diff --git "a/designv11-69.json" "b/designv11-69.json" new file mode 100644--- /dev/null +++ "b/designv11-69.json" @@ -0,0 +1,9580 @@ +[ + { + "image_filename": "designv11_69_0002576_nems.2007.352170-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002576_nems.2007.352170-Figure2-1.png", + "caption": "Figure 2. A. 3x3 fluorescence detection \u201cchip\u201d B. Magnification of one excitation/detection unit prior to self-assembly process.", + "texts": [ + " The fluorescent label absorbs this energy and releases it again at a lower energy level. This emission energy is captured by the photodetectors below. Our integrated fluorescence detection chip is currently a 3 x 3 array of individually controllable excitation/detection units. Each unit is made up of a single circular excitation source, which is an AlGaAs LED, surrounded by 8 square pn-junction photodetectors. The shape differences of the elements allows the self-assembly process to utilize shape recognition to guide the elements to their final locations. Fig. 2 shows the circular and square receptor site locations prior to self-assembly. The template substrate can be any material which can withstand typical microfabrication processes. These materials include glass, plastics [5], ceramics, and so on. We have chosen glass as our substrate because of its chemical compatibility and ease of use during fabrication. The size of the glass substrate was chosen to be the same as a 1\u201dx3\u201d glass microscope slide for easy integration with existing instrumentation and for its familiarity to researchers in biology and chemistry disciplines" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000737_00368790410558239-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000737_00368790410558239-Figure1-1.png", + "caption": "Figure 1 Radial bearing experimental apparatus", + "texts": [ + " In addition, the minimum oil film thickness can be theoretically as follows: hmin \u00bc cd 2 \u00fe 2:35 So \u00f03\u00de where d is the hub diameter, and the Sommerfeld number: The analysis of the effects of surface roughness of shafts Cem Sinanog\u0306lu Industrial Lubrication and Tribology Volume 56 \u00b7 Number 6 \u00b7 2004 \u00b7 324\u2013333 D ow nl oa de d by G eo rg e M as on U ni ve rs ity A t 0 0: 36 1 2 M ar ch 2 01 6 (P T ) So \u00bc Pc2 h0v \u00f04\u00de where c \u00bc d=d; d \u00bc \u00f0D 2 d\u00de; h0 is kinematics viscosity of oil and v is the angular velocity. The Sommerfeld number is used to give information about the stability or instability of the shaft. The radial bearing used in the experiments is a model journal bearing (Figure 1). This bearing is suitable for demonstrations of the effect of the more important variables such as speed, viscosity and load on the pressure distribution in a journal bearing. The bearing is manufactured from clear perspex, thus making the oil film profile clearly visible when viewed from the front of the assembly. The pressure distribution of the oil film can be compared with the Sommerfeld function. The Sommerfeld function is given as (Kurban and Sinanog\u0306lu, 2000); DP \u00bc P 2 Ps \u00bc 2 6mr2v d2\u00f02 \u00fe 12\u00de 1 sin u \u00f02 \u00fe 1 cos u\u00de \u00f01 \u00fe 1 cos u\u00de2 \u00f05\u00de where P is the pressure value on the film oil at u angular displacement (clockwise), Ps the static pressure in hydrodynamic lubrication and in the experiment it was kept constant at 735 mm-oil, d the radial clearance, 1 the relative eccentricity, r the shaft radius, v the angular velocity and m the dynamic viscosity of oil", + " Different rotational speeds (1,250, 1,750 and 2,000 rpm) were employed. Mobil 0W-40 synthetic oil with a viscosity of 864 kg/m3 was used as lubricant. Relative clearance of the bearing as c \u00bc 1:5 \u00a3 1023 is selected for light metal bearing, bearing-shaft systems are treated at H7/e8 tolerance. Diameter of hub is 54.8 mm, diameter of the bearing bush is 55 mm, bearing width is 70 mm and weight of bearing is 650 g. The pressure variations were recorded with the 16 tubes and the results were plotted. The radial bearing shown in Figure 1 in this section essentially consists of a clear perspex journal bearing mounted freely on a steel journal shaft (A). The large diameter journal shaft is directly fixed onto a motor shaft. The speed of the The analysis of the effects of surface roughness of shafts Cem Sinanog\u0306lu Industrial Lubrication and Tribology Volume 56 \u00b7 Number 6 \u00b7 2004 \u00b7 324\u2013333 D ow nl oa de d by G eo rg e M as on U ni ve rs ity A t 0 0: 36 1 2 M ar ch 2 01 6 (P T ) motor shaft (B) is accurately controlled by the standard tecquipment control unit, which is mounted within and in front of the main framework", + " The required loads were added on to the shaft, at the bottom of and then a angular displacement was formed in the bearing. When the manometer levels were settled down, the pressure reading on 16 manometers were taken. Initially, oil tank was fixed at 735 mm levels (oil supply head Ps \u00bc 735 mm) and therefore, in graphics P the reading pressure, and Ps the static pressure. The positive pressure difference values correspond to local bearing load capacity. The pressure, which is constant due to axial direction, indicated by 1, 2,. . .,5 tubes are placed along the bearing