increases, the growth rate increases, and the maximum stability region moves inward toward the smaller Vmax . As long as the state trajectory is confined inside the predicted Vmax , the system is stable. At time t2 5 243 3 1023 s when gp 5 3000 s21, the ELM is sufficient to displace the plasma outside the region Vmax " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001148_ac951051v-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001148_ac951051v-Figure1-1.png", + "caption": "Figure 1. Schematic diagram of the optically transparent thin-layer electrochemical cell (OTTLE). The thin layer, windows, and cell walls have been exaggerated for clarity.", + "texts": [ + " A specific example of such a reaction is the electrochemically induced polymerization of a soluble monomer to form an insoluble electroactive polymer film, which deposits on the working electrode. In this case, it is useful to obtain spectral data not only as a function of time but also as a function of potential under application of a voltage ramp (i.e., cyclic voltammetric conditions). In an effort to further develop the capabilities of fast-scanning spectroelectrochemistry, we have designed an OTTLE cell, shown in Figure 1, that employs a small-diameter cylindrical thin layer that is open and accessible to bulk solution along the entire thinlayer circumference in a manner similar to those of earlier designs that attempted to minimize uncompensated resistance.2a-c The auxiliary electrode follows the entire edge of the thin layer, establishing a more symmetrical geometric relation between the working and counter electrodes. The cell has been evaluated under potential scan conditions. In nonaqueous solvent (acetonitrile) at scan rates as high as 25 mV/s, near-Gaussian current peaks with peak separations on the order of 165 mV or less are routinely obtainable from millimolar solutions of electroactive species" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000194_s0263574797000489-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000194_s0263574797000489-Figure1-1.png", + "caption": "Fig . 1 . Schematic diagram of the Delta robot", + "texts": [], + "surrounding_texts": [ + "For simplicity , the measurement device in the example is assumed to be capable of measuring all 6 components of the endpoint position expressed in the world frame . A set of n measurement points in the workspace is chosen and readings are taken from the measurement device when the robot is moved to these points . It is appropriate to consider which set of measurement points in the Fig . 2 . The fixed-leg Stewart Platform . robot\u2019s workspace guarantees a complete solution for the error parameters dI . This is a question of obser y ability . O ( M ) 5 4 L s 1 ? s 2 ? ? ? s L 4 n (16) where s 1 , s 2 , . . . , s L are the singular values of the identification Jacobian . Menq and Borm 1 1 present an observability index (16) based on the singular value decomposition of the identification Jacobian M (15) . It is normally advantageous to choose measurement points that maximise the observability index . Borm and Menq 1 2 use a numerical optimisation procedure to search for optimum measurement positions . For the example presented here , a similar numerical optimisation was performed using a steepest ascent method . It was observed that the measurement points generally migrate towards the edges of the workspace and toward singular configurations as the optimisation progresses . This is because the best increase in the Menq and Borm observability index is achieved by increasing the greatest singular value of the identification Jacobian . This is not a great problem with serial manipulators since there are no singularities of the type where mobility is gained however parallel robots often possess singularities of this type . When the robot is moved close to a singular configuration some terms of the direct Jacobian matrix become very large . This appears in the singular value decomposition as a very large greatest singular value . The other singular values are less prone to inflation . The near-singular points caused numerical range dif ficulties during the identification phase . Similar dif ficulties were encountered when using either the least singular value or the condition number as the observability index . A possible solution to this problem would be to set an arbitrary limit on the 2-norm of the direct Jacobian matrix of the manipulator at each measurement point . This is equivalent to creating an imaginary sphere around each singular point which the measurement points cannot move into during optimisation . The data for this example were generated numerically . A set of values was chosen for the dI error parameters . The object was to recover these values from the data using the identification technique . Once the measurement data has been obtained , the values of the error parameters dI must be determined . This is a model fitting exercise . The simplest approach to solving for dI is to use a linear least squares solution . Linear least squares estimation works satisfactorily if the identification Jacobian is well conditioned , the observability is relatively high , and the calibration model is suf ficiently linear . For situations where the condition number is poor , singular value decomposition 1 3 is often used . For improved performance , 1 4 it is possible to use a weighted least squares solution which weights the contribution of each data point to the result according to the inverse of its uncertainty . Non-linear least squares solutions are generally more computationally demanding but they often give a more accurate result . A practical technique is the LevenbergMarquardt algorithm . 1 3 The technique involves minimising a cost function (17) \u03c7 2 5 O ( x act i 2 G # ( x nom i , e \u0303 )) ? diag ( C ) 2 1 (17) where G ( x nom i , e \u0303 ) is the direct geometric solution and C is the covariance matrix of the experimental data . Matrix C may be known a-priori or it may be estimated from the experimental data . For this example , the simulated data were processed using MatLab . 1 5 The number of measurement points used was 50 . The results for noisy and noise free data heave been analysed by both linear and non-linear methods and presented in Table I ." + ] + }, + { + "image_filename": "designv11_65_0003236_a:1020985609601-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003236_a:1020985609601-Figure1-1.png", + "caption": "Fig. 1. Scheme of the flow cell. Holes for connection scews are not shown. Dotted lines denote the mutual orientation of the components when assembling. Polytetrafluoroethylene membrane is used as a sealing layer.", + "texts": [ + " The electrodes were kept in the dry state or in the working buffer solution at 4 \u00b0 C. Flow-injection measurements. A flow cell with a 30- \u00b5 L working chamber was made of polymethylmethacrylate. It consisted of two parts bolted by two M6 bolts. The upper part (0.7 mm in diameter) included channels for the reagent supply and removal and screwed joints for connecting elements. The cholinesterase sensor was fixed with a sealing layer in a rectangular-shaped hollow located in the channel junction in the lower part of the cell (Fig. 1). All the junctions were implemented with polytetrafluoroethylene tubes (i.d. 0.5 mm). The circulation of solutions and the dosing of reagents were provided with a BR-1 flow-analyzer unit (AO Khimavtomatika, Moscow). The unit was equipped with a peristaltic pump and a loop for injection of a substrate and inhibitor. The loop volume was 45 \u00b5 L; fill-up time, 30 s; and flow rate, 0.1 mL/min. The manifold for biosensor-response measurements is given in Fig. 2. An anodic-oxidation current of thiocholine measured at +560 mV against an Ag/AgCl electrode served as a measured signal (biosensor response)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002291_robot.1993.291985-Figure2-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002291_robot.1993.291985-Figure2-1.png", + "caption": "Figure 2: definition of the endeffector's frames", + "texts": [ + " The six voltages delivered by the force sensor form a vector V which, after multiplication by a calibration matrix C, gives the forces and moments vector H exerted on the sensor at a coordinate frame Z, located at the sensor's center, and expressed in &: '\"Hz, = CV For the force control that we want to implement, we have to measure the forces and moments exerted by the robot's tool on the environment and express them in the frame &, attached to this tool, that is & H h . This vector can be obtained from by: I: '\"Ax, is a rotation matrix which represents the orientation of I;, with respect to h. ' C P ~ is the vector which contains the coordinates of the origin of Z n in Z,. x represents the cross product operator. Figure 2 shows Z, and Z, (attached to the sensor and the tool, respectively) as they are defined in our application. Our UNIMATE controller uses the VAL 11 software as a robot programming language. In VAL, the position of a frame C with respect to a given reference frame Z, is described by the cartesian coordinates X, Y, Z of the origin of Z in Zr. The orientation of z is described by three angles denoted as 0, A, T. There exists a relationship between these angles and the Z-X-Z Euler angles a, p, y, as follows: 0 = a; A = p - 90\"; T = y + 90\" (1) Once Z is defined, it is possible to make the robot's tool frame Z n coincide with Z" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000561_41.793351-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000561_41.793351-Figure1-1.png", + "caption": "Fig. 1. Two-link flexible-joint manipulator.", + "texts": [ + " Remark 2: The control gains and in the fast controller (32) are designed such that the poles of the fast subsystems with their nominal values are placed far in the left hand of the complex plane. This may always be ensured for all , such that the uncertainty in (34) does not affect the stability. In other words, one can, in principle, ensure this by imposing a bound on the variations in , , and and by designing and taking this constraint into consideration. To illustrate the application of the proposed adaptive control scheme, consider a two-link manipulator with flexible joints, as shown in Fig. 1. The elasticity in the joints is modeled as a torsional spring of stiffness . The equations of motion for such a manipulator can be written [15], [21] as (44) (45) (46) (47) where are the entries of the inertia matrix (a) (b) and are the mass and moment of inertia of the th link, respectively, , are the length and the distance from the joint axis to the center of mass of the th link, respectively, and , , , , where is the gravitational constant. The manipulator consists of actuators of inertia connected through gear boxes of ratio . Here, we consider a system with the nominal parameters given by , , , , , and . Equations (44)\u2013(47) may be represented in the standard singular perturbation form as (see [15] for details) (48) (49) where with (43) (a) (b) (c) (d) Fig. 6. Closed-loop response of the full-order system with parameter variations (solid line: desired; dashed line: actual). (a) Position of link 1. (b) Position of link 2. (c) Velocity of link 1. (d) Velocity of link 2. The manipulator of Fig. 1 was assumed to be initially at rest with and . The initial conditions of the fast variables were set to arbitrary nonzero values to ensure that the fast dynamics will be excited. The initial values of the feedback gains for the slow subsystems were chosen to be and for . For the fast subsystems, we chose and for . The desired reference trajectories are given by (50) (a) (b) (c) (d) Fig. 8. Closed-loop response of the full-order system with s(t) = 0 (solid line: desired; dashed line: actual). (a) Position of link 1" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001273_rob.4620080205-Figure2-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001273_rob.4620080205-Figure2-1.png", + "caption": "Figure 2. Illustration of the normal driving-axis coordinate system.", + "texts": [ + " The resulting formulation is more efficient than the original one based on the Denavit-Hartenberg notation, and made the implementation on an INTEL 8086/8087 microprocessor (5MHz clock) under 17ms for the Stanford manipulator. It is also showni3 that the reformulated Walker and Orin's composite body method based on the normal driving-axis coordinate system can save 45 multiplications and 28 additions for a general open chain in comparison with the one based on the Denavit-Hartenberg notation. This experience is adopted again to establish an algorithm for the formulations derived above. In the normal driving-axis coordinate system (Fig. 2), the z-axis of a bodyfixed frame is the driving axis of the corresponding link, i.e., uji) = [O 0 l l T . sin p i , di cos pi ] T , where bi , di , pi and 8; are the parameters of the coordinate system and are shown in Figure 2. Note, d; = di + qi, Oi = e l , ifjoint i is translational; otherwise di = d / , Oi = 0; + qi; i.e., d ; and 8 : are the null-position values of di and 8;, respectively. And the distance from the origin of frame Ei-l to frame Ei is ;-:di-l) = [bi, - di 208 Journal of Robotic Systems-1991 Since it is recommended to assign the base frame Eo coincident with frame El in the null-position configurati~n,\u2019~ the distance between the origins of Eo and El is then zero. Thus, The terms, which are constant and can be calculated in advance, are listed in the part \u201cinitialization\u201d of the algorithm shown in Figure 3" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002201_iemdc.2001.939386-Figure4-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002201_iemdc.2001.939386-Figure4-1.png", + "caption": "Fig. 4. Modification to Stator Bars and Gap Support", + "texts": [ + " The original chamfer was replaced by an inlet radius in the circumferential direction and a smaller radius in the radial direction. The shape was interpolated between the radial and the circumferential radius by a smooth transition. Standard Motor Modification Fig. 3. Modification to Rotor Bars C. Modification to the Stator Bars At the downstream face of the stator bars a wedge was added. While the flow undergoes a sudden expansion in the original configuration, this wedge allows a smooth diffuserlike deceleration of the coolant. The modification is shown in Fig. 4. Bars D. Modification to the Gap Supports in the Stator The shape of the rotor bar was modified as shown in Fig. 3. Instead of the sharp edges facing towards the center of rotation a smooth transition was applied. CFD calculations had shown a significant reduction of pressure loss due to this modification. The losses related to the separation and flow contraction between the rotor bars and the gap supports could be reduced significantly. The radius of the inner end of the gap supports was reduced close to the stator inner diameter. The gap supports were optimized by CFD calculations to provide the most uniform distribution of flow to the two passages windward and leeward of the stator bar while keeping the losses at a minimum. The new configuration is shown in Fig. 4. IV. TEST-RIG DESCRIPTION In order to investigate the effects of the suggested improvements in detail, a rotating test rig was built. The entire facility is shown in Fig. 6, a closeup of the test section is shown in Fig. 5. This test rig consisted of a periodic element of the motor modeled at a scale of 1:l. Two rotor cooling gaps and one stator cooling gap were modeled. The distribution of mass flow in the stator-rotor gap was controlled by two labyrinth seals on the Drive End (DE) and Non-Drive End (NDE) of the stator-rotor gap" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001285_1.2833938-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001285_1.2833938-Figure1-1.png", + "caption": "Fig. 1 Hydrostatic journal bearing with self-controlled restrictors em ploying a floating disk", + "texts": [ + " Furthermore, Kazimierski (1992) modified the bearing structure to improve the easiness of manufacturing the bearings. Yoshimoto (1987) studied the static characteristics of aerostatic journal bearings which had an inner bush supported by O-ring to improve static bearing stiffness. Mizumoto et al. (1987) proposed a hydrostatically-controlled restrictor using a floating ring. Yoshimoto (1990) has proposed a hydrostatic journal bearing with a selfcontrolled restrictor employing a floating disk, as shown in Fig. 1. It was subsequently demonstrated that the proposed bearing could achieve a very high static bearing stiffness (nearly infinite Contributed by the Tribology Division for publication in the JOURNAL OF TRIBOLOGY . Manuscript received by the Tribology Division June 30, 1996; revised inanuscript received March 18, 1998. Associate Technical Editor: M. J. Braun. Stiffness) over a wide range of imposed load independent of supply pressure and viscosity of operating fluid. However, as mentioned by Mori et al", + "org/about-asme/terms-of-use loating Disk ^W\\r Fixed Restrictor (Capillary) -iMv- Self-Controlled Restrictor Fig. 2 Typical directions of tlie step load ploying a floating disk are investigated theoretically and experi mentally. Then, the influences of various design parameters such as supply pressure, viscosity, the magnitude of the step load, and the imposed static load on the step response character istics are discussed. Furthermore, theoretical results are com pared with experimental results in order to confirm their vaUdity. Figure 1 shows the geometrical configuration of the proposed journal bearing. The proposed bearing is circumferentially di vided into four pads. A self-controlled restrictor is installed in pads 1 and 2, respectively. The pocket in the pads 3 and 4 has a capillary restrictor, as shown in Fig. 1. Therefore, the proposed bearing is asymmetric with respect to the bearing center and it can be considered that the bearing characteristics depend on the load directions. Figure 2 shows three typical load directions, y\\, y2, and y^. In the yi direction, the load is imposed between pads with a self-controlled restrictors. In the ^2 direction, the load is imposed just on a pad with a self-controlled restrictor, and is imposed between a pad with a self-controlled restrictor and one with a capillary restrictor in the y^ direction", + " 3 Theoretical Analysis and Results In order to theoretically obtain the step response characteris tics of the proposed bearing, the coordinate system as shown N o m e n c l a t u r e Co = average bearing clearance c = outer gap of a floating disk du d2 = diameter of capillary 1 and 2 d^ = diameter of a fixed restrictor (capillary) D = diameter of bearing e^, By = displacements of the shaft in the X and y directions Ae,, Aey = displacements of the shaft in the X and y directions from the equilibrium position = flow controlling gap (see Fig. 1 ) (7= 1,2) = Co( 1 + E;,- sin 6 + By COS 9): bearing clearance Co, H2 = h2/co: dimensionless bearing clearance U = 1, 2) lengths of capillary 1 and 2 length of a fixed restrictor (capillary) = bearing width and the width of pocket 2 \u2022 mass of a floating disk (J = h 2) / ! l , //\u201e = /, L,L, /M2 = mass of a shaft \u00ab], n2 = the number of capillary 1 and 2 Pij, P2i = pressures in flow controlling gap and bearing clearance (/ = 1 - 4 ) , U = 1, 2) Pa = ambient pressure Ps = supply pressure Pmi = pressure in the pocket 2 of the pad (/ = 1 ~ 4) Pij = pressure in the pocket 1 of a self-controlled restrictor (j = 1,2) P = (P - Pa)l{Ps - Pa) = dimensionless pressure r = radial coordinate in a selfcontrolled restrictor rb, r^, Trf = radii (see Fig. 1) W = w/{LD{p,, \u2014 Pa)} = dimen- sionless load capacity Ws = wj {LD(ps \u2014 Pa)} = dimen- sionless step load yi, y2, y3 = three typical directions of an imposed load (see Fig. 2) z = axial coordinate Za, Zb = values of axial coordinate at the edge of pocket 2 e = eccentricity 8;, = e^/co. By = By/Co = ecccntrlcities in the x and y directions &a,()x,92 = circumferential angles per taining to pocket 2 da, &b- values of circumferential co ordinate at the edges of pocket 2 /u = viscosity of lubricant T = (ps ~ Pa)t/1^ = dimen- sionless time (p = attitude angle Subscripts i = 1, 2, 3, 4 = parameters pertaining to the pad number j = 1,2 = parameters pertaining to the number of selfcontrolled restrictors 316 / Vol" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003702_gt2004-53860-Figure7-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003702_gt2004-53860-Figure7-1.png", + "caption": "Fig. 7 Static load deflection mechanism for full damper", + "texts": [ + " Parameters such as foil thickness, bump height, pitch, length and material (including interface friction coefficient), were varied in order to achieve the desired stiffness and damping values. Based on the desired stiffness value and maximum expected motion, the compliant bumps were made of 0.152 mm thick Inconel foil. Figure 6 shows an unwrapped section of the full compliant foil journal bearing damper sub-assembly. Static Load Deflection Mechanism In order to conduct a static load deflection test on the full damper once installed in the weldment, a mechanism as shown in Fig. 7 was designed and fabricated. This mechanism used two precision coil springs and a bar to apply a known load to the inner ring/disk assembly. The displacement of the bump foil assembly, due to the applied external load, was measured using two high precision mechanical indicators (precision as high as 50pm). Additionally, in various tests, an eddy current displacement sensor was used to measure inner ring motion to verify the readings taken with the dial indicators. Previous testing of a smaller 150 mm diameter compliant foil damper assembly was completed using a computer controlled and automated load deflection tester 4 Copyright \u00a9 2004 by ASME l=/data/conferences/gt2004/71222/ on 07/07/2017 Terms of Use: http://www" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003848_2005-01-3779-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003848_2005-01-3779-Figure1-1.png", + "caption": "Figure 1. A typical rotary lip seal.", + "texts": [ + " The following specific results are included in this paper: Seal friction torque testing system development Drive axle pinion seal and wheel seal friction baseline data Effect of speed and temperature on pinion seal friction Effect of coatings on pinion seal friction Pinion seal friction torque calculation and comparison to test data Seal friction torque in relation to bearing friction torque and overall axle efficiency 2005-01-3779 Seal Friction Effect on Drive Axle Efficiency Hong Lin, Douglas C. Burke, Robert R. Binoniemi, Leo Wenstrup and Thomas Woodard Dana Corporation Rotary lip seals are used in drive axles both to retain oil and to exclude contaminants from entering. The primary performance requirement for a rotary lip seal is sealing, i.e. preventing oil leakage. Other major performance parameters include seal life and friction. Figure 1 shows a typical rotary lip seal. Rotary lip seals have some distinctive application advantages: low initial cost, easy installation, small space requirement, and effective sealing. Main limitations may include: life, leakage, and the need for oil lubrication. Similar to any other tribology systems, the lip seal-shaft system is quite complex even though a lip seal is a low cost component. Lip seal performance, namely sealing, life and friction, depends on many factors, such as seal lip design, seal material and finish, shaft material and finish, seal/shaft geometry and alignment, seal installation, and operating conditions (temperature, speed, load, and lubrication)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003278_ias.1989.96735-Figure2-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003278_ias.1989.96735-Figure2-1.png", + "caption": "Fig. 2. Simplified equivalent circuit of a n I P M mo-", + "texts": [ + "OO 0 1989 IEEE Lcc = Laao + La1 + L92cos(26 - 120\") and the stator-to-stator mutual inductances are, -1 2 Lab = Lba = + Lg2COS(26 - 120\") -1 2 Lbc = Lcb = + LgzCOS(28) -1 Lac = Lca = Y L a a o + Lg2~0~(26 + 120') where, 6 = electrical rotor angle Laao = component of the self inductance due to the space fundamental air-gap flux. L,I = additional component due to the armature leakage flux. Lgz = component of the self inductance due to rotor position dependent flux. The direct and the quadrature a x i s inductances are, If the switching frequency is high ( > l o kHz), then during one switching period the variation of inductance with the rotor position can be neglected. With this assumption the instantaneous voltage equation for the phase 'a' of the IPM motor can be obtained from the equivalent circuit of Fig. 2. (9) di, dt Va = R,i , + Laa- + E a where, V, = terminal voltage of phase a i, = current through phase a Laa = synchronous inductance of phase a E, = the back-EMF of phase a R , = the resistance of phase a = Laa - Lab Similar equation can be easily written for the phases b and c. Inductance Calculation fiom the Hysteresis Current Controller The phase inductance of an IPM motor can be calculated analytically from the instantaneous voltage and current information. Fig. 3 shows the current waveform obtained when a hysteresis current regulator is used", + "- (V - E - R a +(I,(t)ltl - A 1 h ) e - W (13) Direct Method of Inductance Calculation (Va - Ea) (1 - e - - ) ir(t)(tl+At + AIh = ____ Instead of calculating the phase inductance indirectly, the phase inductance can be evaluated directly in the following manner. If the switching frequency is high enough (> 10 kHz), then the variation of inductance with the rotor position can be neglected. The phase voltage \u2018V,\u2019 in the equivalent circuit of Ra (I4) +(i,(t)ltl - A 1 h ) e - W The equation (14) can be simplified to express Laa as, -RaAt (15) the Fig. 2 can be written as, rn( $2) di, L., = Va = Raia + Laa- dt + E a where, Analogous equations can be obtained for the trajectory b-c of the current wave form when the current is decreasing. The equations are, The phase inductance La, can now be expressed as, V, - R,i, - Ea La, = &- d t where, where, and The equation (21) is easier to evaluate than equations (15) and (17). Moreover, equation (21) gives a single expression for the inductance calculation instead of two relations of the hysteresis controller" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002735_amp-120025079-Figure5-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002735_amp-120025079-Figure5-1.png", + "caption": "Figure 5. (a) Cage used for applying Re coatings on 18 graphite cores, and (b) polished cores with surface finish tl > max(tol, to2), and xl , 0 < x~ -< t~, such that x ( tl; tol) = x ( tl; to~) = Xl; JOTA: VOL. 51, NO. 1, OCTOBER 1986 153 x)-plane, emanating from the distinct points (to~, too and (to2 , /02) on the UP of the target set [namely, the segment connecting the origin to the point (4~, 4~ )], meet at the point (q , x~). Thus, nonuniqueness implies the existence in the (t, x)-plane of a triangular domain formed by a pair of intersecting trajectories and based on the segment [ to~, to2] of the half line {(t, x) [ x = t, t - 0 } , namely, the triangle to~, to2, (q , x~) in Fig. 4. Hence, either (a) or (b) below hold: t (a) there exists a sequence { ok}k=3~ [to~, to2], such that the triangles formed by the trajectories in the (t, x)-plane emanating from the points (tok, tok) and (tOk+l, tOk+O, which meet at (ilk, Xlk) and are based on the segments [tog, tOk+l], k = 1, 2 , . . . , are such that lim (qk, Xlk) C {(t, X)[ t = X, t --> 0}; (97) k~co (b) we have the situation in which there is a triangle To~, ~2, (71, ~ ) , where to1-< t-o~ < to -< to2, such that all the trajectories emanating from the points (to, to), where toe Iron, to2], meet at the vertex point (tl , x~); see Fig. 4, where tl must now be replaced by ?~, etc; then [see Eq. (94)], x(t'l; to)= 21, for all to C [to,, to2]. (98) Now, the optimal flow field is locally well-behaved near the UP of the target set. Indeed, it is easy to show that, explicitly, (d~*ldr) l ,=r=-a[(k-1) / r<][s in g / ( 1 - oz cos ~)], 0 - < ~ < 6 , and therefore no two adjacent flow lines would ever intersect in the neighborhood of the target, which is contrary to assertion (97). Hence, case (a) cannot arise. Consider now the function of to, ,/tAr, : [0, q$] ~ [0, tl], J/,/~,(to) --a x(T,; to), where, as before, x(" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000095_0957-4158(94)e0025-l-Figure4-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000095_0957-4158(94)e0025-l-Figure4-1.png", + "caption": "Fig. 4. Mechatronic tension compensator.", + "texts": [ + " Firstly, because the helix angle at which the yarn is wound varies in a complex manner, partly because of the practical limitations of manufacturing the traverse cam which distributes the yarn along the cone, and secondly because the compensation required changes as the winding operation progresses and the package increases in size (Fig. 3). These problems have not prevented purely mechanical solutions from being engineered but they are complex and costly and generally contain sliding contact surfaces (e.g. cams) which need to be carefully shielded from the fibres and dust of the spinning mill. The mechanics of a mechatronic positively driven tension compensator can be much simpler. Figure 4 shows such a device [9]. The simple two-bollard principle of the passive spring compensator is retained, but the bollard disc is now positively driven by a small stepping motor, the motion of which is controlled by a single-chip microprocessor of the 8051 family. An open-loop control strategy is dictated currently in this case by the non-availability of any cheap and reliable method of sensing tension in the running yarn. The control system is illustrated in Fig. 5. The motion of the disc is synchronised with the oscillation of the yarn traverse-guide using an optical 102 T " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002998_esej:20020106-Figure5-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002998_esej:20020106-Figure5-1.png", + "caption": "Fig. 5 Feedback Instruments' Project Kit 40-100", + "texts": [ + " The recruitment of a new TCS Associate revived the TCS Programme and a sales and marketing meeting was held to discuss the hture of the Mechamnics Project Kit. The kit was marketed very much as it was. It was felt that the complexity of out-of-circuit programming and W erasure of microcontroller devices was too onerous for certain market segments. A simpler device was thus selected (PIC16F877) with flash memory and in-circuit programmability, connections being made by screw terminals. The communications module was also omitted in this version (Fig. 5). The &rst time the module was provided to the MEng students, the results were a disaster. The students could ENGINEERING SCIENCE AND EDUCATION JOURNAL FEBRUmY 2002 not see the point in the lectures about how microcontrollers work or in what they were trying to do in the laboratory. They did not engage sdciently in the project and, despite regular exhortation, underestimated the amount of work required to deliver a working vehicle. Student feedback clearly indicated their dissatisfaction with the module and the lecturing staff were also disappointed with the outcome. The developed teaching material was thus scrapped and the module delivery reviewed. The current delivery of the mechatronics module at Loughborough has run successfdly on two MEng cohorts of which the largest was 32 students. The vehicle has been developed beyond the kit shown in Fig. 5 and is now in its 9th generation. Students are hooked by a module delivery that ensures they have a moving vehicle by the end of day one. Although all student teams are provided with the same basic kit, there is s t i l l room for creativity and their designs are varied (Fig. 6). Students will fiequently use the workshops to construct their own chassis or sensing systems and a few add body shells. The teaching method has changed dramatically. The students now get only one lecture to set the scene and define deliverables" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000406_105971239900700201-Figure9-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000406_105971239900700201-Figure9-1.png", + "caption": "Figure 9. Schematic representation of the one-legged hopping robot.", + "texts": [ + " Small random values are added to the coupling weights during learning to avoid a stagnation in learning by the eventually obtained value 4x~~~ z 0 - Figure 7 shows the timeaveraged height of the robot X-0 , which indicates that ~y - 0.95 [m] was obtained by the learning. The obtained value was almost same as the maximum value obtained in the previous simulation (Figure 5). This value would be the maximum height allowed by this physical system. hopping height xd = 0.9 [m] was achieved. at UNIV OF CALIFORNIA SANTA CRUZ on April 2, 2015adb.sagepub.comDownloaded from 145 3.3 Control experiment of a hopping robot The proposed learning rule was applied to the control of an actual one-dimensional hopping robot (Figure 8). Figure 9 shows an overview of the robot. The robot moves along two vertically situated poles. A DC-motor on the trunk makes a force on the table through arms which consist of a crank system. Because a spring is fixed under the table, the robot can continue hopping if the arms push the table at an appropriate time in a hopping cycle. The angle of the arm 0 is measured by a potentio-meter, and the acceleration of the table a is obtained by an acceleration sensor. Because the motor speed ~ is approximately proportional to the current I applied to the motor in steady state, i" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002629_iros.2001.973411-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002629_iros.2001.973411-Figure1-1.png", + "caption": "Figure 1: Dexterous Shoot composed of A p p d and Kick", + "texts": [ + " For example, the element motions may switch from one to another very slowly in order to achieve 8* steadily. Such a slow sequential composite motion cannot be desirable. To speed up the behavior the criterion should be modified, for example, to J 7 ( 4 = 118O0(a) - 8*112 + PT,, P > 0 (4) where T7 is the necessary time to complete the composite motion. With this criterion next motion begins without waiting for the completion of the previous motion. This is effective not only for speed-up but also for smooth motion switching. The resultant behavior is smooth and speedy, i.e., dexterous. Figure 1 shows a dexterous ball shooting behavior (Shoot) considered in this work. Shoot is a sequential composition of approaching (-) and kicking (Kick) (Figure 1 upper left and right, respectively). If Kick begins &er Appxxh completes the precise positionning, the result is awkward and inefficient(Figure 1 lower left). Kick should smoothly switch to h to achieve a dexterous Shoot (Figure 1 lower right). A motion switching is practically a motion merging. Suppose that motion i smoothly switches to motion i+ l at time t o , and let qi(ti) be the target posture in motion i at time ti. And let the motion switching be done by shiRing the target posture as q(t) = fsw(t)qi(ti) + (1 - fSw(t))qi+l(ti+l) where t = ti+' = ti-to , &(t) is the motion switching function defined, for example, as By choosing appropreate t* and T*, a smooth motion switching is obtained as shown in Figure 2. 3 Knowledge array network The solution (7 ,a) of problem (3), i" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001492_iros.1997.649095-Figure6-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001492_iros.1997.649095-Figure6-1.png", + "caption": "Fig. 6 foot end trajectory (divided to six parts)", + "texts": [ + " In section 4, the robot is consisted with four wits and in section 5 the robot is consisted with five units. Fig. 2 shows a unit of the robot. A unit has one computer for control, four motors to move the two legs and interface circuits for communication and control. Two methods of the leg control me used in the experiments mentioned in the following sections. In experiment at section 4, the one cycle motion divided to four part. h d in experiment at section 5, the one cycle motion divided to six parts to stabilize the robot posture. In the experiment in section 4, the marks on the trajectory in Fig. 6 (a) show the points o f the parts division. The both sides legs of a unit positioned with two 404 parts diffemce on the trajectory by each other. That means the both side legs moved on reversed phase. In this experiment, the walking robot is consists with four units. During the walking, the first unit and 4th unit move their legs on same phase. second unit and 3rd unit move their legs on same phase too. These two groups of units move on reversed phase by each other. The reason why the robot used these gait control is that when the second and third units move on same phase motion, the yaw torque to turn each unit are canceled by each other. Then the robot keep the posture in the walking. In experiment at section 5 , the one cycle motion divided to six part. Fig. 6 (b) shows the foot end trajectory against the unit like Fig. 6 (a). The marks on the trajectory show the points of the parts division. The both side legs of a unit have a three parts difference on the trajectory. Though each leg moves on reversed phase of each other. In this experiment, the robot is consisted with five units. The reason to use these method that when the unit volume is uneven number, the yaw torque of each units is not canceled. If the leg motion like the four units robot, the five units robot does not keep the posture. With this leg control, a leg is contacted on the grand in the five part of the whole six part" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001428_s0167-8922(00)80155-8-Figure3-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001428_s0167-8922(00)80155-8-Figure3-1.png", + "caption": "Figure 3 Grid locations for tempera ture m e a s u r e m e n t", + "texts": [ + " The microscope is focussed onto the steel ball surface within the contact, and uses an lnSb detector to sample the level of IR emission from the focus area. The microscope is able to focus on a spot approximately 11 ~tm in diameter. Since the Hertzian diameter is typically 300 lttm, this means that up to 27 non-overlapping emission and thus temperature readings are possible across the contact. In practice, the microscope is made to fully traverse the contact, to determine IR emission over a grid of locations from the inlet to the exit region of the contact, as shown in figure 3. A transmission filter was used to remove any IR emission from the hydrocarbon oil film present, as shown in figure 4. The IR emission from the sapphire can be neglected, which means that all measured emission originates from the steel surface. This emission was directly related to surface temperature by means of a calibration in which IR emission was monitored through a sapphire window from a polished steel surface at a series of temperatures. Figure 5 shows a typical steel ball temperature rise map from a contact lubricated with mineral oil at a bulk lubricant temperature of 60\u00b0C" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001696_s0263-8223(01)00168-4-Figure2-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001696_s0263-8223(01)00168-4-Figure2-1.png", + "caption": "Fig. 2. (a) Spring subject to diametrical compressive line load F . (b) Reduced loading system of a semi-spring subject to diametrical compressive line load F .", + "texts": [ + " Castigiliano\u2019s theorem and strain energy expression were applied together to develop the stiffness equations of the spring. The strain energy U per unit width of the spring is given by Tse and Lai [13] U \u00bc Z n 1 2 N M ( )T A0 B0 C0 D0 \" #T N M ( ) dn; \u00f01\u00de where n is the surface area of the mid-surface. \u00bdA0 is the in-plane compliance matrix, \u00bdB0 and \u00bdC0 are the coupling compliance matrices and \u00bdD0 is the flexural compliance matrix. \u00bdM and \u00bdN are the moment resultants and stress resultants, respectively. The composite spring is subject to unidirectional compressive line load \u2018\u2018F \u2019\u2019 in y-direction as shown in Fig. 2(a). According to Castigliano\u2019s theorem, in finding the deflection of curved beams and similar structures, only strain energy due to bending need normally be ta- ken into consideration. The strain energy expression for the complete spring derived from the semi-spring model as shown in Fig. 2(b) is U \u00bc 2L\u00f0UBE \u00fe UEF \u00fe UFC\u00de \u00bc 2L 1 2 Z n D11D66 D2 16 jDj M2 dn BE \u00fe 1 2 Z n D11D66 D2 16 jDj M2 dn EF \u00fe 1 2 Z n D11D66 D2 16 jDj M2 dn FC ; \u00f02\u00de where UBE;UEF and UFC are the strain energy in the three portions BE, EF and FC, respectively. jDj \u00bc D11\u00f0D22D66 D2 26\u00de D12\u00f0D21D66 D16D26\u00de \u00fe D16\u00f0D21D26 D16D22\u00de; where Dij are elements of the flexural stiffness matrix \u00bdD ; M is the bending moment per unit width, L is the width of the spring. By substituting the moment resultants into Eq. (2), we have U \u00bc L D11D66 D2 16 jDj Z w=2 0 1 L MB ( F 2 s 2 Lds \u00fe Z p a a 1 L MB F 2 R sin h 2 LRdh \u00fe Z w=2 0 1 L MB F 2 s 2 Lds ) ; \u00f03\u00de where s is the distance from the middle of the flat contact surface in x-direction, w the length of the flat contact surface, R the mean radius of the spring, F the applied unidirectional line load, MB the bending moment per unit width at the middle of the flat contact surface and h is the angular coordinates of the spring" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000021_922404-Figure7-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000021_922404-Figure7-1.png", + "caption": "Figure 7 . A die configuration", + "texts": [ + " In this case, the machine and die must be moved alternately to generate a path with no collisions. At present, the programmer determines the free path manually. ~ecausethe shape, size and location of the structural elements differ from one panel to another, there is a multitude of escape trajectories. Consequently, a general path planning method, able to find all these trajectories, is required. Implicitly, the die is oriented at 90\" to the direction of travel but, for difficult positions, this can be changed (See figure 7). A system configuration is therefore defined by the value of the five machine axis setpoints plus the die position and orientation. Then a collisionfree path is generated. DESCRIPTION OF THE MACHINE KINEMATICS - The kinematics of the automatic riveting machine are complex. The machine has 5 degrees of freedom which is necessary to position the machine datum perpendicular to the position of the point to be riveted. The machine includes a C-frame capable of moving along the x-axis. The tool head is mounted at the top of this frame and a die at the bottom" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002398_cdc.1995.480218-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002398_cdc.1995.480218-Figure1-1.png", + "caption": "Figure 1: A two-link robot arm with joint angles 6 = (61,82), joint torques r = (TI] r ~ ) ] end-effector position 2, desired end-effector position xd, link lengths 11 and 1 2 , and link masses ml and m2.", + "texts": [ + " 0 Proof: This is a straightforward application of the implicit tracking theorem, Theorem 2 of [?I. See also [6]. 0 Remark 4.2 Equations (21) are the equations of motion for the manipulator. Equations (24) provide exponentially convergent estimates f ( t ) and 6( t ) of r,(t) and 6 , ( t ) , Equations (22) and (23) determine A the input r as a function of 6 , f', and 6 . 5 . A Two-Link Example In this section we work through an example of the application of Proposition 4.1 to the control of a simple model of a two-link robotic arm diagrammed in Figure 1. The links of the robot arm are assumed rigid and of length 11 and 1 2 . The masses of each link are assumed, for simplicity, to be point masses ml and m2 located at the distal ends of link 1 and link 2 respectively. The desired position of the end-effector at time t is z d ( t ) . The actual position is z ( t ) . We wish to make the end-effector (end of the second link) track a prescribed trajectory zd(t ) in the Euclidean plane. The joint-space of the arm is parameterized by 6 E T2 where T2 is the two-torus" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002529_00037-1-Figure3-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002529_00037-1-Figure3-1.png", + "caption": "Figure 3 Dimensioned sketch of an assembled Wyoming combined-loading compression (CLC) test fixture (after Adams and Welsh, 1997).", + "texts": [ + " Thus, various attempts have been made over the years to apply a combination of end and shear loading (Ewins, 1971; Purslow and Collings, 1972; Port, 1982; Hsiao et al., 1995). While the concept was sound and these methods were a technical success, the test fixtures and/or procedures used were not sufficiently attractive to become used extensively. More recently, Adams and Welsh (1997) developed what they term the Wyoming Combined Loading Compression (CLC) test method, and associated fixture shown in Figure 3. The test fixture is relatively simple and easy to use. A straight-sided [90/0] crossply laminate can be tested without tabs since flame-sprayed gripping surfaces are used. The clamping forces are attained by tightening four bolts in each fixture half. By controlling the torque applied to the bolts, the ratio of end loading to shear loading can be controlled. Since only relatively low bolt torques are required (2.2\u00b13.4N-m), the stress concentrations induced by the clamping forces are low, as opposed to shear-loading methods", + " This fixture performs very well when used with low strength materials. It was not intended for use with high strength materials, but could have been if specimen tabs were added. Without tabs the specimen would have end-crushed. In 1996, some 15 years later, recognizing the value both of using an untabbed specimen and combined loading, researchers at the University of Wyoming (Adams and Welsh, 1997) added flame-sprayed gripping surfaces and a few other minor refinements to the ELSS fixture, as shown in Figure 3. Increased bolt torques were used so that a combined loading rather than pure end-loading was achieved. The straight-sided untabbed specimen was retained. The intention was to test [90/0] laminates in compression successfully, and then to \u00aabackout\u00ba the unidirectional ply axial compressive strength, as discussed previously in Section 5.06.6.1. The fixture has been renamed the Wyoming Combined Loading Compression (CLC) test fixture and is receiving considerable attention in the late 1990s (Wegner and Adams, 1998; Adams and Welsh, 1997)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002097_0022-4898(85)90047-3-Figure2-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002097_0022-4898(85)90047-3-Figure2-1.png", + "caption": "FIG. 2. Scheme of the stand with the impact pendulum.", + "texts": [ + " The cutting tools - - wedge shaped (apex angle 35 \u00b0) knives made of aluminium alloy, height 80 mm with three width variants 60, 80 and 100 mm used singly and in combinations to yield cutting depths of 80, 160 and 240 ram. The scheme of the hydraulic stand is shown in Fig. 1. The stand, a frame-type structure, comprises a massive steel shaft (reciprocating in roller guides), whose forepart carries a suitable adapter through which the cutting tool with its transducers is mounted; the back part is connected to the piston rod of a dismountable hydraulic drive, which is replaced by the pendulum in the dynamic tests as shown in Fig. 2. Hydraulic drive ~ F ra m e-t Yite ~i rUhatf~ r e ~ Guide roller Adapter ~ - I ] ~ r F\u00b0rce 7\u00b0,\u00b0:::, Bin Soil El(;. 1. Scheme of the hydraulic stand. The bin is mounted on the forepart of the frame and filled with soil in layers of 0.2 m, compacted to natural density (determined with a standard penetrometer [23]). The physical properties and sieve analysis of the soil are listed in Tables I and 2. The transducer device consists of a pair of supports (upper and lower) for the knife or set of knives; the sum of the indications represents the resultant frontal resistance force, and their difference yields the position of the latter along the vertical" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003034_cdc.1994.411508-Figure3-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003034_cdc.1994.411508-Figure3-1.png", + "caption": "Fig. 3 - Position and orientation errors in a curved path", + "texts": [ + " In this sense, v , = v , + m x m , v , = v , + a x m (2) The coordinates of point S can be determined from geometrical relations, as a tan 6, + b tan 6, CM = , M S = a + b tan 6, - tan 6, tan s, - tan s, (3) Equation (3) implies that the radius for a circular path is minimum when S, and S, assume their maximum absolute values and have different signs. Moreover, it may be seen by inspection that the steering wheel with larger angle (absolute value) is dominant. That is, when abs($) > abs(&), if the front wheel is to the right the vehicle turns right no matter whether the rear wheel is to the right or left. The effect of the other steering wheel is to change the radius and the centre of rotation. 2452 The angular speed w may be defined in terms of the steering angles of the wheels in the form of (4) V, w =-(tan 6, - tan 6,) a + b Figure 3 shows the position and orientation errors after a small motion of the vehicle for the period At. The effect of the curvature change can be sitive or negative. The lateral displacement depends on e forward displacement and the orientation change. In this sense, the lateral speed of point M in the XPY coordinates is yM = w, = V, A0 (5) where V, is the forward speed. For point C, the rate of lateral displacement, therefore, is given by We are concerned about the position error and its variation in the course of motion. It follows from equation (6) and figure 3 that td = V, E, + (MC) w f k (7) where 1 stands for the effect of the path curvature and it is zero in the case of a straight path. 3. Structure of a Controller Noting that the angular speed is the rate of change of onentation and denoting by w part of the orientation error due to path curvature, the equations of motion in the state space form can be derived from eauations (3). and (8) as follows:The open loop system has two poles at zero and is unstable, but it can be stabilized b an appro riate feedback law to determine the two inputs, tan 6: and tan $" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002560_icsmc.1995.538488-Figure4-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002560_icsmc.1995.538488-Figure4-1.png", + "caption": "Figure 4: Configuration of the wrist singularity for PUMA manipulator.", + "texts": [ + " 3, it h a p pens when the Wrist locates at the yl -a1 plane. Then, one of the axes in z1, or z2 or y3 can be chosen as the singular direction. Here, the d:rection in y3 is chosen. Therefore, when the interior singularity occurs, J11 viewed in coordinate 3 becomes From (ZO), the linear singular direction of the forearm interior singularity is clearly in parallel with y3 axis. 2.2.2 Wrist Singularity The wrist singularity can be identified by checking the determinant of the matrix OJ22 in (14) as (21) Referring to Fig. 4, it can be seen that the wrist singularity happens when the z3 and a5 axes are collinear. In this configuration, the angular velocity about the x 5 axis (i.e., 5 f l ~ ) is unachievable. Thus, 'J22 will have the structure A det('J22) = 7u = -S5 = 0. 0 0 0 5 ~ 2 2 = [ y y ] Consequently, the singular direction of the wrist singularity is about the x5 direction. After analyzing all of the unescapable singularities of nonredundant manipulators, considerations and resolution methods for the Inverse Kinematics singularity problem will be presented in the next section" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003597_trib2004-64334-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003597_trib2004-64334-Figure1-1.png", + "caption": "Fig. 1. A schematic diagram of linear microball bearings [2].", + "texts": [ + " The contributions of this paper include: a) verification of rolling motions of microballs by directly observing trajectories of all bearing elements (stator, slider, balls); b) observation of the hysteresis between the friction and the relative velocity; and c) investigation of the influences of the ball number and oxide formation on the frictional behavior. 1 * Corresponding authors. X. Tan is currently affiliated with the Department of Electrical and Computer Engineering, Michigan State University, East Lansing, MI, 48824. Email: xtan@egr.msu.edu (X. Tan), ghodssi@eng.umd.edu (R. Ghodssi). nloaded From: https://proceedings.asmedigitalcollection.asme.org on 01/04/2019 Terms of Use: A schematic of linear microball bearings is shown in Fig.1. The bearing consists of two silicon plates (slider and stator) and stainless steel microballs of diameter 285 \u00b5m. Two parallel Vgrooves, which house the balls, are etched on the plates using potassium hydroxide (KOH). In the experiments the stator of the bearing is mounted on an oscillating platform driven by a servo motor. The only force acting on the slider is the friction applied by the microballs. By collecting motion data of the bearing elements through a CCD camera, one can infer the friction dynamics inside the system" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003455_jmes_jour_1980_022_045_02-Figure11-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003455_jmes_jour_1980_022_045_02-Figure11-1.png", + "caption": "Fig. 11. The piece of paper to make the model shown in Fig. 12", + "texts": [], + "surrounding_texts": [ + "THE FOLDING OF DEVELOPABLE FLAT SHEETS O F METAL 237\ndS\u2019\ndd\n01\nFig. 8. The geometrical relationship for r 00 (this means that the right hand portion of the sheet in Fig. 6(a) remains a plane), p = O for any R and 8 and the sheet cannot be folded. (b) When R-tco and 8 # ~ / 2 (the folding line is now a straight line, not parallel to the generator of the cylinder), p=n for any Y and 8. Hence, in this case, two surfaces coincide.\nTable 2. The angles of fold, according to equation (1 6)\nPlr 0 0.1 0 . 2 0 .3 0 .4 0 .5 0.6 0.7 0.8 0.9 1.0\n(Ded 0 11.4 22.6 33.4 43.6 53.1 61.9 70.0 77.3 84.0 90 (Rad) 0 0.199 0.395 0,583 0.761 0.927 1,081 1.221 1.349 1,466 1 3 7 1\nP I P 1 .5 2 2.5 3 4 5 6 7 8 10 00\np(Deg) 112 6 126.9 136.4 143.1 151.9 157.4 161.1 163.8 165.8 168.6 180 Wad) 1.966 2.214 2.381 2.498 2.652 2.747 2.811 2.858 2.893 2 942 3.142\nJouindl Mechaoical Enginreling ScienLe g) IMechE 1980 Vol22 No 5 1980\nat University of Bath - The Library on June 5, 2016jms.sagepub.comDownloaded from", + "238 W. JOHNSON AND T. X. YU\nThus, using (16), we obtain p= 151.9\", when B = O ;\nand p= 113.7\", when 1 9 = 0 * = 5 1 . 8 O .\nThe above result is exemplified in the paper-model shown in Fig. 12.\n(a)\n(C)\nJournal Mechanical Engineering Science Q TMechE 1980\n3.5 Folding on a cone along a circular arc As an example of the application of expression (16), consider a sector of a circle folded into a cone, along a circular arc (see Fig. 13). Here, 8=0 along the folding line, so that R cos2 8 = R , Because Y = R . tan a, where CY is the semi-vertex angle, then from ( I 6),\n~- 2. tan-' (cot a) = T - 201 R cos2 0 /3 = 2 tan-'\nr\nTo exemplify this case, take a piece of paper, cut it as a sector of a circle and draw on it a circular arc, as in Fig. 13(a). Fold it slightly along the circular arc and then make the top portion close-up to become the surface of a cone; the shape shown in Fig. 13(b) is naturally produced without buckling.\n3.6 In the curved line folding process, the plastic work done includes two parts-that dissipated in the folding line, and that dissipated by producing the curved surfaces; that is, The plastic work done in folding\nw= wz+ ws= W l f w,1+ wsz (18) where I denotes the folding line and sl and s2 denote the curved surfaces produced.\nIn this paper we only consider the first part of the plastic work, that is, we only discuss the energy dissipated\n(4\nFig. 12. A paper model for R=4r\nVol22 No 5 1980\nat University of Bath - The Library on June 5, 2016jms.sagepub.comDownloaded from", + "239 THE FOLDING OF DEVELOPABLE FLAT SHEETS OF METAL\n4 FOLDING ALONG A CURVED LINE /\nFig. 13. Folding on a cone along a circular arc\nplastically on the folding line. Referring to ( I ) , i n the general case of curved line folding, we have\nWr = \\, Mp/3 d.7 = M , s, ds (1 9)\ni.e., Wr is proportional to the integral of the angle of fold along the whole folding line.\nWhen the folding line is as shown in Fig. 6(a), ds= R do, /3 is given by (16) and\nW I - ~ M , \\: 2 tan (: cos2 8 ) R d0\nwhere\n(21) When Rlr and 8\": are given, I(%*) requires to be calculated numerically.\n4.1 The general case of folding on a cylinder In the general case of folding along a curved line, both R and 0 vary along it.\nAssume that the equation of the folding line in the initially flat sheet is\ny = y ( x ) for -X r r /2 (5) There is no benefit if the deviation of the new direction is larger 1than 90 \u00b0 from the goal direction. Configuration Uncertainty Task (CUT) For all n confirmed features lying within Rmax, the benefits Bc(~b)depend upon the size and shape of the uncertainty ellipse associated with the covariance matrix Q(k). Given the parameters of the ellipse (a, b, a), the estimated benefit of each confirmed feature Bc(~b) is proportional to the maximal projection wc of the ellipse onto the feature normal Pr\" (See Fig. 5.) t =Po - a, 1 w~= ~-v/a2(1 + cos(2t)) +b2(1 - cos(2/)) . (6) Feature Map Task (FMT) For all m tentative features lying within Rmax, the benefits Bm(cb) depend upon the difference of their plausibility and uncertainty and the respective thresholds which promote tentative to confirmed features. The smaller the difference the larger the benefit since these features can then be used for localisation purposes. Another point of view (not discussed here) would be to treat the benefits as the expected information profits" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002394_robot.2001.933236-Figure7-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002394_robot.2001.933236-Figure7-1.png", + "caption": "Fig. 7. Modified path that avoids the obstacles.", + "texts": [ + "24m and R, = O.lm respectively, see Fig. 5. The criterion that the additional coefficient must meet a,, b;; + P,, b4 + Y,, > 0 The roots of the second order polynomials in b4, involved in Eqs. (22) and (24), are plotted for each obstacle, resulting in Fig. 6. This figure also illustrates the possible forms of the root plots. The selection of the value of the additional coefficient is now an easy task. For example, by selecting b, = -0.12, the obstacles are avoided for all orientations in the employed range. Fig. 7 depicts the modified path which avoids all obstacles. The corresponding system inputs are shown in Fig. 8. As shown in Figs. 7 and 8, the resulting path is smooth with a continuous curvature profile, and requires no excessive input velocities. If the obstacle distribution is such that one coefficient cannot lead to a modified path that avoids all obstacles, then more coefficients should be used, depending on the distribution of the obstacles. In this case, N inequalities will result representing the distance from each obstacle, although they will be of the form (26) ,=4 for i = l , " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003638_s021812740401103x-Figure7-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003638_s021812740401103x-Figure7-1.png", + "caption": "Fig. 7. The six kinds of enveloping surfaces \u2202Vh,l in \u03c9-space, for Lagrange tops. The two red surfaces as well as the blue and green are computed for A1 = 1, \u03b1 = 1.5 (\u03b1-range III), the magenta and yellow surfaces for A1 = 2, \u03b1 = 0.505 (\u03b1-range I). In the upper row, the values (h, l) are (4.5, 3), (0.85, 0.3), (1.1, 0.3); in the lower row they are (3.1, 3), (1.851, 1.85), (2.307, 2.1). The direction \u03c9 = e1 points to the left. All surfaces possess rotational symmetry with respect to e1; one quarter has been cut away to make the inside visible.", + "texts": [ + " We shall see that, in general, the topology of P2 h,l and \u2202Vh,l depends not only on the topology of E3 h,l but also on details discussed in Secs. 4 and 6. For the Lagrange case, it turns out that all manifolds P2 h,l are two-tori: sin- gle T 2 in the red, blue and green regions, pairs of T 2 in the yellow and magenta regions. The projection of P2 h,l to \u03c9-space possesses, in general, up to three singular points on the axis where l = A\u03c9 is collinear with r. For the Lagrange case this means there may be up to three singular points on the e1-axis, see Fig. 7. Their number is 0 in the green region of the bifurcation diagram, 1 for both components of \u2202Vh,l in the yellow region, 2 in the blue region, 1 and 0 for the components related to S3 and S1\u00d7S2 in the magenta region. The red region deserves special attention. It is separated in two parts by the line h3 = 27l2/8A1, 0 \u2264 h \u2264 3/2, see Eq. (49) and Fig. 6 where the bifurcation sets \u03a3 are shown for the same parameters as in Fig. 3, together with the extra line (color magenta). This line is not related to singularities of E3 h,l or P2 h,l; rather it reflects a singularity of the projection. For values (h, l) in the red region to the left or above the line, the surfaces \u2202Vh, l have one singular point; for (h, l) to right and below the line, they have three, cf. the two red pictures in Fig. 7. The envelopes enclose the regions of admissible angular velocities. Trajectories \u03c9(t) remain in Vh,l and are tangent to \u2202Vh,l, where dl2/dt = 0. The outer part of the envelopes corresponds to local maxima of l2, the inner part to local minima. The extra constant of motion ls = \u3008l, r\u3009 of the Lagrange case implies that the trajectories are restricted to planes \u03c91 = ls/A1. The intersections of these planes with \u2202Vh,l are the invariant foliation of \u2202Vh,l by projections of Liouville\u2013Arnold tori from the reduced phase space to \u03c9-space" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003634_cdc.2004.1430205-Figure2-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003634_cdc.2004.1430205-Figure2-1.png", + "caption": "Fig. 2. The configurations (up to SE(2)-translation) from which the hovercraft system is controllable", + "texts": [ + " One result extends Theorem 9 to a situation where BYq0 is semidefinite. The other result is a structural result for two-input systems with one decoupling vector field. This is an interesting application of the vector-valued quadratic form technology. Then we look at the controllability of the hovercraft. As we shall see, the system \u03a3hc has quite a complicated structure as concerns its controllability. To summarise, we shall show that the system is accessible from all configurations and that the only configurations from which the system is STLC are those shown in Figure 2, and any SE(2)-translation of the configurations shown. Let us first provide some explicit expressions for the vector fields Y1, Y2, and some of their symmetric products so that we can see how the analysis might proceed. The SE(2)-invariance of the system is useful here since we may without loss of generality evaluate all symmetric products at (x, y, \u03b8, \u03c6) = (0, 0, 0, \u03c6), essentially meaning we evaluate at the group identity. Thus in the expressions immediately below this simplification is tacitly made", + " What\u2019s more, Lemma 4 gives the nice interpretation of the integral curves of Y2 as being zero SE(2)-momentum moves for the system. Thus we are indeed in the case described by Theorem 15. It turns out that we are in case (v) of the theorem, and that it is possible to explicitly describe the analytic subset S. The following result summarises this. Proposition 19: Let q0 = (0, 0, 0, \u03c6). The following statements hold: (i) \u03a3hc is STLC from q0 if sin(2\u03c6) = 0; (ii) \u03a3hc is not STLCC from q0 if sin(2\u03c6) = 0. Thus the system is STLC only from the configurations shown in Figure 2. Remarks 20: 1) The system has the interesting property of possibly being STLC at a point (points where sin(2\u03c6) = 0), but of not being STLCC at points in a neighbourhood at the point. A similar, but simpler, example was examined in [16]. 