axis. At the experimental work (Figure 1), the masses (H) on the (J) shaft can be placed on different positions. The pressure values on the journal bearing were measured from 3-6-7- 8-9-10-11-12-13-14-15 and 16 th tubes. The tube number also indicates the angles, i.e. 3 (0-3608), 6 (308), 7 (608), 8 (908), 10 (1508), 11 (1808), 12 (2108), 13 (2408), 14 (2708), 15 (3008) and 16 (3308). The variations of pressure are shown in Figure 4 (Case 1) versus the angular position for different angular velocity values, i.e. 1,250, 1,750 and 2,000 rpm (steel shaft)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003379_j.camwa.2010.09.024-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003379_j.camwa.2010.09.024-Figure1-1.png", + "caption": "Fig. 1. A nonholonomic system with friction.", + "texts": [ + " The basic problem solving for constraint forces may come down to the problem of the unique solution of the LCP. By LCP theory, we can obtain the following conclusion. For each\u039b, only when B\u2217 is a positive matrix, the system has a unique solution solving for constraint reaction forces, and the unique solution is \u03bb = B\u22121 1 \u039b or \u03bb = \u2212B\u22121 2 \u039b. When B\u2217 has non-positive principal minors, there may exist a unique solution for some \u039b, but there also exist no finite solutions or multiple solutions for other \u039b, which is called singularity. As an application, we consider the system as shown in Fig. 1. The slider A moves along a rough slideway and the ball B and the slider A are hinged by a massless rigid rod AB. The lengths of the two rods are expressed as 2l. Let (x, y) be the coordinate of the sliderA, \u03b81 and \u03b82 be the angles between the two rods and the vertical line, respectively. The system is controlled by cos \u03b81x\u0307A + l\u03b8\u03072 = 0. Now we will show the process obtaining the LCP model for a concrete example as follows. Since Coulomb friction at the sliderA is considered, we select (xA, yA, \u03b81, \u03b82) as the generalized coordinates of the system" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003779_gt2010-22877-Figure8-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003779_gt2010-22877-Figure8-1.png", + "caption": "Figure 8 shows two-dimensional streamline distribution and the static pressure contours of the labyrinth seal. The unit of the pressure contours in Figures 8-12 is Pa. The static pressure value decreases step by step from the inlet to the outlet in the referenced labyrinth seal. The leakage fluid passes the clearance between the first fin and rotor surface and impinges on the step of the rotor surface. The leakage jet is then deflected and directed toward the wall of the cavity and formed an anti-clockwise recirculation zone. The leakage fluid flows across the gap between the two fins and the rotor surface and impinges on the second long fin. The jet is deflected and divided into two recirculation zones due to being directed toward the rotor surface and the bottom of the labyrinth chamber. One clockwise and one anti-clockwise recirculation region are generated in the labyrinth chamber. A similar flow pattern is observed in the sequence cavities. Due to this flow pattern, the kinetic energy of the leakage flow is dissipated into heat energy.", + "texts": [ + " In the flow direction, boundary conditions at the flow inlet and exit boundaries were needed. The inlet boundary was placed at the seal entrance, and total pressure, total temperature and turbulence quantities were defined, while the averaged static pressure was specified at the outlet of the seal. The stationary walls were defined to be adiabatic. The rotor surface is assumed adiabatic, which rotates at a speed of 3000rpm. As to the brush seal, the non-Darcian porous-medium model was utilized for bristle packs. Fig. 8 Static Pressure contours and streamline distribution of the referenced labyrinth seal The first long fin is retained and the subsequent three long fins are redesigned into the bristle pack based on the referenced labyrinth seal. The leakage flow characteristics of the brush seals 1 and 2 with four sealing clearances were investigated. The local flow field in the first bristle pack region with different sealing gaps for brush seals 1 and 2 is discussed because of the similar flow patterns as in sequence bristle packs" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002054_bfb0119386-Figure6-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002054_bfb0119386-Figure6-1.png", + "caption": "Fig. 6: Reference objects in test bed", + "texts": [ + " Due to temperature influence, the ultrasonic sensors have problems with their measuring accuracy because of the changing density of the air. All sensor principles were theoretically examined in order to determine if smoke or heat bias the function or accuracy of the sensor and to proof that they are suitable for the application. Furthermore the difficulty of data processing was checked. After theoretical analysis, several sensors were chosen to create a concept for the sensor system and for extensive live tests. 