2) It is interesting to contrast our hovercraft model with the simpler ones considered in, for example, [1], [3], [4], and described briefly in Remark 5. In this case, the system is STLC from every point in Q [1], and is furthermore kinematically controllable [4]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000356_0021-8928(96)00040-8-Figure5-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000356_0021-8928(96)00040-8-Figure5-1.png", + "caption": "Fig. 5.", + "texts": [ + " In this connection it would be of interest to investigate the mechanical meaning of the effect of a parallel transfer of the vector along a dosed contour on the manifold of positions of a rigid body with one fixed point (group SO(3)). In conclusion we note that the application of the solid-angle theorems formulated above is outside the framework not only of Euler's case but also of rigid-body dynamics in general. These theorems can be used in all cases when the problem reduces to analysing the mutual angular position of two trihedra. For example, suppose that the mean length of a thin inextensible tape, which is absolutely flexible in one direction and absolutely rigid in the other, is bent into a certain spatial curve (Fig. 5). The orientation of the initial section of the tape with respect to the end section can be determined as the orientation of trihedra connected with these sections. If the initial trihedron is shifted along the mean line so that the vector i touches the median line, and the vector k is situated along the normal to the tape, the projection of the angular velocity of the trihedron onto the k axis is zero (the condition for absolute rigidity of the tape in the i, j plane) and the answer to the problem in question is given by Theorem 2, in which 0h(t) =- 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003656_s00397-004-0377-4-Figure15-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003656_s00397-004-0377-4-Figure15-1.png", + "caption": "Fig. 15 Statistics of the likelihood of convergence to attractors from the nematic rest state at a fixed concentration N=6, for the K1 and W bi-stable region with Pe=2.3. The lighter points represent initial director configurations that will converge to K1, whereas the black points converge to W", + "texts": [ + " In the bi-stable W-LR regime, the likelihood of LR varies between 15% and 32%, but then jumps to 41% likelihood when bi-stable with T limit cycles. Out-of-plane bi-stable and tri-stable statistics From Fig. 2 and Table 2, we have a bi-stable K1 and W regime. Whenever the tilted kayaking limit cycle K2 is stable, it always has its bi-stable twin, on the other side of the shearing plane. We also have a tri-stable region of K1 and K \u00fe; 2 . We now measure the statistical likelihood of converging to each attractor from nematic equilibrium data. Case 1. When Pe=2.3 and N=6, the K1 and W bistable regime, Fig. 15 shows the statistical likelihood of convergence to K1 and W as the initial nematic director at rest is sampled across the sphere. \u2013 If the initial directors ~n have polar angles h0\u202185.5 , then all the data converge to the in-plane W attractor. Therefore, initial director orientations that are sufficiently close to the shearing plane are attracted to the W attractor. This has important consequences since plate preparations in Couette cells can strongly prejudice the director orientation of the rest state" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002119_1097-4563(200012)17:12<643::aid-rob1>3.0.co;2-8-Figure16-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002119_1097-4563(200012)17:12<643::aid-rob1>3.0.co;2-8-Figure16-1.png", + "caption": "Figure 16. Rotation angle between Gauss frames for two objects in contact.", + "texts": [ + "z u , v g Q u , v 21 The Gauss map for an arbitrary surface map can be \u017d .calculated as the cross product of x u, v and \u017d .y u, v . \u017d . \u017d \u017d .. \u017d . \u017d . \u017d .z u , v g Q u , v x u , v y u , v 22 When two objects are in contact, a Gauss frame is defined at the point of contact on each object. The normals at the point of contact are collinear and have opposite signs. The x and y axes of the two Gauss frames lie in the same plane but will likely not be aligned and will be offset by some rotation \u017d .angle around the z axis. Figure 16 shows the Gauss frames for two objects in contact. The indi- \u017d .Figure 14. Surface map Q u, v . vidual Gauss frames are shown and then they are superimposed, showing the angle between them. The metric of a surface is used to measure the changes in the position on the surface for a differential change in the parameters u and v. The general equation for the metric of the surface is \u017d . Q u , v 0u \u017d .M 231 \u017d . 0 Q u , vv Curvature at a point on a surface is a measure of the rate of change of the surface normal in a particular direction at a point on the surface14 and essentially characterizes the local shape of the surface in an intrinsic manner, independent of scale and orientation" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000888_0301-679x(91)90050-j-Figure2-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000888_0301-679x(91)90050-j-Figure2-1.png", + "caption": "Fig 2 Qualitative velocity profiles shown by arrows; 6- distribution indicated by A B C and EFG, and 6*- distributions shown by broken lines", + "texts": [ + " However , the an t i symmetry necessitates considerat ion of all the four regions. This is easily visualized for the Newtonian case (n = 1) as follows. Equat ions (1) and (2) can be solved for n = 1 and m to be constant or a funct ion of x alone; one can obtain u = Ui + ( U 2 - U 1 ) h - 2m ( l(1) dp = 6 ( U , + U a ) m ( h _ h , ) / h 3 dx ( l l ) It can be seen that u is a linear function of y at points where dp /dx = 0. At these points, the velocity gradient Ou/Oy = ( U 2 - U ~ ) / h l which can never be equal to zero, since U2 4= U~. It can also be seen f rom Fig 2 that for each x, Ou/Oy can vanish at one point of y ( = 5 , 0 <~ 5 ~< h) in the regions -~c < x < -x~ and -xL < x < x2. Thus these regions can be divided into four sub-regions separa ted by the 8-profile, having velocities u~, u2, u3 and u4 as shown in Fig 2. For any e > 0, let us and u~, be the velocities in the regions -x~ - ~ ~< x ~< -x~ + ee and x~ - e3 ~< x ~ x2 respectively, 240 August 91 Vol 24 No 4 then according to the velocity profile, shown in Fig 2, one may observe the following: 0u~ >0, a0, 0~ to than for t < to (e.g. MD1 in Figure 3). to = t shall be the current instant of time and f(to) the corresponding function value. Starting with p0=(to, f(to)), we compute a set of N + I points P\u201d\u2019, N E (0, 1, ..., N } on f with the property that they are equiarcide, i.e., the curve length A1 on f between P\u2019 and PI-\u2019 is constant for all i E { I , 2 , ..., N } . * \u2019 9 Figure 6 shows an example with N = 5 before, during and after the state transition. Figure 6a corresponds to the situation before the transition, i.e., to < to. For a piecewise linear function f, the P\u2019 are equidistant not only in the curve length but also in time, i", + " The thin black line represents a trigger signal computed with N = 10 equiarcide curve segments and a detection threshold k = 2.5. Here, the transition could be detected reliably in several repetitions of the experiment. The length AI of the single arcs is selected automatically for each sample, depending on the local noise level. The details of this selection, however, are beyond the scope of this paper. The last example demonstrates a case where a detection based on the MSL residuum (as well as several other methods) fails. The experiment is a transition sequence EE -+ V/E -+ V/F, similar to Figure 3 with motion direction MDz being perpendicular to gravity. As DLO, the spring-steal ruler described above is used. In this case, the detection is especially critical due to the following reasons: First, the force decrease at the transition to V/F is rather small. Second (and more important), there are remarkable distortions due to slip-stick and rebounding. and the finite resolution of the FMS Figure 9 shows the moment signal measured by the FMS with the transition occurring at io = 630. As indicated by the trigger signal, the transition is recognized by the equiarcide-points method with a remarkable delay Ai = 160 4sec" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003483_imece2004-59492-Figure6-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003483_imece2004-59492-Figure6-1.png", + "caption": "Fig. 6: Different protuberance values", + "texts": [ + " 5 shows the case of undercutting between curves t1 and t3. Intersection B is known because it is generated by the limits of the two rack profiles r1 and r3, while the intersection A is evaluated with the previously mentioned numerical method, using \u03c6u instead of \u03c612 in eq. (35). wnloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 02/07/2016 Te Particular cases According to the values of the protuberance parameters, two particular cases can be obtained by the enveloping process. Figure 6 shows two situations, which are theoretical cases and are not included in standard protuberance parameters. In the first case (fig. 6(a)) the protuberance segment r2 generates the small involute curve t2 on the tooth in same way as r3 forms t3. So two intersection A and B must be calculate using numerical method. In the second case the highest values of protuberance r2 annihilates the actual envelope of the segment r3. A criterion to understand when these two situations occur is to evaluate the intersection \u03c6p between curves t1 and t3. If the segment r2 does not envelop the real tooth flank the following condition must be true: 31p\u03c6 \u03c6> (38) If the protuberance is such that the segment r3 does not exist, the point correspondent to \u03c631 is above the tangent line to the circle tip. This implies: 31 32\u03b8 \u03b8> (39) Equations (38) and (39) take into account the curves orientations as in fig. 6. 6 Copyright \u00a9 2004 by ASME rms of Use: http://www.asme.org/about-asme/terms-of-use D 2D-Gear generator program All the developed equations were implemented in a FORTRAN code. The code is able to generate an output script file for AutoCAD 2002 containing instructions to draw a complete 2D gear. The input file is insert using dialog box and must contain the following rack and gear information: - Module; - Number of teeth; - Pressure angle; - Addendum modification; - Rack fillet radius; - Protuberance; - Protuberance angle; Figure 7 shows different teeth obtained from a rack without protuberance, with pressure angle 20\u00b0, module 1 mm and tooth fillet 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002872_1.1456456-Figure7-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002872_1.1456456-Figure7-1.png", + "caption": "Fig. 7 Free-body diagram of the composite cylinder at elevated temperature", + "texts": [ + " 6, which is similar to the preload model in Harris @11#. The inside and outside cylinders model the bearing outer race and the hub, respectively. In this model, subscript 1 refers to the inner cylinder and subscript 2 refers to the outer cylinder. In addition, E is the Young\u2019s modulus, n is the Poisson\u2019s ratio, a is the coefficient of thermal expansion, ri is the inner radius, and ro is the outer radius. When the temperature is increased by DT , an uniform radial stress field p develops between the inner and outer cylinders; see Fig. 7. Note that if a2.a1 , p.0 ~i.e., tensile stress!. In general, the radial displacement u of each cylinder consists of two components up and uT , or more explicitly, u1~r !5up11uT1 , u2~r !5up21uT2 , (1) where up is the displacement resulting from the radial stress p, and uT is the displacement from the thermal expansion. Consider the inner cylinder first. According to linear elasticity @12#, when a thick elastic cylinder is subjected to tensile radial stress p at the outer radius, the radial displacement is up15A1r1 B1 r , (2) ology gy" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002756_0954405011519402-Figure2-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002756_0954405011519402-Figure2-1.png", + "caption": "Fig. 2 Laser\u00b1powder interaction surface", + "texts": [ + " If not, the surface roughness will be related to the powder particle size. The \u00aenal sintered density s is assumed to be dependent on the pulse energy El and the surface energy density el: s \u02c6 s\u2026El, el\u2020 \u20265\u2020 Particle size and size distribution can also have a strong in\u00afuence on the sintered density. They are not varied in this work and can therefore be treated as constant material properties. In the determination of the sintered depth ld it is necessary to investigate the powder surface for a small surface area dA exposed to the laser hatching pattern (Fig. 2). B11800 # IMechE 2001 Proc Instn Mech Engrs Vol 215 Part B at Northeastern University on November 28, 2014pib.sagepub.comDownloaded from A surface energy density ew is assumed to be wasted (i.e. not utilized by a material phase change). The mass molten for the small area dA scanned is dmm and the mass vaporized dmv. If two e ciencies, a heating e ciency \u00b2h \u02c6 \u2026el \u00a1 ew\u2020=el and a melting e ciency \u00b2m \u02c6 dmm=\u2026dmv \u2021 dmm\u2020, are de\u00aened, the following can be written: mm \u02c6 dmm dA \u02c6 \u00b2hel \u20261=\u00b2m \u00a1 1\u2020hsg \u2021 hsf \u20266\u2020 which is termed the molten mass area density" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001226_20.717826-Figure4-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001226_20.717826-Figure4-1.png", + "caption": "Fig. 4. Distribution of flux density, I A = I B = 0 A.", + "texts": [ + " Permanent magnets are modeled by using two parameters, namely, magnetization and permeability. Here, we assume permeability in the magnet is equal to that in the air for simplicity. Fig. 3 shows the mesh for the stator iron of the target motor. The number of nodes is 11550, and the number of elements is 59904. Periodic boundary condition is set at r-z plane[2], and Dirichlet boundary condition is set at other surfaces. Static torque is calculated using Maxwell's stress tensor and the difference of magnetic energy. 111. ANALYSIS OF RESULTS A. Flux Distr.ibution Fig. 4 shows magnetic flux distributions without exciting current. Small flux density (B<0.4 T) is ignored in these figures. Fig. 4a shows the flux density distributions in the case that rotor position 81,L=150, that is, the magnets of the rotor are at a position corresponding to the teeth of phase A. Large flux density appears in the teeth of phase A. Fig. 4b shows the flux density distributions in the case that 8,,,=22.5\", that is, the magnets of the rotor are at a position corresponding to the teeth of phase B. Then, large flux density appears in the teeth of phase B. In the case that 8,=Oo and 7.5\", the flux densities have the same amplitude as and the opposite direction to those in the case that 8,,=15\" and 22.5\", respectively. Fig. 5 shows magnetic flux distributions with exciting current I ~ = 0 . 3 3 A. Excitation of phase A produces flux density in the lower section of the motor" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002212_1350650011543484-Figure6-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002212_1350650011543484-Figure6-1.png", + "caption": "Fig. 6 Diagrammatic view of oblique impact", + "texts": [ + " Initially, the effect of off-centre normal impact was studied, since no traction force was exerted and this provided a good test case to study the bending effect. In Figs 6a and b, the key parameters and their effects under oblique impact are illustrated. The top block was subjected to the normal force Fn and traction force Ft of the ball where the impact spot was assumed to be a distance d away from the centre line. The height and width of the block above the transducer were assumed to be 2L and 2R respectively. Because of the off-centre normal and traction forces, the top block rotated around its centre of mass, as shown in very exaggerated form in Fig. 6b. The effect of bending on the measured tangential force under oblique impact and different amounts of offset is shown in Fig 7. The highest tangential force was measured when the impact position was at the centre of the plate (i.e. d in Fig. 6a was zero), although a small negative force was measured during the first tens of microseconds. As the impact point moved to the rear side of the plate (i.e. negative d), the initial negative force decreased in magnitude, until it flattened off and eventually became positive at the start of impact. It was noted that, if the impact point was displaced from the geometrical centre of the impact plate by more than 1 mm, vibrations of the force plate swamped the tangential force trace. The experimental observations as shown in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003496_s0022-3913(75)80139-9-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003496_s0022-3913(75)80139-9-Figure1-1.png", + "caption": "Fig. 1. Hanau XP-51 articulator: (A) horizontal condylar guide thumbscrew, (B) Bennett guide thumb nut, (C) adjustable posterior wall thumb nut, (D) auditory pin, (E) third reference locator, (F) disposable mounting ring, (G) individual incisal guide table, (H) custom-built incisal guide table, (I) incisal guide pin, and (J) face-bow support groove on incisal guide pin.", + "texts": [ + " In this article, an articulator will be discussed in which design and construction are based on the theory that an intercondylar adjustment is not required. A posterior ,,vail adjustment in the condylar housing will compensate for the lack of an intercondylar adjustment. *Assistant Professor, Department of Removable Prosthodontics. **Director of Engineering and Development, Teledyne Dental, Hanau Division, Buffalo, N.Y. 10 Voh, mc33 New semiadjustable articulator. Part I 11 Number 1 The Hanau XP-51 articulator is of the arcon type. It has a fixed distance between the condylar posts of 90 ram. (Fig. 1). The adjustable guide within the condylar housings are (A) the Bennett guides, (B) the horizontal condylar guides, and (C) an adjustable posterior wall (Fig. 2). This posterior wall can be adjusted to the position of the working-side condyle in lateral movement. When adjusted, this wall contacts and forms a continuous guiding surface for the working-side condyle. Further, the posterior wall combines to compensate for the lack of an intercondylar distance adjustment on the articulator. The angle so formed is called the \"compensating angle" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000288_0094-114x(95)00086-e-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000288_0094-114x(95)00086-e-Figure1-1.png", + "caption": "Fig. 1. Cam mechanism: (a) physical system [7], (b) the corresponding dynamic model.", + "texts": [ + " This model was further modified by Hanachi and Freudenstein [4] by introducing additional features such as the inclusion of Coulomb friction, the hydraulic lifter and damping due to nested springs. Pasin [5] used a model in which the pushrod and the valve spring are continuous elements to calculate the natural frequencies of a force closed cam mechanism. Tiimer and Onltisoy [6] calculated the valve spring preset using Pisano and Freudenstein's model in which they considered only the separation of the follower from the cam. In this paper two different models are presented to investigate the separation phenomena in the force closed cam mechanism illustrated in Fig. 1 (a). The first model is similar to the one used by Pasin. Pasin's model does not include damping, however in the present study damping of the system is represented by a viscous damper. Compared to Pisano and Freudenstein's model, this model considers both the valve spring and the pushrod as continuous systems. In the second model, as the stiffness of the valve spring is less than that of the pushrod, the pushrod is considered to be rigid. 487 In a cam-follower mechanism train, in addition to the separation of the tappet from the cam, the separation of other members of the mechanism which are kept in contact through the valve spring force, must be investigated" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000132_70.488954-Figure5-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000132_70.488954-Figure5-1.png", + "caption": "Fig. 5. (a) Possible unexpected results due to the position accumulation error. (b) Incremental collision test considering the position accumulation error.", + "texts": [ + " It can be seen as an attractive force exerted by the C-obstacle keeping the contact until the attraction force exceeds the sticking factor. 1 ) Evaluation of the PositiodOrientation Uncertainty: The position uncertainty must be considered differently depending on the space where the robot navigates. While the robot moves in free space, a positioning error is accumulated due to the control error associated to each velocity commanded. Such error accumulation can cause the robot to collide with an unexpected obstacle or to m i s s a target one when the nominal path is executed by a real robot (Fig. 5(a)). To cope with this problem, the radius p of the position uncertainty disk is augmented on each iteration according to the control error 7. A simple vicinity collision test between the disk and the neighbor C-obstacles is performed on each iteration to determine possible eventual collisions due to control errors (Fig. 5(b)). When a \u201csafe\u201d collision occurs, the position uncertainty disk becomes a segment over the contacting surface. When the robot moves along the surface, this segment continues to grow because its exact position is unknown. The position uncertainty is completely elimnated when the robot detects a vertex (because its position is accurately known). The orientation uncertainty $ increases while the robot rotates in free space. This uncertainty can only be reduced when the robot is in contact with an obstacle" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003682_s10237-004-0050-y-Figure3-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003682_s10237-004-0050-y-Figure3-1.png", + "caption": "Fig. 3 Sketch of the simulated system: a thin tube is subjected to a time-dependent internal pressure p; two families of reinforcing fibers are defined along two skewed directions with symmetric angles + / and / with respect to the longitudinal tube direction", + "texts": [ + " A sheet of preserved bovine pericardium is wrapped in order to obtain a thin tube with internal and external radii Ri = 1 mm and Re = 1.01 mm, respectively, and thickness U = 0.01 mm in its reference configuration. This example is aimed at simulating the inflation of a vascular patch made of bovine pericardium used in vascular surgery. A cylindrical reference system is assumed by setting the reference coordinates (r, h, z) along the radial, circumferential and longitudinal directions of the tube (Fig. 3). Let kr, kh, kz denote the associated stretches of the tube. The preferred and cross-preferred directions of the collagen fibers within the sheet are aligned, in the undeformed tube configuration, to two directions represented by the unit vectors NaT \u00bc 0 cosu sinu\u00bd and NbT \u00bc 0 cosu sinu\u00bd ; u denotes the angle between the preferred fiber direction and the tube longitudinal direction; and it is assumed to be equal to p/4. The tube is subjected to a time-dependent internal pressure p(t) and to a prescribed longitudinal constant stretch, kz = 1" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002855_s0261-3069(01)00038-3-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002855_s0261-3069(01)00038-3-Figure1-1.png", + "caption": "Fig. 1. Schematic of filament feed and deposition mechanism for FDC.", + "texts": [ + " The liquefier extrudes a continuous road of material through a nozzle (typically 254\u2013762 mm) and deposits it onto a fixtureless platform. The build strategy (tool path) of an object to be manufactured is based on a CAD file and is generated by the machine software (in this case Quickslice ). The movement of the liquefier in the X- and Y-directions is hence controlled by the computer. When deposition of the first layer is completed, the fixtureless platform indexes down, and a second layer is built on top of the first layer. This process continues until the whole part is completed. This process is schematically illustrated in Fig. 1. *Corresponding author. Using Honeywell\u2019s GS44 Si N as the ceramic mate-3 4 rial, it has been demonstrated that ceramic green bodies with complex shapes can be successfully fabricated by FDC w5x. Fig. 2 shows FDC\u2019ed Si N parts. The suc-3 4 cessful production of high performance ceramics by the FDC process requires that no defects exist in the green parts. Build defects, however, such as missing roads, poorly bonded roadsylayers or sub-perimeter voids are sometimes encountered in FDC parts built under nonoptimized conditions" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003437_0954406042690498-Figure5-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003437_0954406042690498-Figure5-1.png", + "caption": "Fig. 5 (a) Cutting edge of the blade for cutting the gear tooth surface. (b) Relationship between coordinate systems S4 and S6", + "texts": [ + " For du \u00bc 0, the principal curvature k\u00f03\u00deI and direction e \u00f03\u00de I, 3 can be determined by k\u00f03\u00deI \u00bc dn3 ? dr3 dr3 ? dr3 du\u00bc0 , e \u00f03\u00de I, 3 \u00bc dr3ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi dr3 ? dr3 p du\u00bc0 \u00f04\u00de For db \u00bc 0, the principal curvature k\u00f03\u00deII and direction e \u00f03\u00de II, 3 can be determined by k\u00f03\u00deII \u00bc dn3 ? dr3 dr3 ? dr3 db\u00bc0 , e \u00f03\u00de II, 3 \u00bc n36e \u00f03\u00de I, 3 \u00f05\u00de The cutting edge for cutting the gear tooth surface is also designed as a parabolic curve with a parabolic parameter k2 as shown in Fig. 5a. This cutting edge, denoted by Gg, is rigidly connected by coordinate systems S8\u00f0o8; x8, y8, z8\u00de and S6\u00f0o6; x6, y6, z6\u00de, and its equation is represented in S8 by fx8, y8, z8g \u00bc fv, k2v2, 0g \u00f06\u00de The cutting edge Gg forms a surface of revolution after rotating about the axis of the disc. The coordinate system S4\u00f0o4; x4, y4, z4\u00de is applied to connect rigidly to the surface of revolution and the coordinate system S6 is provided with pure rotation about the axis of the disc with parameter y. As shown in Fig. 5b, S6 coincides with S4 when y \u00bc 0. The surface of revolution formed by Gg C04304 # IMechE 2004 Proc. Instn Mech. Engrs Vol. 218 Part C: J. Mechanical Engineering Science is denoted by S4, whose equation can be obtained by transferring the coordinates of Gg from S8 to S3 according to R4\u00f0v, y\u00de \u00bc M46\u00f0y\u00deM68R8\u00f0v\u00de \u00f07\u00de where R4\u00f0v, y\u00de \u00bc fx4, y4, z4, 1gT, R8\u00f0v\u00de \u00bc fx8, y8, z8, 1gT M46\u00f0y\u00de \u00bc Tx\u00f0 rg\u00deRy\u00f0 y\u00deTx\u00f0rg\u00de, M68 \u00bc M57 The unit normal vector to S4 can then be determined by n4\u00f0v, y\u00de \u00bc qyr46qvr4 kqyr46qvr4k \u00f08\u00de where r4 \u00bc fx4, y4, z4gT" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002477_iedm.1997.650441-FigureI-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002477_iedm.1997.650441-FigureI-1.png", + "caption": "Figure I . Schematic of a display backplane of thickness d bent around a radius of curvature r.", + "texts": [], + "surrounding_texts": [ + "Introduction\nThe use of altexnative substrates to glass is one area of increasing interest to the flat-panel display industry. The primary issues which alternative substrates are meant to address are a reduction in the weight of the display and an alleviation of the problem of display breakage during manufacturing and use. Many products, such as cellular telephones, personal digital assistants, hand-held electronic games, and countless other portable electronic devices which currently use glass as a substrate material may benefit from a lighter, more durable substrate. When this substrate material is flexible, in addition to being lightweight and rugged, displays may find new applications in areas which have traditionally been considered too harsh and severe. The use of a flexible, even foldable, substrate material such as thin stainless-steel foil, which is lightweight and rugged, opens up the possibility for new display products such as foldable intelligent maps. The use of this flexible but opaque substrate would nlxessarily eliminate the use of backlighting. Emissive or reflective displays, however, remain viable options for such backplanes. The goal of our work is to demonstrate the feasibility of such a display.\nThe foldability of thin-film displays may be limited by the stress E in the thin-film layer that develops when the substrate is bent. Fig. 1 shows the substrate and thin film composite of total thickness d wrapped around a radius of curvature r. When r >> d, the stress E is given by\nY d 2 r E=-.-\nwhere Y is Young's modulus (1). With the stress being proportional to the d/r ratio, very thin foil substrates become highly desirable if a flexible display is the goal. Because the\nsubstrate also provides mechanical integrity to the display, strong materials are best. For these reasons, we have chosen stainless steel foil.\nExperiment\nAn active-matrix thin film display requires the integration of a light-emitting device with a switch (2). In our experiment, we integrate organic light-emitting diodes with amorphous silicon thin film transistors on flexible, ultra-thin substrates of 25 pm thick stainless steel foil. A schematic cross-section of the integrated TFT/OLED structure on the steel foil is shown in Fig. 2(a), and its equivalent circuit is shown in Fig. 2(b). As-rolled 304 stainless steel foil serves as the substrate. Planarization with 0.5 pm thick spin-on glass removes the short-wavelength roughness of 0.3pm rms. This planarization, necessary for high transistor yield, functions as primary insulation. Further insulation is provided by a 0.5 pm thick plasma-enhanced CVD SiN layer. The TFrs are made in the inverted-staggered, back-channel etch configuration with 120 nm thick Cr gates, 400 nm PI?CVD gate SiN dielectric, 150 nm a-Si:H channel, and 50 nm n+ aSi contacts, followed by 120 nm Cr source/drain contac.ts (3). The channel length and width are 42 pm and 776 pm, respectively. We employ such large TFTs, and also large contact pads, to ease probing and diagnosis on bent or rolled substrates.\nFollowing the fabrication of TFTs, OLEDs were made on the surface of the Cr source/drain contact pads of 2 mm x 2\". Conventional OLEDs are built on transparent sublstrates\n20.6.1 0-78034100-7/97/$10.00 0 1997 IEEE IEDM 97-535", + "coated with a transparent hole-injecting anode contact such as indium tin oxide (ITO), so that light can be emitted through the substrate, because the top contact is an opaque electroninjecting metal cathode. Because of the opacity of the steel substrate, we developed the top-emitting structure as shown in Fig. 2(a), in which the high work function metal Pt functions as the reflective bottom anode and a semitransparent cathode is applied on top. OLEDs were fabricated by sequential e-beam deposition and patterning of 400 8, Pt anode contacts, spin-coating of a continuous layer of 1700 8, active luminescent molecularly doped polymer (MDP), followed by the e-beam deposition of a 140 A semitransparent Ag top cathode. The overlap of the anode and cathode contact areas determines the active OLED device area, a 250 pm diameter dot, without the need to separately isolate the organic layers. All OLED fabrication steps were performed at room temperature and are compatible with finished TFTs (4).\nThe active organic material used is a single-layer MDP thin film. The hole-transport matrix polymer poly(Nvinylcarbazole) (PVK) contains dispersed electron-transport molecules 2(-4-biphenyl)-5;(4-tert-butyl-phenyl)- 1,3,4- oxadiazole (PBD) and a small amount of green fluorescent dye, Coumarin 6 (C6), which provides efficient emission centers (4).\nResults\nA. Characterization of TFTs\nThe TFT/OLED structures were fabricated on 38 mm x 38 mm steel foils. To test the mechanical resiliency of the TFTs alone (before OLED integration), one of these foils was scissor-cut into strips 4 mm wide and 38 mm long. The asprocessed TFTs had threshold voltages VT = 3.5 V, field effect mobilities yE = 0.9 cm2N.s, OFF currents IOW U Nym, and ON currents ION = lo-* Npm, as shown below in Figure 3(a). These strips were later rolled along their length and held for one minute under both convex (TFTs facing outward) and concave (TFTs facing inward) bending, and tested for topological integrity and electrical performance.\n20.6.2 536-IEDM 97" + ] + }, + { + "image_filename": "designv11_65_0001902_3.20413-Figure7-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001902_3.20413-Figure7-1.png", + "caption": "Fig. 7 Transition profile and shipboard landing task.", + "texts": [ + " The landing pad is presented in both horizontal and vertical aspects. In the horizontal aspect, the pad symbol is geometrically similar to the Spruance -class destroyer pad Simulation Conduct Facility The VMS provides large-amplitude vertical and longitudinal translation capability, and provides high-fidelity motion cues. A continuous, three-window, computer-generated imaging (CGI) system (no window mullions) provides unobstructed views of outside scenes. Simulation Tasks Shipboard transition and landing (Fig. 7), and land-based precision hover tasks were used for the evaluation of the candidate control systems. The shipboard tasks consisted of three distinct phases: transition to hover, translation to a point over the ship's landing pad, and final, descent to landing (the term \"landing\" includes the last two phases). Transition The transition starts at a speed of 120 knots relative to the ship, along a reference flightpath with a glide-slope of \u2014 3 deg. The reference deceleration is approximately 3 ft/s2, stepping down to approximately 1" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002190_70.850647-Figure4-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002190_70.850647-Figure4-1.png", + "caption": "Fig. 4. Definition of work-space during obstacle evasion.", + "texts": [ + " Consider a robotic manipulator that must develop a PPO from a given initial pose specified by the position of one of the end-effector points and oriented with respect to a certain coordinate frame, to a final pose specified likewise. Consider that the end-effector path must evade the point an arbitrary obstacle placed in the work environment. Let be the radius of the spherical space into the work space, which must be evaded by the end effector, to ensure that any collision with the obstacle would be impossible. Realize that the three points defined by vectors lie in unique plane named in the work space, as is shown in Fig. 4. Hence, the above-defined sphere becomes a circle inside such a plane and the problem is simplified. The objective will be to find a point located at the perimeter of the circle and to constrain the trajectory to pass over it without entering into the circle zone. The equations of the straight lines and can be defined with the pair of points and , respectively. The line will represent the shortest and simplest path between and The equation of the straight line will be constructed containing the point perpendicular to and cutting it at The point could be obtained as follows: (21) with (22) On the same basis, the equation for the straight line will be (23) with (24) The points and are obtained by evaluating the vector function in and as shown in Fig. 4. These are the farthest and closest points to the line , respectively. The next step will be to impose the trajectory to pass over a selected point between and The selection of the most suitable point, will depend on the physical characteristics of the workspace and also on the desired trajectory. The objective will be to impose the end-effector trajectory to reach a configuration in Cartesian space, with a known velocity at a specific time that passes over the point To define the space an arbitrary but reachable orientation of the manipulator is utilized. To determine the joint variables corresponding to an inverse kinematics procedure is performed. The expressions (19) and (20), adequately handled, may be solved to obtain the values of the function and its derivatives defined as (25) Note that the configuration must lie between the initial and final configuration of the path which imposes a limitation on the radius of Fig. 4. To modify the hyperbolic trajectory the parameters and of (6), must be established, for all joint variables at a simultaneous time Depending on the position of the obstacle, one of the next three blocks of conditions should be selected as in (26), shown at the bottom of the next page, with (I) when is near the take-off positions, (II) when is in the middle positions, and (III) when is near the landing positions, as shown in Fig. 5(a) and (b). Fig. 6 shows an example of the dialog box to modify the parameter to obtain the desired shape of the trajectories" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000577_aim.1999.803236-Figure3-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000577_aim.1999.803236-Figure3-1.png", + "caption": "Figure 3: The shaft dynamometer for the human-powered submarine as it was demonstrated at the conclusion of the capstone design course, Spring Quarter 1998.", + "texts": [ + " 2, a pulse timing technique was used to infer both angular velocity and shaft torque (based on shaft twist) sixteen times per revolution of the shaft. From these measurements, the other output variables of interest were computed. To converge on an optimal design, the students used a model relating the mechanical and computational design parameters to the expected accuracy and bandwidth of the dynamometer. The final design included a scheme that used an automatic lookup table to compensate for the manufacturing tolerances in the magnetic encoders. As shown in Fig. 3, the laboratory prototype was demonstrated on a lathe equipped with a F\u2019rony brake, where the shaft could be subjected to constant speeds and reproducible torques. One year later, a second mechatronics project team built on the previous effort to complete a fully autonomous and submersible prototype, including features such as firmware code, battery power, data management, and capabilities for uploading stored data to a PC. 1. This project was suggested by one of the students participating in the mechatronics thread" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000149_00423119608969196-Figure10-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000149_00423119608969196-Figure10-1.png", + "caption": "Figure 10: The i th Contact Patch", + "texts": [ + "9 the relation between twisting angle 0 and the direction of velocity a t the leading edge and the trailing edge can be expressed as: - sin y -1 r From this relations the twisting angle P can be calculated as: sin y 0 = 2 tand1 ( F ~ ) Local twisting angle is proportional to the distance from the leading edge. Therefore the local twisting angle/& at i th small patch can be expressed as: DISCRETE BRUSH TIRE MODEL 209 Local twisting moment is calculated considering the lo~lgitudinal force distribution at the it11 contact patch (Fig.10). C.w!Az - where G,,, = is the twisting stiffness of the i th contact patch and C, is the longitudinal stiffness of tread rubber per unit square. The moment Mzr can be calculated by summing up the local moment from the leading edge to the trailing edge. The moment around the vertical axis can be calculated by summing up the moment Mzl(without tread base deformation) and the moment M,,. 4'.3 Calculated Results The lateral forceFy and the moment M, around the vertical axis versus a are calculated under the camber angle 0,20 and 40[deg]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003612_jmes_jour_1980_022_017_02-Figure4-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003612_jmes_jour_1980_022_017_02-Figure4-1.png", + "caption": "Fig. 4. Experimental rig", + "texts": [], + "surrounding_texts": [ + "Lubricant film thickness was obtained using an optical interference technique. The principal features of the experimental apparatus are shown in Figs 3 and 4. A glass plate covered with a layer of transparent silicone rubber was mounted on the stage of a microscope. A polished steel cylinder was released from a specified height and was allowed to fall through oil onto the rubber layer. Monochromatic or white light was shone 6 3 5 2 a 4 1 h L . N .. 0 v $ 3 Ill k N . N 2 M 3 solid (broken lines) ; v=0.5 from below through the microscope objective and reflected back to the eyepiece, display screen or 35 mm camera. The interference fringes formed from the thin oil film between the upper cylinder and the rubber layer were used to determine oil film thickness during the approach of the two surfaces. In order to produce clear, well defined fringes, the rubber layer was coated with a thin semi-reflecting layer of aluminium. Journal Mechanical Engineering Science 0 IMechE 1980 Vol22 No 2 1980 at TUFTS UNIV on November 30, 2015jms.sagepub.comDownloaded from 68 C. J. CUDWORTH AND J. F. MYKURA A series of normal approach tests, using different loads and different thicknesses of elastic layer, was carried out in order to obtain data on the variation of oil film thickness with time and to compare this with the predictions of the theory. A stop watch was used to measure time after release, and photographs of the interference fringes were taken by hand at successive intervals of time on 35 mm film. Subsequently, using the sequence of photographs, the order of the central fringe at a given time was evaluated and the film thickness calculated. It should be noted that the initial clearances used in the experiments were of the order of 0.1 mm and in general the maximum film thickness recorded using the interferometric technique was less than 1 pm. It was therefore only the closing stages of approach, i.e., less than 1 per cent of the initial clearance, that were studied in this series of experiments." + ] + }, + { + "image_filename": "designv11_65_0002395_pesc.1994.349727-Figure5-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002395_pesc.1994.349727-Figure5-1.png", + "caption": "Fig. 5 : Selection of flux vector, with @ inside region I , when Vsd.Vsq > 0", + "texts": [ + " 2, the selection of voltage vector depends only on the two hysteresis outputs, which are related to specific values of v, and vq (for example, v, > 0 if we want to rise flux amplitude, and > 0 for a fast increase of torque in the positive direction). Let's consider the hysteresis outputs as unitary values, which signs are the desired signs for v, and v q . Now let's examine vector selection if stator flux vector is inside region I. Vector B will be selected if v, and vsq are decided to be positive (fig. 5) . However, the real values of vu and vq depend on the actual position of vector flux, as it's shown in figs. 5 and 6. Fig. 6 indicates also that the region I could be completely shifted by 30\", up to the new Vdmin = V,.sin( Ba ) Vdmax Vqmin Vqmax = V,,.cos( Ba ) = V,, sin( Ba + 60\" ) = V,,.cos( Ba + 60\" ) ( V,, is the amplitude of inverter voltage vectors ) 0 -40 -h A ' e'D 4'0 $0 - - L , , . .--.. -. - . { I-}....-\u20ac1 +--..-e)-- --1_--1 \u20ac- Previous equations indicate that the rise of Ba increases Vdlnh and Vdmax, and decreases Vqmin and Vqmax" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001030_ac9963300397-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001030_ac9963300397-Figure1-1.png", + "caption": "Fig. 1 Measuring cell made with a C02 sensing electrode and stirring bar reactor using sorghum seeds as biocatalyst. ( a ) Global view and (b) profile view of the reactor.", + "texts": [ + " The standard solutions of oxalic acid were prepared by dissolving an appropriate amount in the electrolyte solution that was used to carry out the measurements. In pH effect studies the oxalic acid solutions were prepared in buffers with adjusted pH. The buffers used were: for pH from 1.0 to 1.8, KCI-HC1; pH 2.5, glycine-HCI and pH from 3.1 to 5.2, succinic acidNaOH. The sorghum seeds were washed with distilled water, cut into four pieces and immersed in a 0.10 mol 1-1 KCI solution. This material was used to construct the reactor. The device, constructed for the measurements, illustrated in Fig. 1, is similar to that used by Guilbault.15 An OP-9353 Radelkis electrode sensitive for C02 was used in this system. In the stirring bar type enzymic reactor, 10 seeds of sorghum, previously cut and soaked with KC1 were covered with Nylon gauze, over which the electrode was fixed. This system was introduced into a thermostated glass cell to carry out the measurements. The volume of electrolyte or buffer used in the cell was 5.0 ml, previously saturated with oxygen. The device was conditioned until a constant potential was attained and then aliquots of oxalic acid samples were added using a Finnpipette micropipette" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002623_a:1015265514820-Figure6-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002623_a:1015265514820-Figure6-1.png", + "caption": "Figure 6. The robot model used in simulation experiments. The sensors P1 and P2 have a sensing angle of 90\u25e6. Each of the ring sensors has a sensing angle of 45\u25e6.", + "texts": [ + " These steps are executed by the corresponding controller subsystems. The algorithm terminates when the reactor determines that the robot has reached its destination. We have implemented this architecture as a C++ program running on the industrial PC of the robot. The resulting code is very efficient, resulting in a \u22483 ms execution time for the controller algorithm loop. The effectiveness of the navigation method was tested in a series of simulation experiments. All these experiments used the robot model shown in Fig. 6. The robot was equipped with a ring of eight proximity sensors located at 45\u25e6 intervals and two proximity sensors on the front. The sensor ring was used for Qmap construction and positioning. The two sensors at the front (P1 and P2) were used for obstacle avoidance. P1 and P2 had a smaller range than the ring sensors. The simulator assumed that each obstacle in the environment map had a stimulus region associated with it. This was a constant set by the user, which reflected the range of each proximity sensor" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002685_robot.1999.772524-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002685_robot.1999.772524-Figure1-1.png", + "caption": "Figure 1: Two serial manipulators, with different geometry but the same instantaneous twist and wrench spaces.", + "texts": [ + " The physical interpretation of G follows straightforwardly from the definition of a dual basis: the i th column of G corresponds to the wrench on the end-effector that requires a torque at the i th joint but no torques at the other joints, in order to keep the manipulator in static equilibrium (and neglecting gravity!), [2]. In summary: the i th columns in J and G are each other\u2019s dual, given the geometry of the serial manipulator. The reasoning in the previous paragraphs relies on the position and orientation of the joint axes in space; it does not, however, rely on how these axes are interconnected! For example, both manipulators in Fig. 1 have the same Jacobian matrices. This remark yields the following definition (of little practical value?) , which generates an equivalence relation in the set of all possible serial manipulators: Definition 1 Two serial manipulators are instantaneously dual if they have the same type of joints at the same position and orientation in space. It is clear that manipulators, that are not only instantaneously dual, but dual for all possible joint positions, must, kinematically speaking, be functional copies of each other" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002911_s00348-002-0578-5-Figure3-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002911_s00348-002-0578-5-Figure3-1.png", + "caption": "Fig. 3. Experimental arrangement: liquid bridge cell (A), CCD cameras (B and C), and illumination system (D)", + "texts": [ + " When a>0, the interface deformation at the stability limit caused by both effects is similar (in both cases the liquid column necks at the top half and bulges at the bottom one). Since both effects add up, the stable slenderness decreases (curve A\u2013B). If both branches corresponding to the same value of B are considered, obviously the stable region is that located between both curves, whereas points outside this region represent unstable liquid bridges. 3 Experimental setup The experimental setup used to perform the experiments (Fig. 3) consists of the following main elements: liquid bridge cell (A), CCD cameras (B and C), illumination system (D), image recording system (not shown in the sketch) and associated software. The liquid bridge cell consists of a three-axis table; the displacement along each of the different axes is controlled through micrometric screws. The accuracy in the displacement along each of the three axes is \u00b10.01 mm. The liquid bridge is formed between the end faces of two solid cylindrical supports. The bottom support is mounted vertically on the horizontal platform of the three-axis table, and it can be displaced both along the x-axis and the y-axis" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002629_iros.2001.973411-Figure8-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002629_iros.2001.973411-Figure8-1.png", + "caption": "Figure 8: Kick behavior", + "texts": [ + " The main purpose is to realize behavior evolution, thus the environment is simplified and necessary information is supposed to be given, the goal position is fixed and known, for example. It would be significant to learn Swhg or Step by optimally combining PTP motions. But these motions are already realized and it is time consuming in real robot experiments. Therefore Step and Swkg are used as element motions to compose Appmecn and Kick(Figure 7). As an optimization technique the Simplex method for non-linear functions is used. First Kick is optimized for several typical ball positions psK = (ZB, w) in a given region OK (Figure 8). Kick is formulated as a Swkg motion along a predefined foot tip trajectory TK(aK) specified by a parameter vector aK E Rs, and then aK is optimized. Kick is evaluated by - *(aK) 11 . The resultant o p timal parameters for the case po = (3000,0), OK = [180,300] x [-400, -2801, J1 = Ja = 4 , are stored as a knowledge array VK = {wgai1} E R(313, Ja, J I ) as shown in Figure 9. With this empirical knowledge, VK, the robot can adaptively realbe sub-optimal Kick for arbitrary bK E OK even for inexperienced ball positions" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002119_1097-4563(200012)17:12<643::aid-rob1>3.0.co;2-8-Figure2-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002119_1097-4563(200012)17:12<643::aid-rob1>3.0.co;2-8-Figure2-1.png", + "caption": "Figure 2. Angular rotation of 2 for a planar surface and a sphere of radius 1.", + "texts": [ + " Calculate the curvature form please see the Appendix for these geometric definitions and .concepts based on Montana\u2019s kinematic equation of relative rolling motion, using the recorded values of , , and p.x y \u017d7. Estimate the radii of curvature at the contact point and in directions that correspond to the .x and y axes of the probe frame of the object from the curvature form. The and parameters are calculated byx y measuring the net rotation with respect to the x and y axes, respectively, of the initial contact frame on the probe. Figure 2 shows a 90 rotation of the finger \u017d . on planar and spherical surfaces. In the planar case, the surface normal points in the same direction in both the initial and final positions. In the spherical case, the direction of the surface normal has rotated 45 since the sphere has the same radius as the fingertip. In the case of convex surfaces, the angular rotation of the surface normal will be less than that for a planar surface for the same . Intuitively speaking, the curvature estimation is based on the difference in the change in the contact \u017dpoint for a known angular rotation of the finger see " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001102_tt.3020030402-Figure10-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001102_tt.3020030402-Figure10-1.png", + "caption": "Figure 10 Schematic view of rotary stick-slip tester", + "texts": [ + " RepeatabiZity of the stick-slip number The repeatability of the test was checked by performing 5 tests on 4 different lubricants. All tests were conducted using dead weights as normal load. The stick-slip numbers for products A and B were determined by the leaf springs; the stick-slip numbers for products C and F were determined using the load cell. The test results are presented in Table 2. From this table, it follows that the repeatability of the stick-slip tests is fairly good. ROTARY STICKSLlP TESTER: Test rig A schematic view of the rotary stick-slip tester is shown in Figure 10. The test rig consists basically of a rotating bath (l), driven by a frequency controlled AC engine, in which a ring (2) is mounted, and against which two shoes (3), which are connected to a fixed arm (4), are pressed. The rotation of the bath causes sliding of the ring against the shoes. The normal load is transferred to the shoes by means of the arms and levers (5) on which dead weights (6) are placed. The very small displacement of the arm caused by the friction force between the shoes and the rotating ring is measured by a load cell (7) which connects the arm (4) to the frame of the test rig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002212_1350650011543484-Figure9-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002212_1350650011543484-Figure9-1.png", + "caption": "Fig. 9 Traction force impact apparatus", + "texts": [ + " It was further concluded that sound measurements could be achieved only by decoupling the measurement of normal and the horizontal forces and their bending moments. A separate apparatus was therefore devised for measuring the traction forces based upon a single-component piezoelectric force transducer. J04300 # IMechE 2001 Proc Instn Mech Engrs Vol 215 Part J at UQ Library on June 17, 2015pij.sagepub.comDownloaded from Since the normal force component during impact could be measured accurately with the three-component piezoelectric apparatus, as demonstrated in Section 3, a new apparatus shown in Fig. 9 was designed to measure only the traction force. The traction force was measured with a single-component piezoelectric transducer (Kistler 9031A) whose signal was amplified and displayed as described earlier. The load washer was positioned in line with the dropping ball, i.e. the centre-line of the transducer coincided with the impact position. The tungsten carbide impact plate was mounted on six, vertically positioned aluminium alloy blades. These vertical blades were stiff enough to sustain the high normal impact force without buckling, but very flexible compared with the frictional force transducer" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002206_robot.1996.506166-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002206_robot.1996.506166-Figure1-1.png", + "caption": "Figure 1: Process states in insertion of hose into plug", + "texts": [ + " Humans then provide an alternating motion such as a push-pull motion or a twisting motion to a hose rather than a straight motion so that the hose nozzle can be inserted into the plug completely without excess deformation of the hose. Investigating the above insertion process with respect to the contact between a hose and a plug, it turns out that the successful human insertion process consists of three states, that is, (a) approach state, (b) contact state, and (c) insertion state, as shown in Figure 1. The hose is apart from the plug without the contact between them and the hose approaches toward the plug during the approach state. The hose is in contact with the plug but the nozzle of the hose is not inserted into the plug completely in the contact state. In the insertion state, the nozzle of the hose is completely inserted into the plug. From the above discussion, we find that humans can adapt their motion control to the process states. Note that this adaptation requires the recognition of current process states or the recognition of transitions among process states" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003643_s00466-004-0614-9-Figure6-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003643_s00466-004-0614-9-Figure6-1.png", + "caption": "Fig. 6 Geometry of the membrane", + "texts": [ + " Continuing the backward movement the membrane switch to a regularly wrinkled state again. The wrinkling direction now is completely different from a of the forward directed process. 4.2 Twisted annular membrane This example earlier was considered in literature for elastic membranes, see e.g., Reissner [8], Roddeman [10] or Schoop, Taenzer and Hornig [12]. Here additionally plastic effects are taken into account. Numerical simulation as well as experimental investigations were made. A annular aluminum sheet, fixed at the outer edge, rotates at the inner edge by an angle # (see Fig. 6). The inner radius ri and outer radius ro remain constant during this deformation process. The twist is stopped at # \u00bc 1:82 \u00bc 0:03176 rad and reversed till # \u00bc 0 . The data of the membrane are: E-modulus E \u00bc 70000 N mm2 Poisson ratio v \u00bc 0:3 yield strain y0 \u00bc 0:002 (isotropic hardening) membrane thickness h \u00bc 0:01 mm inner radius ri \u00bc 45 mm outer radius ro \u00bc 125 mm FE-simulation: The elastic-plastic material model from Besdo in conjunction with the isotropic hardening law Eq. (37) will be deployed" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002699_iemdc.2003.1210679-Figure3-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002699_iemdc.2003.1210679-Figure3-1.png", + "caption": "Fig. 3. Several roton: with davble cage A. Spiral sheet mton E C", + "texts": [], + "surrounding_texts": [ + "nomn's TYPE\nI INTRODUCTION\nThe rotors of conventional asynchronous motors are formed by magnetic sheets packed above the shaft of the machine. The rotating magnetic field created by the stator, makes some elecbomagnetic sources parallel to the shaft and so that upright to the rotor sheets. Those currents can not go through if the sheets stay electrically isolated between them, being necessary the intervention of the conventional squirrel cage rings to close the electric circuit thus the rotor currents can be circulated.\nTypical configurations of three-phase aspchronous motors are described in the following points:\n1) Siigle cage winding with hi& resistance and minimal indncbmce.\n2) Deep slot cage, with low resistance and progmsive m c e with the slip.\n3) Low resistance and low reactance slot. 4) Finally, there are additional configurations of double and\ntriple cage combining characteristlcs of the previous\nMASS meum KINETIC n0mn-s N (LF\") ENERGY DIMENSIONS\ndispositions, with different resistance and reactance values for each cage, obtaining torque-speed curves for several applications. Normally the double disposition, interior cage of minimal resistance and high reactance for slips lesser than the one relative to rated toque, is combined with an exterior cage of high resistance and minimal reactance that works for bigger slips rotor stopped inclusive.\n07803-7817-21031$17.03 02033 IEEE 1688", + "Ill THREE-PHASE ASYNCHRONOUS TORQUE-MOTOR angular shape, in order to make a difference between the both zones, one where active currents go through, and the other\nThe so called \"torque-motor\", are three-phase which is used to receive the possible returning currents (A retnnhg currents proposal). asynchronous motors with the stator built in a classical way,\nhut its rotor is different fiom traditional ones, because it is formed by a hard sweet i nn cylinder (with small hysteresis l-, raT FA1.-, cycle). With this kind of rotor we get a highcr useful section for the fluxes circulation of the nux. duc 10 lack of slots s a m g a 5% adding factor by isolating sheets. One of the irfr If /3- /3- E /r drawbacks of that kind of motor is thefact that the lines of magnetic field go through its core reaching considerable depths, all that affects to the existence of emf at quite depth and therefore thev are weaker.\nFig 2. S e \" afa\"thnophax aryochranovs \"orque-mauo\"\nThe immediate effect of a certain nurent circulating to a higher depth is its minor participation in the generation of torque, due to the following causes: 1) Minor distaace between the curtents and the shat? of the\nmotor. 2) The higher reactonce results in lower power factor\nalthough the losses in the copper have the same valuer as if they were currents that circulate by the periphery. In figure 2, that pbenomenon is illustrated, showing thnt the. deeper currents are weaker, and have also some delay in the direction of the displacement.\nIV MOTOR WlTH SPIRAL SHEET ROTOR\nFig. 4.. Plaio developed section ofthe sheds dispaition\nIn spite of this, the returning currents can he established in\nOption A: Through short-circuit rin5. With that kind of conslaction the only rotor resistance that must he considered, is the one corresponding to the outside of the iron sheets that form the rotor. This is due to the shortcircuit ring can theoretically he built in a big section, beiig its resistance totally despised comparing with the resistance offered by the superficial sheets' layers. This solution is not suitable for manufactnring, however it gives the best results.\ntwo manners: -\nrings. Forming a rotor with spiral shape sheets, distributed in a\nradial disposition around the shaft, it is possible to generate magnetic fields stay more in the rotor's periphery, inducing peripheral emf, and nurents along the same sheets, that are only active in their periphery.\nOption B: This second option, consists of a part of the same sheets being a pathmy for back currents that do not generate toque. In this model, the zone of active currents that generate torque are placed in the sheets' periphery. Becanse of these ones are affeded by the magnetic-field, leaving enough sheet section as a r e m way, they do not generate torque currents, without increasing, the resistince of the rotor winding, beiig the last one of a very bigh section. This second option is considerably of a simpler construction, and a good adhesion is enough to consolidate mechanically the sheets to the machine's & ~~ ~ ~ ~~ ~\nV PROTOTYPE DESCRIPTION The peripheral currents of this rotor have more section to\ncucdate, compared with a normal cage rotor's Current. The\ndisposition of the sheets. Instead of being shaped in an\nSeveral types of rotor have been CO-cted, each one is figure 4, shows a developed plain representation of the formed by a of magnetic sheets d i a l y asposed with", + "the head directly above the shall. Those sheets wmpletely wver the outside cylindrical surface of the shaft.\nSince the sheets' density is constant, in order to eliminate kee space that would result of increasing the radial distance, the sheets have been enrolled above the others, as if the rotor would be formed by a group of sheets spinning as it can be seen in the figure 6.\nBy this way, the generated flux in the stator windings, that i s the rotating magnetic field, falls in an inclined way into the sheets, producing some eddy currents which circulate in presence of the mentioned flux, generate =me antagonistic toques, m e outside torque is bigger than the inside torque, hecause the radial distance is much higher, which results in a net movement of rotation of the same value to the difference between the mentioned torques, as we can see in the next f i P .\nVI CONSTRUCTION\nIn order to construct the proposed prototype rotor, a radically different process from the construction of a conventional squirrel cage rotor, is followed.\nIt has been indicated that in those newly rotor, the sheets don't have any slot, so that in order to have them perfectly joined, and not leave air spaces between them, its shape must have a perfect demte profile. This profile is reached by the shaping with a tool or matrix to deform every piece or sheet, giving the precise shape that is described as follows.\nIf we consider a rotor of interior radio R, formed by n sheets of a thickness e, being this thickness very small compared to R, and with an outside radio RE, we will have the trigonometric ratios in the figure 8\nA. Calculus of the length of the sheer in function the interior radio R. and the exterior radio RE.\nIf we make equal the perimeter of the sheets' growth in the inside radio R, and outside radio RE; we can present an equation system such as the following:\nen cos a 2n.R = e n 2.11.~ = -\nR P If we divide both equations we can obtain: - = cosa (2)\nBesides, from the differential triangle, knowing that both angles \"a\" ate the same, because of the perpendicular sides, we can write:\ndl = - dp sothat ,=e (3) cos a R\nIf we integrate now, we get:\n(4) (R: - R ~ )\nRE\n[L]=[$] sothat L = - R 2.R\nThis is an exnression that relates the l e n d of everv sheet with I the interioiradio R and the exterioroniRE." + ] + }, + { + "image_filename": "designv11_65_0003924_0022-4898(71)90023-1-Figure4-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003924_0022-4898(71)90023-1-Figure4-1.png", + "caption": "FIG. 4. Model drag force vs. model depth for varying depths of model 1 propelled at 3.27 ft/min (logarithmic scale).", + "texts": [ + " The output of this gauge was fed into a millivoltmeter or a sensitive X - Y plotter. To achieve uniformity before each run, the sand in the tank, Hatford Loam (particle density -- 2.67 bulk density = 1-55, and the angle of internal friction measured by the triaxial test, ? -- 30.5 \u00b0) was vibrated with a concrete vibrator and levelled with a template to a constant height. I f density and depth were the force determining parameters it would be expected from dimensional arguments that the force would vary with the depth of the model cubed. Figure 4, a logarithmic plot of force vs. depth for five depths (12 in., 15 in., 18 in., 21 in. and 24 in.) shows that this variation is not obeyed precisely. The maximum slope 2.8 is between the 12 in. and 15 in. points; the overall slope is 2.3. Error bars, which represent the standard deviation of all the tests for any one parameter, show that the change in slope cannot be regarded as significant in detail. However, the overall slope shows that at some point the controlling factors of the force are not bulk density and depth but pressure and area" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003391_cp:20040369-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003391_cp:20040369-Figure1-1.png", + "caption": "Fig. 1 Finite-element flux-plot of the tested motor, which is a 2-pole line-start split-phase motor having orthogonal main and auxiliary windings. ' . .", + "texts": [ + " Introduction Modem permanent-magnet synchronous brushless machines often have magnetic circuits in which the patterns of saturation are complex and vary with the position of the rotor, Fig. 1. The classical theory of operation relies on such assumptions as the sinusoidal distribution of windings and sinusoidal variation of inductance with rotor position, and has no natural means of representing the strong but localised nonlinear effects that arise in different operating states, [2]. In almost all the literature that deals with the modelling ofthese machines, the magnet flux (represented by the open-circuit voltage E ) is treated as constant, while saturation is represented by cunent-dependent inductances", + " For sinewound machines a simplified theory can be used, as follows, to express the magnetization curves in a form that does not require multi-dimensional curve-fitting, and is therefore suitable for fast calculation. It relies on the definition of a current space vector, which is possible only if the windings are sine-wound. The two phase currents and the effects of their ampere-conductor distributions can then he represented by the magnitude and spatial orientation of the current space vector. 4 , Y lVSl t D ! - . , --- ..... CURRENT \\ 4 The coils d and q are stationay with respect to the rotor, as shown in Fig. 1 1 , and eqns. (1) are in the rotor reference frame. It is common to assume that yo is constant while the synchronous inductances L, and L, vary with current, hut a more general form is Fig, Solution of flux-linkagei current relationship Y , = Yd(id3iq); Y , = Yq(id>iq). (2) - d When multiplied by the frequencyjo, eqn. (1) gives v, = joy, = E + XdId; ( 3 ) vd = jw, = xqIq in which Xd and X, vary with current hut E is constant. On the other hand, the \"circuit\" counterpart of eqn. (2) does not discriminate between E and XdId, but uses V, directly", + " increases the level of saturation 678 It makes no difference whether i , is positive or negative, &cause the addition of current i, will only increase the level of saturation above that which is obtained with any value of id. Therefore the line q in they, curves in Fig. 11 is not along EF but slightly below it. In general current is flowing in both the d and q coils.and we can define a current space-vector i = ie\" = id + j i g , .(4) , This makes it possible to represent the variation ofy, with both id and iq as shown in the y, curves in Fig. 1 I , in which the current magnitude i is plotted along thex-axis and the different curves are obtained with different values of 8. We have already examined four ofthese curves, for 0 = 0, 180\", and 190\". q-mis ~ The q-axis is similar but simpler because there is no magnet flux-linkage. Therefore the yq curves are symmetrical for positive and negative currents i,, which are shown by the curves labelled q and -q. With current in the d-axis q d I = 0 91800 we expect no variation iny, and therefore the curves fqr id both lie along OJ" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002529_00037-1-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002529_00037-1-Figure1-1.png", + "caption": "Figure 1 Tensile test specimen configurations.", + "texts": [ + " Even if longer grip faces are available, maintaining the uniformity of the clamping pressure then becomes problematic. If gripping problems usually the final solution persist is to reduce the thickness of the material being tested. The area of bonded adhesive is maintained and the total cross-sectional area in the gage section is reduced. In fact, axially-loaded unidirectional composites are typically tested in thin composite form. For example, ASTM Standard D 3039 (1995) recommends a thickness of only 1mm. Figure 1(a) shows the tensile specimen overall configuration recommended in ASTM Standard D 3039. An untabbed, dog-boned specimen such as recommended in ASTM Standard D 638 (1996) is shown in Figure 1(b). When performing an axial tension test, it is often desired to determine both the axial modulus and the major Poisson's ratio. Thus strains must be measured in both the axial and transverse directions. A 0/90 two-element biaxial strain gage is very convenient for this purpose. A biaxial strain-gaged extensometer can also be used. However, such extensometers are even more expensive than the single-axis extensometers previously discussed (MTS Systems Corporation, 1997; Instron Corporation, 1994)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002813_0022-1902(81)80577-5-Figure2-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002813_0022-1902(81)80577-5-Figure2-1.png", + "caption": "Fig. 2. Distribution of the species as ac, vs - log [H+]. Solvent mixture (III) Cc, = 4.00; Cmal~ = 4.00; C#y = 4.00; Cr~ = 12.00 (raM). (a) Theoretical distribution, without taking into account mixed complexes; (b) real distribution.", + "texts": [ + "' and therefore log k = log K ' - A w log ([H20]s/[S]). (a) log k for protonation constants of malonic acid; (b) log k for stability constants of Cu(lI)-malonic acid complex formation: (c) log k for protonation constants of glycine; (d) log k for stability constants of Cu(II)-glycine complex formation. Ternary systems. The values of protonation and Cu(II)-complex formation stability constants suggest that binary complexes of malonate and glycinate would exist in somewhat different ranges of -log [H\u00f7]. Figure 2(a) shows a theoretical species distribution (ac, = fraction of total copper present in each complex) vs -log [H+]: this distribution was obtained on the hypothesis that in a solution containing Cu(II), malonic acid and glycine no mixed species were formed, and only binary complexes exist Figure 2(b) shows the species distribution calculated for the same experimental conditions as that of Fig. 2(a), but taking into account the presence of mixed complexes The equilibrium concentration of Cu 2+ is almost unaffected by the formation of the mixed complex, while it seems likely that the ternary complex is formed by addition of glycinate to Cu(mala) It is possible to investigate whether the solvent composition affects the relation between the stability constants of parent binary and ternary complexes. A comparison can be made by the relationship [8]: Cumala Cu A log K = log K(c,ma,a)tgly) - log Kc,g" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002346_robot.1998.677406-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002346_robot.1998.677406-Figure1-1.png", + "caption": "Figure 1. Quadruped model with decoupled tree structure.", + "texts": [ + "00 0 1998 IEEE 1701 system the constant calculations and the construction of topology data should be accomplished before the iterative stages of the simulation. The quadruped model using the decoupled tree structure approach [7], that decouples the closed chain system into a tree-structured open chain system through the introduction of spring/damper systems at the ground, can be interpreted as a multibody system with several possibilities of topologies. The modified velocity transformation method is more effective for the more parallel configured systems[l2]. Thus we choose the most parallel-configured system as shown in Fig. 1. The body of the quadruped is joined to the ground by free joint and each leg is composed of three links with revolute joints. We label the ground as body 0 and number the other bodies in the ascending progression until all bodies are numbered. For convenience' sake we assume that body j is the lower adjacent body of body i and call by joint i the joint between body i and bodyj. At each joint, the joint coordinates as many as its DOF are assigned to describe the relative motion between two joined bodies. They are numbered in a proper order as shown in Fig. 1, which determine the form of the resultant mass matrix. Let us denote by Doeand qi the DOF and the joint coordinate vector of joint i. The body tree array r showing at column i the topology of the chain from body i to the ground and the body exponent array H showing the number of bodies in the chain are constructed as T(1,i) = i T(2,i) = j (1) T ( k , i ) = T(2 ,T(k - 1,i)) and H(i) = k - 1 when T ( k , i ) becomeszero first (2) The coordinate tree array having at column i the coordinate numbers of the joint coordinates related to the kinematic expression of body i and the coordinate exponent array H, showing the number of the coordinates are constructed as follows: } (3) coordinate number of qi 4 q ( ", + " The number of computations for an iteration of the simulation is proportional to N and F3 . It means that the proposed method is an O ( N ) algorithm for parallel structures and an o ( N ' ) algorithm for serial1 structures. Therefore the proposed method may be most profitable for the quadruped model with decoupled tree structure because it has been demonstrated [6] that the best O ( N 3 ) algorithms are faster than the best O ( N ) algorithms for N < 10. For the quadruped model with N:=4 and F = 3 shown in Fig. 1, 6894 times of multiplication and 5454 times of addition (6894m+5454a) are needed for an iteration of the simulation. I[f parallel computatiion is applied to each leg, the number of computations can be reduced to 2064m+1608a and only 29.7% of the execution time will be needed to accomplish the simulation as shown in Table. 3. From the fact we can expect that the real-time simulation can be achieved with the application of parallel computation in the case of At =0.0005s. 5 Conclusions We have applied the dynamic formulation using the modified velocity transformations and the modified Cholesky decomposition for DBE5D matrix to the real-time simulation of the quadruped" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003206_2003-01-1712-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003206_2003-01-1712-Figure1-1.png", + "caption": "Figure 1. Front Suspension Lower Control Arm", + "texts": [ + " However, based on discussions with engineers who had performed nibble evaluations in the past on this and other vehicle programs, it was decided to use a version of a company sanctioned \"nibble test method\", This test method consisted of applying imbalance weights to the road wheels in various configurations, based on a NVH CAE recommended Best Practice Procedure. The specifics of the amount of weight and point of application is described under the Methods section. Front suspension Lower Control Arm assemblies were built up for the nibble studies . Table 1 contains information on the front bushings of the arm (shown in Figure 1) which were used in the evaluations. All of the bushings were manufactured in the U.S. , except for those used in iteration 4, which were manufactured in Germany. The bushings used in trials 1-4 were manufactured with rubber compounds which yielded various levels of damping. The varied damping levels resulted in varied dynamic rates. In general, higher rubber damping, results in a higher dynamic rate, relative to the static rate. Copyright \u00a9 2003 SAE International Figure 2 \u2013 Plan view of LCA Bushing The vehicle was prepared for testing as follows: Tire pressures set to the nominal specification; wheel assemblies balanced to within 4 grams residual imbalance; wheel alignments checked and adjusted, vehicle ride heights checked" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003257_s1474-6670(17)35276-x-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003257_s1474-6670(17)35276-x-Figure1-1.png", + "caption": "Fig. 1. Furuta pendulum.", + "texts": [ + "eywords: Speed Gradient Method. Furuta pendulum. Swing up. Energy-based Design. Non-linear Oscillations. 1. INTRODUCTION The inverted pendulum is a very simple device that displays very interesting behaviour modes, which have attracted the attention of many con trol researchers. In this paper, the interest is f~ cused on the rotating type of inverted pendulum. A schematic representation of the system is shown in Fig. 1. As it is displayed in this figure, () denotes the angle of the pendulum with the upright verti cal and ip denotes the angle of the rotor arm. The rotating pendulum is also known as the Furuta pendulum, and has been studied by many authors such as Furuta himself and Astrom (Wiklund et al., 1993), and others (Bloch et al., 1999). The inverted pendulum gives rise to many interesting control problems. It is a nonlinear underactuated mechanical system that is unstable at the desired position. Furthermore, the actuator limitations produce very complex and interesting behaviours that deserve careful analysis", + " It was shown by A. Shiriaev (Fradkov et al., 1997) that A4 can be weakened at the cost of strengthening A2. namely, let A2 be complemented by the conservativity-like con dition A2'. h(x)T Lfh(x) :::; 0 for all x E no and h(x)T Lfh(x) = 0 if x E no and Lgh(x) = O. Let A4 be replaced by Shiriaev's rank condition dimS(x) ~ 1 '=>;R h7 32 The I/O polynomial R 2 k h7 is given as R X20 n 1 an cosn h7 bn cosn\u00ff1 h7 sin h7 a0 X40 n 0 cntn 7 0; 33 where t7 tan h7=2 : Since tangent half-angle substitutions are repeatedly used, Theorem 3 was used to verify that R is devoid of any extraneous solutions. Hence, with link 1 grounded and link 2 as the input link, link 7 (as well as the mechanism of Fig. 5) has 40 possible assembly con\u00aegurations. For h2 69:0\u00b0 and the numerical data given below, the coe cients as well as the solutions of the 40th degree I/O polynomial are given in Table 5. Consider next the case when link 7 is the input link. The loop-closure equations can be expressed, in condensed form, as follows: Loop1 6; 10; 8; 7 ! f1 h6; h8; ~h10 ! F1 h6; h8 ; Loop2 6; 5; 4; 7 ! f2 h4; ~h5; h6 ! F2 h4; h6 ; Loop3 7; 8; 9; 2 ! f3 h2; h8; ~h9 ! F3 h2; h8 ; Loop4 7; 4; 3; 2 ! f4 h2; ~h3; h4 ! F4 h2; h4 : 34 From the reduced loop-closure equations F1\u00b1F4; using tangent half-angle identities and Theorem 1, the variables h6; h8 and h4 are eliminated in succession as outlined below: F1 ~h6; h8 F2 h4; ~h6 F12 h4; ~h8 F3 h2; ~h8 9=;F12;3 h2; ~h4 F4 h2; ~h4 9>=>;R h2 35 r1 8:50 r2 5:00 r3 4:45 r4 3:00 r5 2:80 r6 3:00 r7 2:90 r8 2:70 r9 2:50 r10 4:90 r1a 4:60 r2a 2:75 r4a 2:20 r6a 5:70 r7a 4:90 r8a 1:75 a 13:5\u00b0 b 30:0\u00b0 c 97:0\u00b0 / 17:5\u00b0 d 19:0\u00b0 g 107:0\u00b0 Here R 2 k h2 is expressed as R X15 n 1 an cosn h2 bn cosn\u00ff1 h2 sin h2 a0 X30 n 0 cntn 2 0; 36 where t2 tan h2=2 : Since tangent half-angle substitutions are used repeatedly, Theorem 3 was used to verify that R is devoid of any extraneous solutions" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001810_0094-114x(84)90051-x-Figure2-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001810_0094-114x(84)90051-x-Figure2-1.png", + "caption": "Fig. 2. Chain load distribution", + "texts": [ + " MARSm!K steel roller chain; one worm gear reducer, one X - Y recorder, one operational amplifier, one + 15 VDC/100 mA constant power supply, one 10-turn potentiometer, a fixture with a magnetic base to hold the potentiometer casing, an aluminum support plate, a specially designed input crank handle, and a standard set of laboratory weights. The four strain gages were used to measure the roller chain load distribution. The strain gages were mounted on the link plates of one pin link, as shown in Fig. 2. The strain gages were axially aligned with the direction of the highest principal stress in the link plate. The strain gaged pin link will be called the a c t i v e l ink. test machine wiring diagram. The single-reduction worm gear reducer was used in rotating the sprocket and in accurately measuring the angle of rotation of the sprocket. The gear-reducer output shaft has a keyway for mounting the test sprocket. The input shaft has a keyway for mounting a manually rotated handle. The X - Y recorder was used to record the experimental results of the roller chain load distribution while the tests were being performed. The X-axis terminals were connected to a 10-turn potentiometer attached to the sprocket-mounting shaft of the reducer. The 10-turn potentiometer was connected to a -+ 15 V, d.c. power supply as shown in Fig. 2. The X-axis output represents the angle of rotation of the sprocket starting from a zero degree position, which will be denoted by to. The zero degree position, to, is the position where the active link meshes fully with the sprocket for the first time as the sprocket rotates. The 180 o position, 480, is the position where the active link starts to disengage from the sprocket. The Y-circuit terminals were connected to the strain gage leads through an operational amplifier, OP-AMP, as shown in Fig. 2. The results recorded in the Y direction are calibrated to represent the actual force on the link which is proportional to the average of the four strain gage readings. Before each test, the test sprocket and chain were cleaned thoroughly. The steel sprocket had a built in hub and was placed on the reducer output shaft and secured with a square key. In conducting these experiments, it was important to guard against chain jumping, which is the condition where the chain rollers climb and skip over sprocket teeth. A 30-newton slack-side load was used to prevent jumping on all experiments, except as otherwise noted, to provide experimental safety and more meaningful experimental results. The 10-turn potentiometer shaft was positioned on the output shaft centerline and attached with a rigid connection. The potentiometer casing was held stationary using the laboratory fixture. The 10- turn potentiometer leads and the strain gage leads were connected according to the wiring diagram, Fig. 2. The set-up and calibration procedures followed are given in [6]. The graph paper on the X - Y recorder platen was replaced and, with the pen in the \" u p \" position, the calibration was checked using the actual test load. The handle was turned to bring the active link into the zero position. In this position the chain was fully loaded on both sides. With the pen on the graph paper, the sprocket was rotated, using the handle on the input shaft, until the active link came off the sprocket on the other side of the chain" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002154_tjmj.1985.4549004-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002154_tjmj.1985.4549004-Figure1-1.png", + "caption": "Fig. 1.", + "texts": [ + " The copyright for the original Japanese article is held by the Magnetic Society of Japan. The copyright for the English version of this article as it appears here is held by the Institute of Electrical and Electronics Engineers, Inc. characteristics including the dependence of the pipe impedance on the current, and the strikingly different impedances of the inner and outer pipes. We conducted many experiments to investigate the impedance characteristics of this type of magnetic coaxial pipe and studied possible applications of the structure. Experimental Results and Discussion Fig. 1 shows the coaxial pipe construction. Fig. 2 shows an example of the current dependence of the impedance Zin of the inner pipe and Zout of the outer pipe. The dotted line shows the current dependence of the permeability pb of the outer surface of the inner pipe. It is clear that the current dependence of IZint exhibits the same tendency as IALb. It is also clear that IZouti << IZini . Fig. 3 shows an example of the measured frequency dependence of both Zin and Zout. There exists a frequency that brings Zout to a minimum" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002201_iemdc.2001.939386-Figure9-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002201_iemdc.2001.939386-Figure9-1.png", + "caption": "Fig. 9. Control Volume for the Rotor Inlet", + "texts": [ + " 7 shows the comparison of the ventilation characteristic of the original and the improved configuration without axial mass flow through the rotor. Fig. 8 shows the same comparison for the tests with high axial mass flow. The pressure rise over the entire system was non-dimensionalized by the dynamic pressure based on the rotational speed at the outer diameter of the rotor. This dimensionless parameter is shown as .II, on the y-axis of these p1ots.p is the dimensionless volume flow rate through the stator plotted on the x-axis(see 1). For the case presented in Fig. 9 this simplifies to: 0 0.05 0.1 0.15 0.2 0.25 - I 'p [ - I Fig. 8. Comparison of the Ventilation Characteristics of the Test Motor With and Without Modifications. High Axial Mass Flow A P P In the worst case the pressure on the compression plate p is constant and equal to p l . In this case of sudden acceleration the pressure drop simplifies to: lp=A * uT The density p was taken in the stator-rotor gap. The area A was chosen as the width of the cooling gap times the circumference of the rotor. Fig", + " This is true both for the case with high axial mass flow and without axial mass flow. The increase of mass flow around the actual motor operating point (small pressure rise, @ M 0 - 0.2) is about 10 % in both cases (about 8 % without axial mass flow, about 13% with high axial mass flow). An average increase of 10 % was predicted for the actual motor. B. Rotor Characteristic The ventilation characteristic of the rotor can be deduced theoretically in two steps. First the conservation of momentum is applied to the control volume shown in Fig. 9. Applying the conservation of momentum in the axial direction in the relative frame of reference for an incompressible fluid and neglecting frictional effects one arrives at: In the ideal case of loss-free acceleration the Bernoulli equation applies: This is assuming that the circumferential component of velocity does not contribute to the delivery of air through the axial bores. In reality the pressure drop over the rotor inlet will be somewhere between the sudden acceleration and the loss-free case", + " In this derivation no use needed to be made of the so-called 'Stoherluste' or shock losses. These losses are artificially introduced into the ventilation characteristics by many authors (e.g. [2], [l], [4]). The shock loss coefficient is usually used to fit the theoretical ventilation 5of8 characteristic to the experimental data. That is why the coefficient assumes values in the wide range from 0.5 to 12. For the case of axial bores in the rotor there is no need to use this approach. Simply applying the conservation For the pressure increase from Position 1 in Fig. 9 (Rotor Inlet) to Position 3 in Fig. 10 (Rotor Exit) 6 and 11 yield: 1 2 equations leads to a relatively accurate determination of the losses associated with the transition of the cooling air p 3 - p1 = - . p . ( U 3 2 - U 2 2 ) + c . m2 (12) into the rotor. The second Part of the Pressure change in the rotor takes place in the rotor cooling gaps. The Control Volume is shown in Fig. 10. In 7 the conservation of rotational moAll pressure drops that depend on the mass flow have been sw\",rized in the last term of 12" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002425_cbo9780511529627.012-Figure9.18-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002425_cbo9780511529627.012-Figure9.18-1.png", + "caption": "Figure 9.18. Ratchet mechanism: (a) pawl's advancing the ratchet; (b) pawl is fully advanced; (c) pawl is retracting. White circles indicate re volute joints.", + "texts": [ + " The constraints are expressed as algebraic equations that relate the part coordinates. For example, a ball rolling down a 45\u00b0 slope obeys the constraint x \u2014 y = r, where x and y are the coordinates of the ball's center point and r is the ball radius. The constraints are a function of the shapes of the touching part features (vertices, edges, and faces); hence they change when one pair of features breaks contact and another makes contact. To illustrate contact analysis and its role in design, consider the ratchet mechanism shown in Figure 9.18. The mechanism has four moving parts and a fixed frame. The driver, link, and ratchet are attached to the frame by revolute joints. The pawl is attached to the link by a revolute joint and is attached to a spring (not shown) that applies a counterclockwise torque around the joint. A motor rotates the driver with constant angular velocity, causing the link pin to move left and right. This causes the link to oscillate around its rotation point, which moves the pawl left and right. The leftward motion pushes a ratchet tooth, which rotates the ratchet counterclockwise", + " A system configuration is free when no parts touch, is blocked when two parts overlap, and is in contact when two parts touch and no parts overlap. The mechanical system configuration space can be obtained by combining the configuration spaces of its pairs (Joskowicz and Sacks, 1991), because the system is a collection of kinematic pairs (Reuleaux, 1875). System configuration spaces allow us to analyze multipart interactions but are difficult to compute. We illustrate how pairwise configuration spaces are used in kinematic analysis and synthesis on the ratchet mechanism of Figure 9.18. There are four interacting pairs: driver-link, link-pawl, pawl-ratchet, and pawl-frame. The link-pawl pair is a revolute joint, and thus has a simple relationship: the pawl is constrained to rotate around the pin axis. The driver-link configuration space is two dimensional because both are pinned by re volute joints to the base. The pawl-frame and pawl-ratchet configuration spaces are three dimensional, because the pawl has three degrees of freedom. Figure 9.22 shows the configuration space of the driver-link pair" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003651_cira.2003.1222168-Figure10-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003651_cira.2003.1222168-Figure10-1.png", + "caption": "Figure 10 Whole arm manipulation using a wall", + "texts": [ + " (e, 9) Figure 8: Computational algorithm of dynamic friction case 1 case 2 . . . 0.2 0.25 0.3 0.35 0.4 0.2 0.25 0.3 0.35 0.4 X.2 Figure B: Simulation results of evolution of object's CoM ( ~ 1 . 1 ~ plane) in the presence of dynamic friction Fiwre 11: Geometric condition 3 Environment Adaptive Manipulation In this section, we consider another manipulation problem using preexist environment. As an example, we treat manipulation of an circular object using a wall to distribute forces as shown in Fig. 10. The system model including the object is given by M&)ii+ho(q,Q) = - J , ( q ) T A , (34) [I 1 where q = [ 81 82 E l T and A, is the reaction force from the wall. We asume that the object always touches the robotic arm. In this case, the control requires the information 5 and its derivative, tactile sensor must be equipped Since this system is under-actuated, we must consider input-output linearization approach. In the following, the outline of the control law is described. The position of center of mass of the object along the wall is expressed as: where p is the slope angle, Here let A, := [cos p sin p ] and multiplying it to the left hand of Eiq" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002478_ias.1998.732271-Figure9-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002478_ias.1998.732271-Figure9-1.png", + "caption": "Fig. 9: Vector diagrams for constant positive speed o.", + "texts": [ + " THE MEASURING SET UP The measurement of the values hd(id, i,) and hq(id, in) can become critical if a suitable procedure is not adopted. In fact, two main problems have to be solved: - compensation of the resistive Ri drops, which vary very fastly during overload even at constant current, because of temperature variation; - compensation of the iron losses, since the hypothesis of a conservative system is assumed. Fig. 8 depicts the approximate equivalent circuit of the reluctance motor, in steady state conditions, while Fig. 9 shows the corrispondent vector diagram. R is the winding resistance, while R, represents the iron loss equivalent resistance, the stator stray inductance-being neglected. The torque producing current vector i, differs from the controlled motor current vector i. If the motor is supplied, at constant positive speed w, with i, and i, = < (* stands for complex-conjugate) motoring and braking conditions occur. v I = R i l + j w h , From (1 2) and (13), ( 14) is derived: - (12) - Since g2 f h, the flux obtained from (14) does not correspond exactly to the set current i, in a conservative context" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002407_robot.1986.1087402-Figure3.1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002407_robot.1986.1087402-Figure3.1-1.png", + "caption": "Fig. 3.1 Straight line MOVE", + "texts": [ + " ROBPRO creates a program file and writes all commands combined with the necessary information into this file. It is possible to specify frames with respect to a chosen coordinate system. In order to drive complex trajectories the specification of via-points is possible. These points are included in the MOVE command. According to the interpolation formula F(r) = D(r) F 1 (3.1 1 with D(0) = I, D(1) =F2F1 r = 0,. ..,I -1 with F as start frame and F as destination frame intermediate frames F(r) are calculated. Fig. 3.1 shows a straight line MOVE, keeping the orientation of the hand constant. Fig. 3.2 shows a MOVE from the same start frame to the same destination frame but using two viapoints. The use of ROBPRO shall be demonstrated by an example, a car wheel assembly. Fig. 3.3 shows the workcell as a topview. The robot has to move from its initial position (PI) to the wheel (P2), grasp it, make a quick move in front of the car (P3) and make a precision move to position P4 to fix the wheel at the car. First all positions have to be given" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002342_isie.2001.931572-Figure2-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002342_isie.2001.931572-Figure2-1.png", + "caption": "Fig. 2. Current waveform under constant dwell angle control,", + "texts": [ + "00~ 2001 IEEE 81 1 In case of fixing input voltage of SRM converter, the torque can be controlled by tum-odoff switching angle. Generally, there are two ways to decided tum-on angle or tum-off angle. Fig. 1 shows the constant dwell angle control method. This method controls the tum-on/off angle keep constantly to the change of the variable speed or the load. When tum on angle is moved to keep the constant speed, effect of negative torque is regardless of speed and load. But because of the high limits of rated power, it may be unstable to drive on overload. Fig. 2 shows that the tum-off angle is fixed and the tum-on angle is tuned for a fluctuation of speed and load by constant torque control method. The fluctuation of efficiency is small until rated power, but if the tum-on angle moves toward for an increase torque, even in the region of decreasing inductance, the current will flow and negative torque will be produced. Thus, the efficiency becomes reduced. Therefore, it is needed to find a proper point the position of tum-on angle and the phase current is determined by constant dwell angle control and torque angle constant control method" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002395_pesc.1994.349727-Figure7-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002395_pesc.1994.349727-Figure7-1.png", + "caption": "Fig. 7 : Selection of flux vector, with @ in the shifted region (I'), when vYr.vq > 0.", + "texts": [], + "surrounding_texts": [ + "There are in the literature many methods based on the direct control of voltage inverter switches, in power systems and induction motor drives [2,3,7,9]. An example of this approach is Takahashi and Noguchi proposal [9] which has been referred as IFAM (Improved Field Acceleration Method) [2]. In IFAM control the inverter switching modes are selected so as to restrict the errors of stator flux and torque within two hysteresis bands and to obtain the fastest torque response. The inputs of the switching table are the two hysteresis outputs and the stator flux angular position. The instantaneous angular position is regarded as a discrete value representing the location of flux vector in one of six switching regions, as it is shown in fig. 2. The method presented in this work generalizes the described scheme proposing the use of \"VARIABLE SWITCHING REGIONS\", and it is realized with data already available. It will be showed that this method allows better control of inverter switching pattern by changing the direct and in quadrature components of voltage vector in field coordinates." + ] + }, + { + "image_filename": "designv11_65_0003359_iros.2003.1248967-Figure6-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003359_iros.2003.1248967-Figure6-1.png", + "caption": "Figure 6. Inputs to the fuzzy controller", + "texts": [ + " In order to perform the sorting and grouping\u2019task autonomously by multiple robots, the control framework is divided into two parts as shown in figure 5 . The fyst part handles the path generation using the Voronoi separation algorithm as explained in the previous section. In this process, top view image of the region of interest is processed in order to identify position and type of objects Figure 5. Diagram of pre-planned path generation and fuzzy control in the proposed system The fuzzy controller has 4 inputs as presented in figure 6. 1. The distance from a predetermined subgoal to a robot p i s t& 2. The distance from the nearest obstacle to the robot @is&) 3. The orientation of the predetermined subgoal with respect to the robot reference kame (e,) 4. The angular difference between the orientation of the nearest obstacle and the goal with respect to the robot\u2019s frame (e,) In figure 6, the x-axis of each robot\u2019s 6ame represents heading of the robot. The goal is placed between two obstacles, 01 and 0 2 . In this figure, the nearest obstacle is 01. The four fuzzy input variables, Dist,, Disk, e,, and Bdi% can be described as follows. 1. Dist, represents the distance between the robot to the goal with the membership functions, p(dg) as depicted in figure I. As seen in fignre 7, three fuzzy sets, \u2018goal\u2019, \u2018close\u2019, and \u2018far\u2019, represent how far the robot is kom the goal. For example if the distance between the robot and the goal is 200 mm" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000785_s0094-114x(98)00074-3-Figure4-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000785_s0094-114x(98)00074-3-Figure4-1.png", + "caption": "Fig. 4. Coordinate systems applied for generation of a screw involute surface.", + "texts": [ + " The observation of inequalities (41) and (42) provides that Rf$0 and is collinear to the tangent to the contact line. Thus R(f) is a regular curve and is an envelope to the contact lines on Sr. Generation of a helical involute gear by a rack-cutter is considered. The approach discussed above is applied for determination of envelope Er to contact lines on the generated screw involute surface Sr. We apply coordinate systems Sr, Sr, and Sf that are rigidly connected to the rack-cutter, the gear, and the frame, respectively (Fig. 4). The generating surface Sr is a plane (Fig. 5) represented as r u; y y cos at y sin at u cos lp u sin lp 24 35: 47 The normal Nr to Sr is represented as Nr @r @u @r @y \u00ff sin at sin lp cos at sin lp \u00ff cos at cos lp 24 35: 48 The equation of meshing is f u; y;f Nr v r \u00ff v r u cos lp sin at y\u00ff rpf 0: 49 The general surface Sr is represented as r u; y;f Mrr f r u; y ; f u; y;f 0; jfuj jfyj 6 0: 50 Eq. (20) of singularities yields gr u; y;f \u00ffu cos lp cos at rpf cos at rp sin at 0: 51 Conditions of Theorem 1 (Section 3) are satis\u00aeed, the singularities on Sr form the helix on the base cylinder of the helical gear and this helix is the envelope to contact lines on the generated surface Sr (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002220_77.828379-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002220_77.828379-Figure1-1.png", + "caption": "Fig. 1. B&ng niorlcl (R) atid x-y mcvh (b) on Uic supcrcottdiictor uad iri this aualysis, Titc", + "texts": [ + " Tokyo, lnpnn where, I1 i s magnetic field strength composed by shieldjag current J flowing in the superconductor and the permanent magnet. The critical current density Jc at applied magnetic ficld is followed by Kim model [ 3 ] . whcre Jco is the critical current density at R=O and Bo is the magnetic flux derrsity which inakes Jc half of Jco. magnetic forcc acting on the superconductor: The Lorents force F written as follows can evaluate the P = 3 X Bex dv where Bex is the external magnetic flux density. Fig. 1 shows the bearing model and the x-y mesh (a-b plane) on the superconductor used in this analysis. The dimension of the superconductor is 46\" in diameter and 1 5 \" in thickncss. The size of permanent magnet is the diameter of 3Omm, 15\" thickness, but the same size coil Pig. 2. magnetic flux distribution (a) ad flow palturns (b) oftlic shielding currcnts in the top layer at 3 mm gap and for ccceiihic position, of which displacement is 4 rntn in radial direclion, and f i e rcsult is for zero-ficld coding case", + " Calculntion results Cor field cooling at the gap of 1 mi, md for eccaitnc case, of which dlsplaccmmt is 4 tm. (a) Maglietic Gcld dstrilutiori. (b) Shielding current flow patterns in Each layer ofn, b, c, d. (c) Shiclding current which flow in the must outer radius ring observed from outside, assumed that Jco o f parallel direction to c-axis is one third of that in a-b plane. B. Calculation Rt.suits center. Fig. 3 is the results for field cooling casc at 13 mm gap, and is lmm gap and for no eccentric operation case. Fig. 3(b) shows fhc current distribution patterns in each l a p A, b, c, d shown in Fig.1 (b), and Pig.l(c) shows Ihc current distribution pattoms in the most outer radius ring observed from outside. Frotn these figures, it is clear that thc most shielding currents flow in the top layer, and sharply decayed in the loww layers. Although numerical oscillations in the shielding current are observed near high m a p t i c field of top laycr, the influence of them is negligible small in levitation property. Gap h (mm) Fig.2 shows magnetic flux distribution and flow patterns of the shielding currents in the top layer at 3 mm gap and for eccentric position, of which displacement is 4 mrn in radial direction, and the result is for zcro-field cooling casc" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003864_50002-2-Figure2.3-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003864_50002-2-Figure2.3-1.png", + "caption": "Fig. 2.3. Definition of position, attitude and motion of the wheel and the forces and moments acting from the road on the wheel. Directions shown are defined as positive.", + "texts": [ + " If the road surface near the contact patch can be approximated by a flat plane (that is, when the smallest considered wavelength of the decomposed surface vertical profile is large with respect to the contact length and its amplitude small) the distance of the wheel centre to the road plane and the angle between wheel plane and the normal to the road surface will suffice in addition to the several slip quantities and the running speed of the wheel. For the definition of the various motion and position input quantities listed in Fig.2.1, it is helpful to consider Fig.2.3. A number of planes have been drawn. The road plane and the wheel-centre-plane (with line of intersection along the unit vector 1) and two planes normal to the road plane, one of which contains the vector I and the other the unit vector s which is defined along the wheel spin axis. From the figure follows the definition of the contact centre C also designated as the point of intersection (of the three planes). The unit vector t lies in the road plane and is directed perpendicular to 1. Vector r forms the connection between wheel centre A and contact centre C. Its length, r, is defined as the loaded radius of the tyre. The position and attitude of the wheel with respect to the inertial triad is completely described by the vectors b + a and s.The road plane is defined at the contact centre by the position vector of that point c and the normal to the road in that point represented by the unit vector n (positive upwards). Figure 2.3 also shows two systems of axes (besides the inertial triad). Firstly, we have introduced the road contact axes system (C, x, y, z) of which the x-axis points forwards along the line of intersection (1), the z-axis points downwards normal to the road plane (-n) and the y-axis points to the right along the transverse unit vector t. Secondly, the wheel axle system of axes (A, ~, ~7, ( ) has been defined with the ~ axis parallel to the x axis, the ~7 axis along the wheel spindle axis (s) and the (axis along the radius (r)", + " Through this a rolling resistance force Fr = M y / r arises which necessarily is accompanied by tangential deformations. We may agree that at the instant of observation, point S, that lies on the slip circle and is attached to the wheel rim, has reached its lowest position, that is\" on the line along the radius vector r. At free rolling, its velocity has then become equal to zero and point S has become the centre of rotation of the motion of the wheel rim. We have at free rolling on a flat road for a wheel in upright position (), = 0) and/or without wheel yaw rate ( ~ - 0), cf. Fig.2.3, a velocity of the wheel centre in forward (x or ~:) direction: V x - r e Q (2.2) with (2 denoting the speed of revolution of the wheel body to be defined hereafter. By using this relationship, the value of the effective rolling radius can be assessed from an experiment. The forward speed and the wheel speed of revolution are both measured while the wheel axle is moved along a straight line over a flat road. Division of both quantities leads to the value of re. The effective rolling radius will be a function of the normal load and the speed of travel", + " The sign of the longitudinal slip x has been chosen such that at driving, when Fx > O, x is positive and at braking, when Fx < O, x is negative. When the wheel is locked (g2 = 0) we obviously have x = - 1. In the literature, the symbol s (or S) is more commonly used to denote the slip ratio. The angular speed of rolling Or more precisely defined for the case of moving over undulated road surfaces, is the time rate of change of the angle between the radius connecting S and A (this radius is thought to be attached to the wheel) and the radius r defined in Fig.2.3 (always lying in the plane normal to the road through the wheel spin axis). Figure 2.5 illustrates the situation. The linear speed of rolling Vr is defined as the velocity with which an imaginary point C* that is positioned on the line along the radius vector r and coincides with point S at the instant of observation, moves forward (in x direction) with respect to point S that is fixed to the wheel rim: V - re~'-~ r (2.6) For a tyre freely rolling over a flat road we have: ~c~ r --~-~ and with ~,~ = 0 in addition: Vr = Vx", + " We obtain the expression in terms of yaw rate and camber angle ~, (cf. Fig.2.6): % ~b - t2 sin?, Vcx v* c x (2.15) The minus sign is introduced again to remain consistent with the definitions of longitudinal and lateral slip (2.11, 2.12). Then, we will have as a result of a positive ~o a positive moment M z. It turns out that then also the resulting side force Fy is positive. The yaw rate ~ is defined as the speed of rotation of the line of intersection (unit vector l) about the z axis normal to the road (cf. Fig.2.3). If side slip does not occur (a --- 0) and the wheel moves over a flat road, equation (2.15) may be written as 1 \"Qr 1 ( p - - - - + ..... s in 7 - - - - + R Vcx R 1 V r sin), (2.16) When the tyre rolls freely (then V,x= O, Vcx = Vr) we obviously obtain: 1 1 (P - ---R + ~r sin7 (2.17) e with 1/R denoting the momentary curvature of the path of C* or approximately of the contact centre C. For a tyre we shall distinguish between spin due to path curvature and spin due to wheel camber. For a homogeneous ball the effect of both input quantities is the same. For further use we define turn slip as - ~ ( - - 1 i f a i s c o n s t a n t ) (2.18) ~Pt- vc; -'- 'g Wheel camber or wheel inclination angle ~, is defined as the angle between the wheel-centre-plane and the normal to the road. With Fig.2.3 we find: sin), - - n ' s (2.19) or on level roads: s i n 7 - s z (2.20) where s z represents the vertical component of the unit vector s along the wheel spin axis. The location of the contact centre C and the magnitude of the wheel radius r result from the road geometry and the position of the wheel axle. We consider the approximate assumption that the road plane is defined by the plane touching the surface at point Q located vertically below the wheel centre A. The position of point Q with respect to the inertial frame (O ~ x ~ yO, z o) is given by vector q", + " For a given tyre the effective rolling radius re is a function of amongst other things the unloaded radius, the radial deflection, the camber angle and the speeA of travel. The vector for the speed of propagation of the contact centre V~ representing the magnitude and direction of the velocity with which point C moves over the road surface, is obtained by differentiation with respect to time of position vector c (2.21): V c - d - /; +d + t : - V + t : (2.26) With V the velocity vector of the wheel centre A (Fig.2.3). The speed of propagation of point C* represented by the vector ~ becomes (cf. Fig.2.5 and assume re/r constant): r Vc* - V + --Let: (2.27) r The velocity vector of point S that is fixed to the wheel body results from ?. V s - V + --Leto \u2022 (2.28) r with o9 being the angular velocity of the wheel body with respect to the inertial frame. On the other hand, this velocity is equal to the speed of point C* minus the linear speed of rolling V s - V c - V 1 (2.29) from which Vr follows: 1/r - l\" ( V c - V s) (2" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002391_ias.1996.557040-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002391_ias.1996.557040-Figure1-1.png", + "caption": "Fig. 1. Schematic representation of the proposed PMAC motor operation.", + "texts": [ + " SPLIT PHASE PMAC MOTORS Synchronous generators with two sets of stator windings (six-phase system) displaced 30 degrees from each other, were suggested in the late 1920\u2019s in an attempt to improve efficiency and alleviate certain limitations associated with the operation of large generators [8-111. In later applications, synchronous machines with two sets of stator windings, each set supplied by a separate converter unit, were used to minimize torque pulsations [ 121. Sensorless operation of induction motors with the help of split phase stator windings is also reported recently in the literature [13]. Six-phase induction motor fed by Current Source Inverter (CSI) drives are used in industrial applications in order to eliminate the sixth harmonic torque pulsation [ 14- 151. Figure 1 illustrates the schematic of the stator windings of the PMAC studied in this paper. However, instead of operating the motor as a six-phase motor, the two split phase windings are connected in series (Fig. 2) and the motor is operated as a 3-phase motor supplied by a current regulated pulse width modulation inverter (CRPWM). The taps are, however, made available for voltage measurements. By measuring the line current and voltages of each coil groups absolute position information of the salient pole PM rotor is estimated. The on-line estimation of the absolute rotor position by the proposed technique, does not require any knowledge of the stator resistance, lending itself to a robust scheme against parameter variation. The following section presents the dynamic model of the PMAC motor including all possible coupling between different phases. WINDINGS OF PMAC MOTOR In Fig. 1, the three-phase stator windings a,b, and c are shifted by an angle such as E with respect to the system a\u2019,b\u2019,c\u2019. In general, this angle is 7c16 radians for all possible practical application involving induction and synchronous machines. However, in the present study this angle is considered to be a variable. The windings of each three-phase are uniformly distributed and have axes which are stationary and 120\u2019 apart. The d-q axes are orthogonal to each other and are rotating with the rotor at synchronous speed, y. Without loss of generality, it is assumed that the q-axis is aligned perfectly with the \u201ca\u201d phase axis at time t = 0. Now let 0 = y t , then 0, = 0,, and 0,. = 0+2. In this study, magnetic saturation is neglected. Moreover, the two sets of windings are assumed similar. The voltage equations of the six stator circuits of Fig. 1 in terms of the machine variables are v, = rsi, +ph, (1) where x = a,b,c; a\u2019,b\u2019,c\u2019, r, is the stator resistance, h, is the total flux linkage by each stator phase, and p is the differential operator dldt. The flux linkages of equation (1) in matrix form are: where x=a,b,c; a\u2019,b\u2019,c\u2019; y=a,b,c; a\u2019,b\u2019,c\u2019; and A, is the flux linkage component of each phase due to the rotor permanent magnet. In other words, ph,, is the induced back emf in each phase due to the permanent magnet field. If a sinusoidal distribution of the magnetic field is considered, then the magnetic flux linkages may be represented as 1 (3) where, h, is the maximum amplitude of the permanent magnet field" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002850_b:inam.0000008220.76535.57-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002850_b:inam.0000008220.76535.57-Figure1-1.png", + "caption": "Fig. 1", + "texts": [ + " A design model of a seismically isolated building may be a solid coupled with a mobile base through yielding constraints. Therefore, methods of rigid-body dynamics may be used to study dynamic loads on seismically isolated buildings due to transient motions of the foundation. In the present paper, we construct a model of variable structure to describe the translational motion of a solid with a spherical recess over a fixed spherical support and integrate the resultant differential equations numerically. 1. Equilibrium Conditions for a Solid on a Fixed Spherical Seismic Damper. Figure 1 depicts a solid with a spherical recess of radius R on a fixed spherical support of radius r. The solid has a variety of similar recesses and spherical supports fixed on a horizontal platform; therefore, the body can undergo purely translational displacements. We assume that the solid contacts with the spherical surface at a point and can slide over the supports under the action of gravity and sliding friction. Let us first analyze the static equilibrium of the solid under gravity, forces of Coulomb friction, and horizontal force that acts at contact points. The friction force vectors lie in planes perpendicular to the normals at contact points, and their absolute values are f1N1, where f1 is the coefficient of Coulomb friction and N1 is the normal reaction (the force of interaction of the contacting bodies in the absence of friction). Let us write the equilibrium conditions for an imponderable body on a spherical support. From Fig. 1, we find \u2212 \u2212 + =N f Fx1 1 0(sin cos )\u03b2 \u03b2 , N f F y1 1 0(cos sin )\u03b2 \u03b2+ \u2212 = , (1.1) where Fy is equal to the weight P of the solid per one seismic damper, Fy = P. By the well-known rule, the force of friction is opposite to the positive velocity of sliding in the absence of the force Fx. If Fx, Fy, and f1 are known, then from Eqs. (1.1) we find the angle \u03b2 and limit normal reaction N1: 1063-7095/03/3909-1093$25.00 \u00a92003 Plenum Publishing Corporation 1093 S. P. Timoshenko Institute of Mechanics, National Academy of Sciences of Ukraine, Kiev", + " We take advantage of the principle of kinetostatics, adding the inertial forces due to the translational motion of the foundation to the equations of system (1.1). Thus, we have F m tx =\u2212 ( )\u03be , F m g ty = +( ( ))\u03b7 . (2.1) Let the angle \u03b2 be a generalized coordinate defining the position of the solid. Moreover, we replace f1 by\u2212 fcsign \u03b2, thus turning the force of static friction at the point 1 into the force of sliding friction. The minus sign implies that when the solid moves with \u03b2 increasing, the force of friction is opposite to the relative velocity of sliding (see Fig. 1). Thus, the kinetostatic equations can be written as \u2212 \u2212 \u2212 =N N f m t mxc1 1sin cos ( ) \u03b2 \u03b2 \u03b2 \u03besign , N N f m g t myc1 1cos sin ( ( )) \u03b2 \u03b2 \u03b2 \u03b7\u2212 \u2212 + =sign . (2.2) In the frame of reference Oxyz with origin at the vertex of the spherical seismic damper, the coordinates of the vertex of the spherical recess, i.e., the point A (Fig. 1), are defined by the following formulas, no matter what position the solid is in: x Rr= sin \u03b2, y Rr= \u2212( cos )1 \u03b2 , Rr R r= \u2212 . (2.3) The coordinates of the point A in the frame Oxyz fully specify the position of the translating solid. Substituting (2.3) into (2.2) and eliminating N1, we obtain the desired differential equation for the unknown function \u03b2 \u03b2= ( )t : Rr g t t f Rrc ( ( ))sin ( )cos (\u03b2 \u03b7 \u03b2 \u03be \u03b2 \u03b2 \u03b2 \u03be+ + + + \u2212sign 2[ ]t g t)sin ( ( ))cos\u03b2 \u03b7 \u03b2+ + = 0 . (2.4) The same equation can be derived from Lagrange equations of the second kind, representing the kinetic and potential energy of the system as ( ) ( )[ ]T m Rr t Rr t= + + + 1 2 2 2 cos ( ) sin ( )\u03b2\u03b2 \u03be \u03b2\u03b2 \u03b7 , \u03a0= \u2212mgRr( cos )1 \u03b2 , and the work done by the force of sliding friction during a virtual displacement \u03b4\u03b2 of the solid is \u03b4 \u03b2 \u03b4\u03b2A N f Rrc=\u2212 1 sign , (2" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000342_mchj.1996.0035-Figure4-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000342_mchj.1996.0035-Figure4-1.png", + "caption": "FIG. 4. Cyclic voltammograms of the glucose sensor at a scan rate of 10 mV/s in 0.1 M phosphate buffer (pH 7.0) in (a) the absence of glucose and (b) the presence of 15 mM glucose.", + "texts": [ + " With glucose absent, the enzyme contributes no response and only 1,19-dimethylferrocene gives a constant peak separation dEp of 60 mV. The peak current shows typical diffusion-limited behavior with a linear relation to the square root of the scan rate. The symmetrical surface waves and fast kinetics demonstrate that the presence of regenerated silk fibroin membrane does not appreciably affect the electrochemical behavior of 1,19-dimethylferrocene. No electrocatalytic oxidation current is found at the 1,19-dimethylferrocene modified electrode when glucose is added to the phosphate buffer. Figure 4 shows typical cyclic voltammetric results for the glucose sensor. In the absence of glucose, the glucose oxidase gives no response and only typical oxidation and reduction peaks for 1,19- dimethylferrocene are observed in Fig. 4a. An increase in the anodic current with a concomitant decrease in the cathodic wave is seen in Fig. 4b upon addition of glucose. The comparison of the voltammograms with and without glucose present demonstrates that 1,19-dimethylferrocene can effectively serve as an electron shuttle between the FAD/ FADH2 centers of glucose oxidase in regenerated silk fibroin membrane and a glassy carbon electrode. The flavin adenine dinucleotide (FAD) of glucose oxidase (GOD) is reduced by the glucose penetrating the membrane: b-D-glucose+ GOD~FAD! \u2192 d-gluconolactone + GOD~FADH2!. Electrons are transferred from the reduced GOD(FADH2) to 1,19-dimethylferrocene ion (DMFc+): GOD~FADH2" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001007_jsvi.1997.0998-Figure12-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001007_jsvi.1997.0998-Figure12-1.png", + "caption": "Figure 12. An example of a belt system with opposing belt movement.", + "texts": [ + " (13) For the combination shown in Figure 11(b), the relationship of the velocity of the belt at pulley a and pulley b is expressed as vb =\u2212va . (14) Therefore, if a belt system consists of n pulleys, the relationship between the velocity of the belt at pulley i and the velocity of the belt at pulley j, when pulley i is located next to pulley j (i, j=1, . . . , n), is expressed by equations (13) and (14). Unless all the equations are satisfied, opposing belt movement will occur. For the belt system in Figure 12, for exmaple, the equations of the velocities of the belts are obtained as v1 = v2, v2 = v7, v7 = v1 for loop L1 v3 = v2 for loop L2 v3 =\u2212v4, v4 = v5, v5 = v6, v6 =\u2212v3 for loop L3 h G G G G J j . (15) v6 = v7 for loop L4 Because these equations are not satisfied, opposing belt movement will occur. Belt systems are usually driven by a three-phase induction motor with a squirrel cage rotor. Therefore, to analyze the transient responses of belt systems, the transient output torque of the induction motor must be examined" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001436_a:1008131702701-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001436_a:1008131702701-Figure1-1.png", + "caption": "Figure 1. Link object Ai and obstacle object Oj .", + "texts": [ + " The general solution of Equation (2) is \u03b8\u0307 = J+x\u0307e + (I \u2212 J+J )g, (3) where J+ = J T(JJ T)\u22121 is the pseudoinverse of J , I \u2208 Rn\u00d7n is the identity matrix, and g \u2208 Rn is an arbitrary vector in the joint-velocity space which can be used to resolve the redundancy at the velocity level in optimizing a suitable performance criterion. We use A and O to denote the manipulator links and obstacles, respectively. The manipulator and obstacles are represented by unions of objects as follows: A = \u22c3 i\u2208IA Ai , O = \u22c3 j\u2208IO Oj , (4) where Ai , i \u2208 IA = {1, . . . , n}, are manipulator links, and Oj , j \u2208 IO = {1, . . . , nO}, are nO obstacles. Consider a link-obstacle pair as shown in Figure 1. The problem is to determine a joint trajectory \u03b8(t) of the manipulator so that its end-effector can move along the desired trajectory while the manipulator A is kept away from obstacles O. Link Ai is assumed to be a cylinder, which can be described as an ellipsoid containing the link centered at yic in the link i coordinate frame described as [20] Ai(\u03b8) = { Ti(\u03b8)y: (y\u2212 yic)TQT i Qi(y\u2212 yic) 6 1 } (5) with Qi = [ 1/ria 0 0 0 1/rib 0 0 0 1/rib ] , (6) where ria and rib are scalar coefficients, and Ti(\u03b8) is a homogeneous transformation matrix for link i coordinate frame which consists of an orthogonal rotation matrix Ri(\u03b8) and a position vector pi(\u03b8) [3]: Ti(\u03b8) = [ Ri(\u03b8) pi(\u03b8) 0 0 0 1 ] " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002607_978-94-017-0657-5_14-Figure3-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002607_978-94-017-0657-5_14-Figure3-1.png", + "caption": "Figure 3. Control constraint", + "texts": [ + " For CABLEV the inversion of the system dynamics is obtained in two steps. First the robot coordinates pare expressed in terms of the prescribed outputs y and their time deriva tives (Maier and Woernle, 2000). With seven robot coordinates p and six platform coordinates y the system is kinematically redundant. This redundancy is here treated by introducing a geometric control constraint between the trolley co ordinates Pg2, Pg3, Pg4, go(p) = Pg2 - ~(Pg3 + Pg4) - C = 0, c = const . (10) According to Fig. 3 this constraint means that the positions of the outer trolleys are always symmetrical to a point P on the intermediate rail with constant distance c to the inner trolley. With (8), (5), (10) a set of ten nonlinear equations is available to calculate the unknowns p E lR? and ,X E IR3 for given y, ~, ~: G;(p,~y)~'x = Ms(Y) ~ + ks(Y,~) - Q;(y) } g(y,p) = 0 . go(p) = 0 (11) The velocities ~ and the accelerations ~ can be expressed in terms of the output derivatives ~, ~ ;tsing (3). System (11) can be solved numerically with respect to p and A" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001075_1.370145-Figure2-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001075_1.370145-Figure2-1.png", + "caption": "FIG. 2. The focal spot is rotated through an angle f, and is scanned along the positive x direction with a constant speed v . It has the same heating effect as keeping the focal spot unrotated and scanning it at an angle fs , along xs direction with a constant speed v , such that fs52f .", + "texts": [ + " Another set of axes x1 and y1 are so chosen that its origin lies at the beam center and the x1 and y1 axes are parallel to the length and width of the rectangular spot, respectively. The laser beam is scanned along the x axis at a constant velocity v . When the beam is rotated by an angle f from the x axis, the x1 and y1 axes are also rotated so that they subtend an angle f with the x and y axes, respectively. This situation is equivalent to keeping the length and width of laser spot parallel to the x and y axes, respectively, and scanning the beam in the xs direction at scanning angle fs 52f ~see Fig. 2!. The relationship between the rotated coordinate system (x1 ,y1) and the original coordinate system ~x,y! is given by:19 x15x cos f1y sin f , ~2a! y15y cos f2x sin f . ~2b! The temperature T(x ,y ,z) is a normalized temperature defined by T5 T12T0 Tm2T0 , ~3! where T1 and T0 are the substrate and ambient temperatures, respectively. The boundary conditions for the above differential equation are T\u21920 as x\u21926` ~4a! T\u21920 as y\u21926` ~4b! T\u21920 as z\u2192` ~4c! k ]T ]z 52 1 Tm2T0 I , at z50. ~4d! The laser irradiance I is given by I5(m50 M (n50 N Imn , M and N are the largest mode numbers in the x1 and y1 directions, respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003142_isic.1989.238662-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003142_isic.1989.238662-Figure1-1.png", + "caption": "Figure 1: (a) A planar cat, (b ) a symmetrical cat.", + "texts": [ + " Forward somersault, or flip, is a gymnastic maneuver in which the performer runs forward, springs off the ground with booth feet, rotates the body forward through 360\u201d degrees, and lands in a balanced posture on one or both feet. Human gymnasts can do a forward flip as an isolated maneuver or as part of a floor routine in which the flip is proceeded and followed by other maneuvers. The best gymnasts can do double and even triple flips. The average teenager can learn to do a forward flip in a few weeks with proper coaching and practice. In this report we study the dynamics of a planar cat (see Figure 1) and then propose several possible strategies that a cat can use to perform forward somersault. The notation used here follows closely that of (lo]. 2 Dynamics Consider a free-fall configuration of the planar cat shown in Figure la. A special caw of this, which we shall call the symmetrical cat, is shown in Figure Ib. Let C, be the inertia reference frame, CO the frame which is fixed to the mass center of the biped and has the same orientation as the inertia frame. Let C , , i = 1,2,3, be the frame fixed to the mass center of body i ", + ", 2 6, (2) be the moment of inertia about the 2- axis, and the mass of body i . Then, the kinetic energy integral (2) can be further simplified to (3) 1 1 h-. - -I,wT t -mi~~i,~12. * - 2 2 where UJ, = e, is the angular velocity of body i. system Summing (3) over i yields the total kinetic energy of the (4) In order to decmple rotational motion from translational motion, we need to express the kinetic energy (4) as the sum of a translational component of the mass center and a rotational component about the mass center of the system. For this, Figure 1 reveals the following kinematic relations. * ti = t t t;, i = 1,2,3. (5) Substituting ( 5 ) into the second component of (4) yields where we have used the fact that i=3 (E mi+:, i ) = o i=l as c b is positioned at the mass center of the system, and i=3 m = C m , i=l is the total mass. forward from Figure 1. Furthermore, the following kinematic relations are straight- Substituting (5) and (8) into the following equation ry = rl - r = t-1 - C C,r,, e, = tn , /m , I d ,=I yields * We can then write (see Figure 1) and Thus, differentiating tf with respect to time 1, yields I l i : l l2 = (\u20ac2 - \u20ac3)2&4 t \u20ac;+I; t \u20ac;&J - 2\u20ac2(\u20ac2 - \u20ac3)ddl C O S & ~ U J I W ~ (12) - 2\u20ac3(Cz - \u20ac3)dd2CoS@31t01w3 t 2\u20ac2\u20ac3d1d2CoSe3Zw2W3 where B,, = B, - 0, denotes the relative orientation of body i with respect to body j , and lli$ = (1 t \u20ac3 - ~ 2 ) ~ d \u2019 ~ : t (1 - C2)2d iW; t 44 W: t - 2t3(~2 - 4 - i)dd, cos e31 w3tt11 - 2\u20ac3(1- C2)dld2 cose32w2w3; (13) - F3 - i ) ( i - C 2 ) n n , co8~21wltu2 Ili$l12 = - (1 + \u20ac2 - \u20ac3)2d2W: t f:di?lJ: t (1 - \u20ac3)\u2019$10: 2(1 + \u20ac2 - \u20ac3)\u20ac2ddl COS82lWlt l12 + 2(1 t e2 - c3)(1 - ~ ~ ) n n ~ c o ~ e ~ ~ u l ~ w ~ (14) - 2~2(1- ~ 3 d i d 2 cm BjZwzt~3" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001092_cnm.1630080604-Figure2-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001092_cnm.1630080604-Figure2-1.png", + "caption": "Figure 2. Balcony beam: E = lo6 Ib/in2, Y = 0", + "texts": [ + " A comparison of the eigenvalues calculated for fully and reduced integrated prismatic elements revealed the presence of six bending modes with eigenvalues lowered to various degrees by reduced integration. Stress interpolation is best performed using the reduced quadrature points (as demonstrated for other elements by Barlow l4 and others) and is applied to the beam element in this manner. The interpolation is therefore quadratic for the cubic beam. To demonstrate the effects of reduced integration and the use of the element in spring vibration analysis a number of examples were performed. Only the cubic element ( N = 4) was used. Example I : Balcony beam The balcony beam problem' (see Figure 2) was used to illustrate the accuracy of force resultants as a function of the quadrature schemes. The displacement and local force results are presented in Table I. It comes as no surprise that the displacement results for reduced VIBRATION ANALYSIS OF COIL SPRINGS 377 integration (URI) are better than those for full integration (FI). The internal force accuracy is, however, largely dependent on the quadrature rule. This fact is clearly illustrated by comparing transverse shear forces. For URI the transverse shear forces are exact, namely 1 " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003585_s0219878904000069-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003585_s0219878904000069-Figure1-1.png", + "caption": "Fig. 1. Passing-C.M. line and vertical operation plane.", + "texts": [ + " The object is heavy enough and there is sufficient friction between the object and environment, so that the object can only roll without slipping on the environment. Based on these assumptions, the purpose of this paper is to estimate the mass and the center of mass of the graspless curved-surface object by dexterous object manipulation on an environment. 2.2. Passing-C.M. line Consider leaning and tipping an object, which is in plane-contact with an environment plane, by a fingertip force acting on the object (see Fig. 1). When the object is leaned, it will contact the environment plane at a point because of the curved edge of the object base plane. At the beginning of leaning, the fingertip force on the object, the object gravity and the reaction force from the contact point are unbalancing one another, and that the contact point will move indefinitely. With that the object rolls without In t. J. I nf . A cq ui si tio n 20 04 .0 1: 47 -5 5. D ow nl oa de d fr om w w w .w or ld sc ie nt if ic .c om by N A N Y A N G T E C H N O L O G IC A L U N IV E R SI T Y o n 04 /2 7/ 15 ", + " In this paper, this straight line will be referred to as the Passing-C.M. Line. The different contact points at the edge of the object base correspond to different passing-C.M. lines. Two or more different oriented passing-C.M. lines intersect at one point, i.e., the center of mass of the object. Hence, the center of mass of object can be estimated by estimating two or more different oriented passing-C.M. lines in an object. 2.3. Parameters of passing-C.M. line When an object is leaned at rest in point-contact with an environment (see Fig. 1), the equilibrium condition of the moment about the contact point can be expressed as lR \u00d7 f + T \u00d7 g = 0, (1) T = mlg, (2) where lR is a distance vector from the contact point to the fingertip which keeps in touch with the object, f is the contact force at the fingertip measured by finger force sensor, g is the gravity. T is the toward-C.M. vector, m is the mass of object, lg is a distance vector from the contact point to the center of mass. According to Eq. (2), vectors T and lg have the same direction but their magnitudes are different" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003067_iros.1998.724635-Figure5-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003067_iros.1998.724635-Figure5-1.png", + "caption": "Fig. 5: Variables in eqs.(15),(16),(17) and (18)", + "texts": [ + " A suficient condition for achieving the palmreaching phase by a horizontal pushing motion of finger links is that the contact points between an object and a finger link become lower than the horizontal line including the contact point between two objects. [ProofJ Eq.(ll for the two objects which are projected to r are de d ned as D F - W , = D B ~ ~ P B ? . (12) DF,At?, is defined as DFYAt?, = [cl 0 - \u20ac2 0 0 0IT where \u20ac1 (> 0) and c2 (> 0) are the displacements of finger links in the horizontal direction. D B ~ and ApB7 are also defined as As shown in Fig.5, when the contact points between the object and each finger link are lower than the horizontal line including the contact point between two objects, we can define Moreover, from geometrical relationships, the following equations are obtained From eqs.(l7) and (181, we can show that object 1 rotates counter-clockwise and that object 2 rotates clockwise. This holds the theorem. 0 While the manipulation of multiple objects by rolling contact seems to be very difficult, the above theorem shows that the palm-reaching phase can be achieved by applying a simple pushing motion that can produce both \u20ac1 and \u20ac2" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002356_acc.2000.879471-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002356_acc.2000.879471-Figure1-1.png", + "caption": "Fig. 1. Symmetric balance beam with magnetic bearings.", + "texts": [ + " However, a very specific one must be constructed, called a diffeomorphism, that twists the original system into a new one with specific properties: unique feedback linearization that is regular and guarantees controllability. This mathematical technique provides an unambiguous way of obtaining the coordinate transformation and a natural way of placing constraints on the actuator gineering Department of the University of Virginia, Charlottesvzle, Virginia, USA. E-mail: peaQvirginia.edu fluxes. 0-7803-551 9-9/00 $1 0.00 Q 2000 AACC 1602 11. SYSTEM MODEL Figure 1 illustrates the system to be considered in this paper. Essentially, a rigid beam of moment J is simply supported at its center of mass by a pivot designated by 0. At a length La to either side of the pivot are horse-shoe electromagnets labeled A1 and A2, which produce forces F1 and F2. Each actuator has a voltage input v1 and 212 and flux state variables $1 and 4 2 . The remaining two states of the system are the beam angle, 8, and angular velocity, 6'. With knowledge of electro-magnetics, it is not difficult to show that the state space form the nonlinear dynamic equations describing the magnetic bearing system are dx d t - Where c1 and c2 are physical constants and go and N axe the nominal magnetic gaps and wire coil turns respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000537_s0263574700015411-Figure2-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000537_s0263574700015411-Figure2-1.png", + "caption": "Fig. 2. The robot PUMA 560 with coordinate systems according to the modified Denavit-Hartenberg notation.", + "texts": [ + " These data are used as input for the forward dynamics program which results in joint accelerations. The latter are exactly the same as the accelerations generated by the trajectory generator at the very beginning of the simulation process. Several examples21 have been run in order to test the software packages. There have been considered the inverse and forward dynamics models for such robots as DDA (Direct Drive Arm) I and II,7 Puma 56016 and others. As a representative example let us consider the dynamic model of the PUMA 560 robot. The PUMA manipulator is shown in Figure 2. Notice that in Figure 2 we have used the modified Denavit-Hartenberg notation and the link numbering system starting from the tip of the manipulator toward the base, according to the convention introduced by Rodriguez. The mass parameters of the links have been enumerated in reference 16 and we are not going to describe these parameters in detail. All joints of the manipulator were moved simultaneously in a time period of 4 seconds. The total movements of joints were respectively: joint 1 in the range from joint 2 in the range from joint 3 in the range from joint 4 in the range from joint 5 in the range from joint 6 in the range from -150\u00b0 to 150\u00b0, -90\u00b0 to 90\u00b0; -130\u00b0 to 130\u00b0, -240\u00b0 to 70\u00b0, -220\u00b0 to 40\u00b0, -150\u00b0 to 150\u00b0, with zero initial conditions for velocities and accelerations" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002993_1.1737377-Figure6-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002993_1.1737377-Figure6-1.png", + "caption": "Fig. 6 A 6R mechanism with one degree of freedom", + "texts": [ + " Hence the mechanism DOF is one. Also the assembling configuration in Fig. 5 is a singular configuration but very likely chosen as initial configuration for the X-mechanism. There are only two independent constraints in this configuration and again a simple rank determination is not admissible to eliminate redundant constraints. It can be shown analytically that C2,3 is equivalent to so(3). 6R Mechanism. An example for which the dimension of the constraint space really depends on the chosen cut joint is the 6R mechanism in Fig. 6 constructed from two planar 3R chains where the respective planes of motion are orthogonal. Possible situations for Ci , j are dim Ci , j55 and dim Ci , j56, where the minimum is achieved for C3,5 as shown in Fig. 6 but also J3 could have been used as cut joint. The constraint vector subspace is the union of two representations of se(2). With d55 the mechanism DOF is estimated as 1 and the algorithm yields fife independent constraints for the six generalized velocities. In fact this is the minimal set of constraints. Bricard Mechanism. This is one of the well known paradoxical mechanisms that disobey the GKC formula but also any attempt to explain their mobility with intersections of algebras. Regardless which of the joint is considered as cut joint dim Ci , j 56 and thus no constraints can be eliminated though the Bricard mechanism has the DOF 1" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001511_ccece.2001.933550-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001511_ccece.2001.933550-Figure1-1.png", + "caption": "Fig. 1 : Generator System", + "texts": [ + " The performance of the induction generator is studied through modeling it by a novel equivalent circuit in a pseudo-stationary abc-dq reference frame. Based on the circuit model a state space mathematical model is develop ed. The model is capable of dealing with the nonlinearities introduced by the used electronic solid-state switches. The performance characteristics have been computed for a wide range of operating conditions through a simulating computer program. 2. PROPOSED CONTROL CIRCUIT The proposed circuit is shown in Fig. 1. Each stator phase has a control circuit that consists of two antiparallel thyristors. This control circuit links the induction generator to the network. The terminal voltage of the generator is controlled by controlling the triggering angle, a of the thyristors. - 0839 - The current begins to flow at this angle and the thyristor is naturally commutated when the current falls down to zero. The current, power factor, active and reactive power of the generator are controlled through the variation of the triggering angle of the thyristors" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001273_rob.4620080205-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001273_rob.4620080205-Figure1-1.png", + "caption": "Figure 1. A composite body.", + "texts": [ + " However, the forces and torques and the center of mass of a composite body are all observed in an inertia frame, namely the base frame Eo. And we want to derive a closed-form formulation for the inertia matrix, although some terms in the formulation are still computed in a recursive form. Suppose that a manipulator has n low-pair joints, which are labeled as joint 1 to n outward from the base. Assign a body-fixed frame on each joint, i.e., frame Ei is fixed on joint i. The distance from the origin of E; to that of Ej is designated as $, and the distance from the origin of Ei to the center of mass of linkj as jp (Fig. 1) . Define a composite bodyj as the union of l inkj to link n. The mass of the composite bodyj is denoted as mj , and the distance from the origin of the base flame to the center of mass of the composite body as rj. Hence where mi is the mass link i. The inertia tensor of the composite body, Jj , results by using Huygeno-Steiner formula14 to obtain 200 Journal of Robotic Systems-1991 where Ii is the inertia tensor of link i, and [ax] denotes a skew-symmetric matrix representing the vector multiplication (ax), i" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003271_6.2002-4761-Figure3-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003271_6.2002-4761-Figure3-1.png", + "caption": "Fig. 3 Admire control surface configuration.", + "texts": [ + " The model parameters B and c in (4), used in (5) and (6), are computed at each sampling instant by linearizing M(x, u) around the current measurement vector, x(t), and the previous control vector, u(t\u2212 T ). In the Admire model, T = 0.02 s. The constrained QP problem (7) is solved at each sampling instant using the sequential least squares solver from ref. 14. The control vector, u = ( u1 . . . u7 )T consists of the commanded deflections for the canard wings (left and right), the elevons (inboard and outboard, left and right), and for the rudder, in radians, see Figure 3 where \u03b4\u2217 denote the actual actuator positions. The actuator constraints 6 American Institute of Aeronautics and Astronautics are given by \u03b4min = (\u221255 \u221255 \u221230 \u221230 \u221230 \u221230 \u221230)T \u03b4max = (25 25 30 30 30 30 30)T \u03b4rate = (50 50 150 150 150 150 100)T measured in degrees, and degrees per second, respectively. At trimmed flight at Mach 0.5, 1000 m, the control effectiveness matrix is given by B = 10\u22122\u00d7( 0.5 \u22120.5 \u22124.9 \u22124.3 4.3 4.9 2.4 8.8 8.8 \u22128.4 \u221213.8 \u221213.8 \u22128.4 0 \u22121.7 1.7 \u22120.5 \u22122.2 2.2 0.5 \u22128.8 ) from which it can be seen, e" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002765_s0045-7949(02)00029-9-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002765_s0045-7949(02)00029-9-Figure1-1.png", + "caption": "Fig. 1. Free-body diagram of simple two-body system.", + "texts": [ + "1) and is equipollent to the generalized reaction force which acts at the origin of the local coordinate system of the body concerned and is given by Qc;i \u00bc F M \u00fe \u00f0Ai uP ;i F\u00de k \u00f04:2a\u00de and the moment is supplemented by the cross product of the position vector of the revolute joint with respect to the new coordinate system and the original force. This may be reorganised as Qc;i \u00bc I 0 uTp;iA T h;i 1 F M ; \u00f04:2b\u00de since \u00f0Ai up;i\u00de h F i k \u00bc uTp;iA T h;iF: \u00f04:3\u00de To illustrate the above, consider a two body system with one body fixed to the ground and the second free to rotate about a revolute joint connecting the bodies. The free-body-diagram of such a system is shown in Fig. 1. The explicit equations of motion in the generalized coordinates for both bodies are as follows: Body 1 m1 \u20acRx1 \u00bc Fx1;2 \u00fe Fx2;1 \u00bc 0; \u00f04:4a\u00de m1 \u20acRy1 \u00bc Fy1;2 \u00fe Fy2;1 \u00bc 0; \u00f04:4b\u00de J1 \u20ach1 \u00bc M1 \u00bc 0: \u00f04:4c\u00de Body 2 m2 \u20acRx2 \u00bc Fx1;2 ; \u00f04:5a\u00de m2 \u20acRy2 \u00bc Fy1;2 \u00fe m2g; \u00f04:5b\u00de J2 \u20ach2 \u00bc Fx1;2 l2 2 sin h2 Fy1;2 l2 2 cos h2 \u00feM2: \u00f04:5c\u00de The constraint equations for this system are C q; t\u00f0 \u00de \u00bc Rx1 Ry1 h1 Rx1 Rx2 \u00fe cos h2\u00f0 \u00de l2 2 Ry1 Ry2 \u00fe sin h2\u00f0 \u00de l2 2 2 6666664 3 7777775 \u00bc 0: \u00f04:6\u00de On construction of Eq. (2.18) and comparison with Eqs", + " 4. Similarly for the wheel loader simulation, the rate of change of torque tends to zero at the half-way point (two seconds of simulation) as a result of the bucket constraint (see Eq. (4.13)) forcing a change in the derivative in the generalized coordinate graphs (not shown). These torques are the required torques to satisfy the imposed driving constraint. From the body constraint forces and torques Qc, one may obtain the joint forces and torques as described in Section 4.1, through the aid of Fig. 1 and Eqs. (4.1), (4.2a), (4.2b), (4.3), (4.4a)\u2013(4.4c), (4.5a)\u2013(4.5c), (4.6), (4.7). Hence from the prescribed bucket trajectories, it is possible to obtain the generalized body and joint constraint forces and torques of connected mechanisms, including hydraulic actuators, in an inverse-type approach. A multibody system approach is used to model the dynamics of excavator digging operations normally conducive to the study of kinematics and dynamics of robotics work. The kinetic energy of the ith body in the system is substituted into the Lagrange equations to arrive at the Newton\u2013Euler equations of motion for the multibody system" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003496_s0022-3913(75)80139-9-Figure5-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003496_s0022-3913(75)80139-9-Figure5-1.png", + "caption": "Fig. 5. Trigonometric values required to make computation of compensating angle.", + "texts": [ + " Both of these factors affect the direct application of the old formula, and they have been incorporated in the charted computations to provide a like effect. Volume33 New semiadjustable articulator. Part 1 15 N u m b e r 1 The compensating angle (Table I\u2022) has been computed for various patient intercondylar distances in combination with various articulator Bennett angles. Four ph);sical dimensions are available: (I) patient's intercondylar distance, (2) articulator intercondylar distance, (3) balancing excursion distance, and (4-) articulator Bennett angle (Fig. 5). Trigonometric calculations utilizing these factors define the patient's (5) mandibular side shift, (6) the articulator Bennett shift, and (7) the articulator compensating angle. S U M M A R Y An attempt has been made to improve the articulator mechanism over existing semiadjustable articulators. The Bennett movement produced on the XP-51 is a straight lateral movement which is generally not the type of movement produced by the patient. The surfaces of the glenoid fossae are not flat as demonstrated on articulators" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001092_cnm.1630080604-Figure4-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001092_cnm.1630080604-Figure4-1.png", + "caption": "Figure 4. Pretwisted beam: E = 200 GPa, G = 76 GPa, J* = I, + I,, ky = k, = 2", + "texts": [ + " The beam is subjected to self-weight y = 2-5t/m3. For the analysis, the angle \\I. is constant and chosen as 0-0. Results are presented in Table 11. Displacement convergence is better for URI, and force accuracy is dramatically better. For coarse meshes, results are also better than those of Reference 7 except for MY min. Example 3. Pretwisted beam To illustrate the ability of the element to represent a strongly pretwisted beam, the Tabarrok pretwisted ~an t i l eve r~~\" with shear load at the tip was used (see Figure 4). Such modelling 378 N. STANDER AND R. J . DU PREEZ VIBRATION ANALYSIS OF COIL SPRINGS 379 capability is often required for slender turbine blades or coil springs in which the wire crosssection is non-circular. In the example the twist angle varies linearly from $ = 0 at the support to I) = - 90\" at the tip. A convergence study was conducted in which one, two and four elements were used to compute tip displacements. Reduced integration was used throughout. Table I11 presents the tip displacements" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003167_60.986445-Figure6-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003167_60.986445-Figure6-1.png", + "caption": "Fig. 6. Pertaining to field calculation.", + "texts": [ + " 5 due to the current \u201c \u201d flowing along , can be expressed, using Biot\u2013Savart law, as (1) where (2) and (3) Similarly, the flux density components , and due to the current flowing through the corresponding conductors are expressed and is obtained as the sum of the above four components. It can be seen from Fig. 5 that the flux density component along the -axis is produced by the current flowing in the conductors and and the components along -axis is produced by the current in the conductors and . For example, using the geometrical relations in Fig. 6, the flux density component along -axis due to current in can be shown to be same as (1) with ( ) replaced by . In a similar manner the flux density components along and directions due to all the sides of the coil can be calculated. The resultant values of the normal, longitudinal and lateral components of the flux-densities due to all the side coils show that the longitudinal component ( ) is negligible while the normal and lateral components ( and ) show little variation with . An explanation of the above observation is that since all the coils (under one side wall) carry currents in the same direction, the fields induced at any field point virtually have infinite pole pitch along the length and hence they do not vary with " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003407_j.carbon.2004.01.021-Figure2-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003407_j.carbon.2004.01.021-Figure2-1.png", + "caption": "Fig. 2. (a) Definition of the director orientation of a uniaxial discotic nematic liquid crystalline material. The director vector n is the average orientation of the unit normal vector u to the disk-like molecules. (b) Schematics of the capillary flow of a uniaxial discotic nematic liquid crystalline material and cylindrical (r, /, z) coordinate system used to describe a generic point P ; showing an uniaxial disc-like molecules with unit normal vector u, director vector n, velocity vector v, velocity gradient rv, and the alignment angle between the director and the axial direction h.", + "texts": [ + " First, we review the elastic and continuum theory of liquid crystals for uniaxial discotic nematic thermotropic liquid crystals, which is the case of the carbonaceous mesophases, then we applied it to capillary Poiseuille flow, followed by discussions of flow-induced macrotextural phenomena and non-Newtonian rheology. nu u u uu (a) In flowing liquid crystal systems both the elastic and viscous stresses are normally important. The continuum theory of elasticity of liquid crystals, developed by Oseen and Frank [4,5], takes into account that external forces can introduce deformations in the relative orientations and can distort the equilibrium configurations of liquid crystals. A schematic representation of the discotic nematic liquid crystalline phase is shown in Fig. 2a, where the unit normal vectors to the molecular discs (u) orient more or less parallel to the director vector n (where, n n \u00bc 1). The elastic free energy density Fd for the flowing li- quid crystal systems is given by: Fd \u00bc K11 2 \u00f0r n\u00de2 \u00fe K22 2 \u00f0n r n\u00de2 \u00fe K33 2 jn r nj2 \u00f01\u00de where the three basic modes of elastic storage are the splay (K11), twist (K22) and bend (K33) modes. The splay and bend distortions are planar, while the twist distortion is along an axis normal to the director orientation. The continuum theory of uniaxial nematic liquids was developed by Ericksen and Leslie [4,5]; according to this theory, for incompressible isothermal conditions, the general conservation of linear and angular momentum are given by the following equations: q ov ot \u00fe v rv \u00bc f \u00fer r \u00f02\u00de q1\u20acn \u00bc G\u00fe g\u00fer p \u00f03\u00de where q is the density, v is the velocity vector, f is the body force per unit volume vector, r is the total stress, q1 is the moment of inertia per unit volume,G is the external director body force vector, g is the intrinsic director body force vector, and p is the director stress tensor", + " \u00f09\u00de where p is the pressure, I is the unit tensor, faig, i \u00bc 1; 2; 3; 4; 5 and 6, are the six Leslie viscosity coefficients that describes an anisotropic liquid, A is the rate of deformation tensor, N is the corotational derivative of the director vector, b is a Lagrange multiplier vector, c1 is the rotational viscosity, c2 is the irrotational torque coefficient, W is the vorticity tensor, k is the reactive parameter, and hk s are the k flow-alignment angles multiples that exist when k < 1. The Poiseuille capillary flow of discotic nematic liquid crystals can be simulated using the macro-scale theory of Ericksen\u2013Leslie (Eqs. (1)\u2013(6)). Assuming that the director orientation vector is confined to the (r, z) plane (see Fig. 2b), and the velocity field v is in the axial direction: n\u00f0r\u00de \u00bc \u00f0sin h\u00f0r\u00de; 0; cos h\u00f0r\u00de\u00de \u00f010\u00de v\u00f0r\u00de \u00bc \u00f00; 0; v\u00f0r\u00de\u00de \u00f011\u00de the resulting governing equations in dimensionless form for the director tilt angle (h) and axial velocity component (~v), under steady state and isothermal conditions, are [6\u20139]: \u00f0cos2 h \u00fe e sin2 h\u00de d2h d~r2 \u00fe 1 ~r dh d~r ! \u00fe sin 2h 2 \u00f0e \" 1\u00de dh d~r 2 1 ~r2 # eH \u00f0h\u00deEr 2eG\u00f0h\u00de ~r \u00bc 0 \u00f012\u00de d~v d~r \u00bc Er 2eG\u00f0h\u00de ~r \u00f013\u00de where eG\u00f0h\u00de \u00bc ~a1 sin 2 h cos2 h \u00fe \u00f0~a5 ~a2\u00de 2 sin2 h \u00fe \u00f0~a3 \u00fe ~a6\u00de cos2 h \u00fe ~a4 \u00f014\u00de 2 2 eH \u00f0h\u00de \u00bc ~a2 sin 2 h ~a3 cos 2 h \u00f015\u00de Er \u00bc R3\u00f0 dp=dz\u00de=K11 \u00f016\u00de e \u00bc K33=K11 \u00f017\u00de ~ai \u00bc ai= g \u00f018\u00de g \u00bc \u00f0g1 \u00fe g2 \u00fe g3\u00de=3 \u00f019\u00de g1 \u00bc \u00f0a3 \u00fe a4 \u00fe a6\u00de=2; g2 \u00bc \u00f0 a2 \u00fe a4 \u00fe a5\u00de=2; g3 \u00bc a4=2 \u00f020a;b;c\u00de e is the ratio of the bend and the splay Frank elastic constants, Er is the ratio of viscous flow effects to longrange elasticity effects (known as the Ericksen number), dp=dz is the given pressure drop in the capillary per unit length, ~r is the dimensionless radial distance (~r \u00bc r=R), ~ai are the dimensionless Leslie viscosity coefficients, g is the average Miesowicz\u2019 viscosity, g1, g2, and g3 are the three Miesowicz\u2019 shear viscosities that describes an anisotropic liquid, R is the capillary radius, and ~v is the scaled velocity (~v \u00bc gRv=K11)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002346_robot.1998.677406-Figure2-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002346_robot.1998.677406-Figure2-1.png", + "caption": "Figure 2. Reference coordinate system.", + "texts": [ + " The body tree array r showing at column i the topology of the chain from body i to the ground and the body exponent array H showing the number of bodies in the chain are constructed as T(1,i) = i T(2,i) = j (1) T ( k , i ) = T(2 ,T(k - 1,i)) and H(i) = k - 1 when T ( k , i ) becomeszero first (2) The coordinate tree array having at column i the coordinate numbers of the joint coordinates related to the kinematic expression of body i and the coordinate exponent array H, showing the number of the coordinates are constructed as follows: } (3) coordinate number of qi 4 q ( . , i ) = and WO H q ( i ) = D\u00b0Fr(k , i ) (4) k = l Table 1 shows the resultant topology data of the quadruped model. Let us denote by qi the vector of the joint coordinates whose coordinate numbers construct (.,i) and call it the relatedjoint coordinate vector of body i. 2.2 Transformation Matrix and Origin Vector For the reference coordinate system we establish the local frame denoted by !TI, at the CG of body i as shown in Fig. 2 . The inertial frame denoted by !TI is fixed to the ground. The transformation matrix from %, to X, is 4 = (.iJO441) ( 5 ) where is the constant matrix determined by the initial orientation of the bodies and R ( q , ) is made of direction cosines resulted from the successive rotation of the joint coordinates. The transformation matrix from %! to !TI is obtained recursively in the order of the size of H(i) as follows: if H(i ) = 1 if H(i ) # 1 (6) 4 = Aji 4 = The origin vector of body i, the vector from the origin of X, to that of XC, describes the relative translation of body i with respect to body j as Dot.; r.. r/ = s /, .. + CUi(k)qi(k) - Si/ = S/, + uiqi - sii (7) k=l where qi(k) denotes the k-th joint coordinate in qi. The vector U ( k ) , the k-th subvector of U i , is the direction cosine of the translational axis corresponding to qi(k) and becomes O,,, for rotational joint coordinates. The vectors si/ and sli denote the vectors from the origins of X i and Sj to joint i respectively (Refer to Fig. 2). They are obtained from the local(constant) ones as si, = AS:. I r l s . = A s ' / I / it (8) U, = A,U: 2.3 Modified Angular Velocity Transformation The angular velocity of body i can be obtained recursively in the order of size of H(i) as follows: if H(i) = 1 0, =my 0, = w, + w,, if H(i) # 1 (9) { where w,/ denotes the relative angular velocity of body i to body j . It can be expressed as Oil = Giq, where Gi = A G' J l Here the matrix G,' which depends on the type of joint i can be obtained analytically" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000869_s0957-4174(97)00057-2-Figure8-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000869_s0957-4174(97)00057-2-Figure8-1.png", + "caption": "FIGURE 8. (a) The initial robot configuration and the desired end-effector configuration. (b) The solution determined by our algorithm,", + "texts": [ + " The collision detection scheme is based on the algorithm by Lozano-Perez (1987) that efficiently corn- putes a collision-free interval for a given robot link as it rotates, keeping all other degrees of freedom fixed. These collision-free intervals are then used by the bouncing technique to convert an arbitrary Manhattan path into a collision-free Manhattan path. Several standard techniques-simplified and hierarchical representations for manipulator links and obstacles, etc.--were employed for efficient collision detection. The robot links, the fixed obstacles, and the payload are each represented by a parallelepiped. Figure 8(a) shows the initial configuration of the arm, and also the desired configuration of the end-effector. No physical obstacles are present in this example. The robot configuration found by our algorithm is shown in Fig. 8(b). Forty-two landmarks were needed to solve this example and the total execution time was 15 s. Figure 9 illustrates a difficult example with a few fixed obstacles present in the environment. The initial configuration of the arm and the desired end-effector location are shown in Fig. 9(a). It should be noted that the desired location of the end-effector is under the table and therefore, fairly constrained. Small moves of the robot must be executed to achieve this (if it is indeed possible). Fifty-eight landmarks were needed to solve this example and total execution time was 30 s" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000076_0045-7825(93)90131-g-Figure2-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000076_0045-7825(93)90131-g-Figure2-1.png", + "caption": "Fig. 2. (a) Analytical solution: Vo(M ) = 0.5 and Vo(N ) = I. (b) Vo(M ) = 0.489, Ve(N ) = 0.980.", + "texts": [ + " 2 dr (a l ) v 3 - - T ' Every component of the stress tensor is zero except o'0,. As ~0:, this component has a linear dependency on r and does not depend on z. Numerical computations were compared with this previous analytical solution. First, we consider a short cylinder (with a ratio LIR = 1) coarsely meshed with different kinds of element. The velocities of the nodes belonging to the OA and BC segments are 0 and 1/2'rr tr/s, respectively. The nodes on segment OB belong to the symmetry axis. Figure 2(a) shows the prescribed velocities and the analytical solution at points M and N. Numerical results with 4 node quadrangles obtained with 4 elements give a rather poor approximation of the solution (Vo(M) = 0.463 and Vo(N ) = 0.928). Obviously, finite element 424 A. Moal el al.. A simulation of the torsion and torsion-tension tests results are improved if we increase the elements number (16 elements) but the solution is still not very accurate (Fig. 2(b)). The analytical solution is reached with a very thin mesh. These computations were done with m = 0.5. Figure 3 shows the influence of the rheological parameter m (strain rate sensitivity) on the numerical results obtained with 4 node elements. The very low values of m introduce a great non-linearity in the rheological equation and elements using linear shape functions are unable to approach the solution. The strain rate sensitivity does not significantly influence the results obtained with quadratic elements" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002664_icsmc.2002.1173330-Figure2-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002664_icsmc.2002.1173330-Figure2-1.png", + "caption": "Fig. 2. A support polygon of a humanoid robot .- .", + "texts": [ + " where xzmp is the x coordinate of ZMP, xcog and zcog are the 1: and z coordinate of COG position respectively, g is gravity constant. B. Support Polygon To realize humanoid's stable posture, it is necessary to control the action force to COG. The area where the humanoid can receive floor reaction force is restricted to the iaside'of the polygon which consists of. sole con- - nected to the floor. This polygon is called the support where G n t e r is center position of. the support polyge nand and xupps-limit is limitaion value of the support polygon as shown in Fig.2. The equation (3) is solved with respect to xcog. - The equation (3) is transformed as follows; (4) where M is mass of the humanoid. Therefore, if the falling force moment does not occur, the relation between the action force to COG and the support polygon is shown by equation (5). fcog(x> Z ~ ( f e o g ( Z ) + M g ) ( x c o g - x \" ~ ~ r r ~ i m . t - x c e n t e r ) 1 (5) The following equation (6) c ~ l l derive similarly. . . . . .- . . . . -. ~ . -. polygon. Ip the.case ofthe huyanoid; the support poly: . . fioi(x) -(fipe(.t),~g~(xc4g-Zlour=~iimit--Zc=ntei) :'- . . -. 1 . - zcog . . . . . . . . . . . . . . . . gon-iszui area shown in'Fig.2 .If ZMP is inside of the support polygon, ZMP con&sta with --the point -.of .the -., -...I floor reaction force, But$ ZMP is outside of .theisup:- .. -. the floor.reaction force of the hum&oid;e.g., ;if ZMP is outside of the support polygon as shown in Fig. 3, a torque occurs in the circumference of a tiptoe. This torque moment is called the falliig force moment. If the - -: (6) -.-- i . . . . . . . . . . . . . . . . - - port.p.olygofi, ZMP does.not .&mi& G t h the.po&+f:''..- i .. .pPSfG -con@ " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000024_s0167-8922(08)70489-9-Figure6-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000024_s0167-8922(08)70489-9-Figure6-1.png", + "caption": "Figure 6. Isometric pressure distribution of the con-rod bearing of a passenger car engine", + "texts": [ + " Interestingly, the higher the engine speed is, the nearer the location of the damage is to the split line. after running at 6000 rpm It is obvious that the conventional rigid bearing theories are deficient in tackling these problems. However, the pinch effect discussed in the foregoing section is able to provide us with some insight of the problem. EHL simulations were thus carried out to understand the problems. The isometric pressure distribution of the passenger car engine connecting rod bearing is shown in Figure 6. This is at 360 degree crank angle, when the maximum inertial loading is produced. It can be seen that the two hydrodynamic pressure peaks are 645 generated in the regions about 20-30 degrees from the with the peak pressure at about 10 degrees from the split lines. The location of the leading pressure peak split. More interestingly, the pressures also vary correlates well with the observed position of damage markedly in the axial direction, with high values occur in the bearing (figure 4). very near to the bearing edge", + " This is of interest since the wear marks are located on the leading side of the bearing. (Engine peed: 16000 rev/min; 359 degree crank angle) (360 degree crank angle, 7000 revlmin) In Figure 7, the pressure profile of a high performance engine connecting rod bearing is presented. The diagram is at 359 degree crank angle position. This is the time when the bearing experiences its maximum peak oil film pressure, and nearly the maximum inertial load. The pressure disaibution is quite different from the one produced at lower engine speed (Figure 6). The peak pressure regions have shifted right to the bearing split line, 4.2. Sapphire seizure test The Sapphire bearing test rig is widely used by bearing manufacturers for bearing material development and production quality assurance. It is also used for tests for other purposes. One of them is to test the seizure resistance of bearing materials. In the test, an axial groove is machined on the lining close to the bearing crown in order to disrupt the formation of oil film and thus induce the occurrence of bearing seizure" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002006_1.528051-Figure4-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002006_1.528051-Figure4-1.png", + "caption": "FIG. 4. The twists (3.1), (3.2), and (3.3).", + "texts": [ + " For each I\" we define a diffeomorphism 1 [ 1\"] called the twist around I\" as follows: it is the identity in the complement of U[ 1\"] and on U[ 1\"] f[t']: ~z'exp[ -2-r=TA(lzl-l)/\u00a3)], for 1\"= aA,bA,cA , (3.1 ) 1[1\"]: ~z.exp[ _2r-TA(2Izl-\u00a3I+\u00a3) _ 2r-T A ( _ 2 Izl -\u00a31 - \u00a3 )] , for 1\"= rAh , SAh , (3.2) f[t']: ~z.exp[ --r=TA (2 Izl-: +\u00a3) (3.3 ) where A is a smooth function such that A(S) = 0 for s\";;O, A(S) = 1Tfor S> 1, and dA /ds>O. This definition agrees with the one given in Ref. 1 for 1 [ a A ] ,J [ b A ], and 1 [ C A ]. The effect of these twists on the annulus is shown in Fig. 4: the spiraling lines are the images under the twists of the intersec tion of the real axis with the annulus. For the orientation-reversing generator we take (the isotopy class of the restriction to ~g.n of) the reflection K 3, assuming that the disks D1, ... ,Dn are placed on ~g in such a way that K 3 (Dn) =Dn' Notice that the total inversion J would not work for odd n. It was proved in Ref. 3 that .n(~g.n) is generated by the following set: {/[aA],J[bA] for A = l, ... ,g; l[cA] for A = 1, ... ,g-l;/[rAh ],/[sAh] for A = 1, " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000046_cbo9780511628863.029-Figure2-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000046_cbo9780511628863.029-Figure2-1.png", + "caption": "Figure 2. Minkowski spacetime in the role of the covering space for Misner space; i and x are Lorentz coordinates that coincide with t and x in copy 0 of Misner space, but are boosted relative to t and x in all other copies.", + "texts": [ + " As Misner showed in his seminal lecture [Mis67], they pass through a chronology horizon of their own (distinct and different from that of the leftward observers), and Cambridge Books Online \u00a9 Cambridge University Press, 2010https://www.cambridge.org/core/terms. https://doi.org/10.1017/CB 9780511628863.029 Downloaded from https://www.cambridge.org/core. University of Warwick, on 18 Sep 2018 at 06:12:08, subject to the Cambridge Core terms of use, available at into a non-chronal region of their own (distinct and different from the leftward one). This pathological behavior can be understood using the covering space of Misner space [HaE73]; Figure 2. This covering space is constructed by lining up a sequence of copies of Misner space, side by side, each one boosted by speed (3 relative to the last one. The copies of the (physically irrelevant) wall are labeled \\VQ, W\\, W2, etc. in Figure 2; and each Misner space is labeled \"copy 1\", \"copy 2\", etc. A representative event P in Misner space is shown in each of the copies. There is actually an infinite number of copies of Misner space and of the wall and of the point P , with the highorder copies asymptoting to the rightward chronology horizon. The typical leftward geodesic L and typical rightward geodesic R are shown in the covering space, along with the two distinct chronology horizons through which they pass. From this covering space it should be clear that there is a complete symmetry between the leftward observers and the rightward ones", + " Cambridge Books Online \u00a9 Cambridge University Press, 2010https://www.cambridge.org/core/terms. https://doi.org/10.1017/CB 9780511628863.029 Downloaded from https://www.cambridge.org/core. University of Warwick, on 18 Sep 2018 at 06:12:08, subject to the Cambridge Core terms of use, available at Intriguingly, there is a third family of timelike geodesies and observers that are intermediate between the leftward and rightward ones. This family consists of geodesies that hit the covering space's origin (the event Q in Fig. 2); an example is the geodesic / . Note that in the covering space there is only one copy of the event Q, whereas there is an infinite number of copies of every other event, e.g. P. This presumably is related to the following remarkable pathology. Although the event Q exists in the covering space, it does not exist in Misner space. There is no way to include it in the spacetime, if one insists that the spacetime be a manifold and one includes the chronal region, and both the left and right chronology horizons, and both the left and right non-chronal regions", + " Those regions cannot be meshed smoothly with Q; and the impossibility of meshing makes Misner space geodesically incomplete: the intermediate geodesies (e.g. / ) all terminate after finite proper time just before reaching the non-existent event Q. As pathological as this may seem, it would have been much more pathological if the terminating geodesies were not a set of measure zero. The intermediate geodesies can be used as the time lines for a coordinate system that treats leftward and rightward geodesies on an equal footing. This coordinate system (T, X, y, z) is related to the Lorentz coordinates (\u00a3, x,y, z) of the covering space (Figure 2) by x = TsinhX; (1) and correspondingly, the metric in this coordinate system is ds2 = -dT2 + T2dX2 + dy2 + dz2 . (2) The boost-related points that are identified to produce Misner space (e.g. the points P in Figure 2) all are on the same hyperboloid i2 - x2 = constant, and therefore are all at the same T, y, z, but different X. The boost that takes one of the P's into the next one is simply a displacement in X by tanh\"1 (3. Therefore, the n'th copy of P is at Xn = XQ + n tanh\"1 /?, and Misner space can be regarded as the space of Equation (2) with X periodic with period tanh\"1 /?. As seen by the intermediate observers, who sit at fixed (X, y, z), the leftward observers circle leftward around Misner space an infinite number of times as they approach their chronology horizon, and similarly for the rightward observers", + " Put differently, Grant space is Minkowski spacetime, closed up in the redirection, with a displacement Ay = a of all rc-directed geodesies when they travel around the space once, and with the space contracting along its x-direction at a rate ^(circumference)/^ = \u2014/?. It should be clear that the lateral displacement Ay = a does not alter the existence of two families of geodesies in spacetime, rightward and leftward, nor of the leftward and rightward chronology horizons and leftward and rightward non-chronal regions, as depicted in Figures lb and 2. What is changed in Figure 2 is that the identified points P in the covering space's successive copies of Grant space are displaced relative to each other by an amount Ay = a into the paper. Correspondingly, the generators of the leftward (or rightward) chronology horizon (null geodesies traveling in the redirection) do not close on themselves; rather, they begin &ti = x = y = \u2014 oo and travel rightward to t = x = 0, y = +oc, where they leave the chronology horizon. Moreover, the chronology horizon possesses no fountains (no smoothly closed null geodesies from which generators spring); and no closed timelike curve passes through any event on the chronology horizon" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001136_s0020-7403(98)00106-4-Figure7-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001136_s0020-7403(98)00106-4-Figure7-1.png", + "caption": "Fig. 7. Lashing spar with associated mass.", + "texts": [ + "g m(5)sin a)/m(5), where l 4 is the coe$cient of friction between the tyres and the deck. The angular acceleration (i.e. roll) of the suspension mass is hG (5)\"(\u00b9(5)#Fz(5)ZG(5)sin h(5)#FSR(BR(5)cos h(5) !FS\u00b8(B\u00b8(5)cos h(5)!(ZG(5)#Z\u00b8(5))sin h(5)) #(ZG(5)#Z\u00b8(5))sin h(5)!Fy(5)ZG(5)cos h(5) !(FS\u00b8#FSR)z(5)k s )/I(5). The similar set of equations can be set up for the trestle mass, but in this case j is zero and the friction acts between the top of the trestle and the underside of the chassis. A lashing spar, with its possible orientation, is shown in Fig. 7. Each of the four spars, which are assumed to be rigid, can rotate relative to the spine and this rotation, plus the #exing of the spine relative to the deck give rise to forces in the lashings. As the de#ections of the chassis are much smaller than the pitch and yaw of the trailer, and as these are also small, the values of / and t, shown in Fig. 6, can be approximated to by using Fig. 5 to give /\"(ZS(6)!ZS(1))/(X(6)!X(1)), t\"(>S(6)!>S(1))/(X(6)!X(1)). These values along with the positions of the centres of mass are used to calculate the position of the ends of the rigid spars and thus C\u00b8(i), etc", + " The components of the lashing forces in the x, y and z directions are proportional to the variables C\u00b8(i), S\u00b8(i) and \u00b9\u00b8(i) for the left-hand lashings and for the right-hand lashings CR(i), SR(i) and \u00b9R(i). The length of the left-hand lashing is given by \u00b8\u00b8(i)2\"S\u00b8(i)2#C\u00b8(i)2#\u00b9\u00b8(i)2 and the force is given by F\u00b8(i)\"(\u00b8\u00b8(i)!R\u00b8(i))ZSz, where R\u00b8(i) is the unstrained length of the lashing. The components of the force F\u00b8(i) in the x, y and z directions are F\u00b8x(i)\" F\u00b8(i)C\u00b8(i) \u00b8\u00b8(i) , etc. The values of the forces acting in the x, y and z directions on the right-hand side of the spine are calculated in a similar manner. Using Fig. 7 it can be seen that the vertical acceleration of the spar mass is given by z( (i)\"(Fz(i)!F\u00b8z(i)!FRz(i))/m(i) and the horizontal acceleration by y( (i)\"(Fy(i)#FRy(i)!F\u00b8y(i))/m(i). The angular acceleration of the mass is h$ (i)\"(Fz(i)ZG(i)sin h(i)#Fy(i)ZG(i)cos h(i) !FRy(i)((ZG(i)#Z\u00b8(i))cos h(i)#B\u00b8(i)sin h(i)) #FRy(i)((ZG(i)#Z\u00b8(i))cos h(i)!BR(i)sin h(i)) #FRz(i)((ZG(i)#Z\u00b8(i))sin h(i)!BR(i)cos h(i)) !F\u00b8z(i)((ZG(i)#Z\u00b8(i))sin h(i)!B\u00b8(i)cos h(i) #\u00b9(i#1)!\u00b9(i))/I(i). As was the case with the other models described in Refs" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0000727_s0167-8922(08)70486-3-Figure2-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0000727_s0167-8922(08)70486-3-Figure2-1.png", + "caption": "Figure 2 Equivalent geometry [3]", + "texts": [ + " Notation U B b semi-major axis of Hertzian contact ellipse spin to roll ratio, 2 0 1 LI semi-minor axis of Hertzian contact ellipse coefficient of minimum film reduction equivalent elastic modulus normal load non-dimensional materials parameter, aE ' non-dimensional film thickness, h / R x film thickness nominal distance of two bodies central film thickness prediction by Hamrock and Dowson [2] minimum film thickness prediction by Hamrock and Dowson [2] minimum film thickness prediction from present study ellipticity ratio of the contact, u / b non-dimensional pressure, p / E' pressure in lubricant film (gauge pressure) asymptotic isoviscous pressure maximum Hertzian pressure principal radii of curvature of equivalent ellipsoid separation between two bodies non-dimensional speed parameter, tlouo/E 'Rx mean velocity component in the x direction mean velocity component in they direction non-dimensional load parameter, F,/ E 'R: elastic deformation non-dimensional Cartesian coordinate, x/ b Cartesian coordinate (rolling direction, minor axis of contact ellipse) non-dimensional Cartesian coordinate, y / a Cartesian coordinate 1 /PI\",rn relaxation factor for SOR method viscosity tl/ 70 viscosity at atmospheric pressure fluid density P/Po fluid density at atmospheric pressure non-dimensional spin angular velocity, (mean) angular velocity of spin W V O / E ' 2. GOVERNING EQUATIONS 2.1. Geometry The contact between an equivalent ellipsoid and a plane shown in Figure 2 was studied, where the equivalent elhpsoid has a pair of principal radii of curvature (Rx, R,,). 60 1 Both surfaces were assumed to be smooth, so that the effect of surface asperities was neglected. The separation in the z direction between two rigid bodies which contact at the origin 0 can be expressed, with paraboloidal approximation, as; x2 y2 s(x,y) = - + - 2R, 2Ry Many machine elements, such as a ball bearing, have elastic contacts in which the direction of rolling or sliding motion coincides with the minor axis (b) of the Hertzian ellipse" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0002371_iros.1991.174511-Figure4-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0002371_iros.1991.174511-Figure4-1.png", + "caption": "Figure 4: Relationship of real space and configuration space friction cones.", + "texts": [ + " The cone is a two-dimensional planar subset of the threedimensional generalized force space. In order to build some intuition. let observe that the configuration space normal models the direction of a real space normal reaction force and its induced reaction torque. This is clear from the cross-product term in the angular component nq of the normal n. Similarly, the edges of the friction cone model the added direction of a real space tangential reaction force and its induced torque. This is apparent from the cross-prodat term in the angular component of the vector vf , See Figure 4. 4. Ambiguities Consider now the generalized friction cone visualized in its plane of residence. See Figure 5. The normal to this plane is given by a configuration space tangent vector that models pure rotation about the contact point. This is because pure rotations about the contact point do not require constraint forces. Let us therefore define t, to be the unit tangent vector that represents pure counterclockwise rotation about the contact point. Note that t, is parallel to the vector ( -ry ,rr , p)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0003950_icems.2005.202484-Figure2-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0003950_icems.2005.202484-Figure2-1.png", + "caption": "Fig. 2. A rotor bar with pear-shape geometry", + "texts": [ + " From Table 1, the results by two methods, a presented multi-layer method and an analytical method, are matched well each other and the relatives errors are approximately equal to zero, which verifies the accuracy of the multilayer method presented in the paper. The analytical formulas are only for rectangular bar, for a slot and bar with more general shape, there exists no analytical method. Multi-layer method presented in the paper is used to solve this problem and the results can be verified by the finite element method. A rotor bar with pear-shape is shown in Fig. 2, which is embedded in a semi-closed slot. The geometrical dimensions of the bar are also marked with the units of cm in Fig. 2. The resistivity of cast aluminum rotor bar is 4.34x10-8 Qm. By the presented method, the magnetic lines of flux are assumed to be in parallel to the slot top line and each layer to be a rectangular bar. For the convenience of obtaining width and height of each thin-layer rectangular bar, the whole rotor bar is firstly divided into 4 sections with semi-round shape, trapezoidal shape, semi-round shape and rectangular shape, respectively. Thin-layer rectangular bars are then sliced from 4 sections" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_65_0001550_s0039-6028(00)00374-5-Figure1-1.png", + "original_path": "designv11-65/openalex_figure/designv11_65_0001550_s0039-6028(00)00374-5-Figure1-1.png", + "caption": "Fig. 1. (a) A block with mass M sliding on a flat substrate. A spring (spring constant ks) is connected to the block and the \u2018free\u2019 end of the block moves with the velocity vs. (b) A block on a tilted substrate.", + "texts": [ + " One of the central topics in sliding friction is to understand the origin Understanding the origin of friction is crucial of the transition from steady sliding at \u2018high\u2019 to modern technologies such as hard disk drives sliding velocities, say vs>vc, to stick\u2013slip motion [1], microelectromechanical systems [2], and other at \u2018low\u2019 sliding velocities (v