4 Experimental validation 4.1 Test bed An experiment container (fig. 6, fig. 7) with the length of 12 meters was built for the tests. Objects with different materials, geometry and surfaces were located in the test bed in order to simulate the conditions of the reference task (fig. 4). The sensors were installed in a sealed box which is moveable and rotatable to simulate the robot's movement. Different smoke densities (according to the density classes - fig. 2) and different temperatures were generated during the test. The data of the sensors were recorded considering the test objects, their characteristics and the different environmental conditions" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000966_978-1-4419-8887-4_6-Figure6.10-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000966_978-1-4419-8887-4_6-Figure6.10-1.png", + "caption": "Figure 6.10. British Transport and Road Research skid resistance tester.", + "texts": [ + " Once this limit has been exceeded and sliding occurs, the force required to maintain movement is lower than that needed to initiate motion. It is easier to keep a sliding object moving than it is to start sliding. A mechanical test method used to measure translational friction is the British Transport and Road Research Laboratory portable skid resistance tester. A standard rubber foot is attached to the end of a pendulum. The pendulum is released from a horizontal position to slide over the court surface. A friction coefficient is determined by the maximum height attained by the foot following sliding over the surface (Fig. 6.10). By using a standard rubber foot rather than samples of typical shoe soles, this test does not provide specific friction values for the actual conditions occurring on the court surface during play. It does, however, provide a standard test procedure for comparison of different surfaces. Chap. 6. Shoe-Surface Interaction in Tennis 139 The rotational friction is influenced by the properties of the court surface and the shoe sole and also by the area of contact between the shoe and surface. Rotational friction can be quantified using a test device that measures the resistance to an applied constant torque (Stuttgart device), or the measurement of the torque required to rotate a weighted foot from a stationary position (BS7044)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000477_s1474-6670(17)37108-2-Figure2.2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000477_s1474-6670(17)37108-2-Figure2.2-1.png", + "caption": "Figure 2.2 Sail Forces Analysis", + "texts": [ + ", 1979) since they were initially investigated by Astrom and Kallstrom, (1967). Autopilot research for smaller motor-powered marine vessels has been undertaken more recently by Vaneck, (1997), Kose and Gosine, (1995) and Polkinghome, et al., (1994), but it appears that adaptive control of yacht motion has so far received little attention. The hydrodynamics of a yacht are determined by its special structure, comprising its main forcing elements: sails, hull, keel and rudder. A typical yacht is shown in Figure 2.1. Figure 2.2 presents the forces acting on a yacht that is beating into the wind. Here Flirt is the lifting force on the yacht due to the wind on the sail, Fdrag is the drag force of sails passing through the wind. and Fs is the sum of Flirt and Fdrag . The coordinate system is defined by having x lying along the heading direction. The symbols used are defined as follows: J:F;,. J:Fy. I:Fb: total driving force, side force and venical buoyancy force; I.Mx. I.My. I.Mz: total heeling moment, pitching moment and yawing moment; M~: damping moment determined by heeling and yawing motion; u, v" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002979_978-3-540-70534-5_7-Figure14.8-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002979_978-3-540-70534-5_7-Figure14.8-1.png", + "caption": "Figure 14.8: Manipulator simulator RoboSim", + "texts": [ + " Unfortunately, for any given manipulator, there is no general or simple way of deriving the inverse kinematics. The algebraic solution for the Puma 560 goes over five pages in [Craig 2003] and there are multiple solutions for most goal poses as the manipulator does have some mechanical redundancies. The use of numeric approximations may be an alternative for this. 14.3 Simulation and Programming There are a number of robot manipulator simulators available either as public domain systems (e.g. RoboSim in Figure 14.8, [Br\u00e4unl 1999]) or as commercial products. All manipulator manufacturers provide simulation systems for application planning with their products and in many cases specialized programming environments as well. After simulation, the first step in programming a real robot is a process called \u201cteaching,\u201d in which the operator manually drives the robot in a certain configuration and then stores this pose by pressing a button on its controls. These poses are then subsequently used as reference points in a robot program, e" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002573_isie.2008.4676902-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002573_isie.2008.4676902-Figure1-1.png", + "caption": "Figure 1: Comparison between the predicted current errors for both DCC approaches (superscripts I and II denote the approach, respectively).", + "texts": [ + "00 '2008 IEEE After impressing an active voltage vector (any of the six possible active vectors in common three phase inverter): 1 2 2 3 2 1 1 3 3 3 1 1 0 3 3 d d DC DC q q s v K V V s v K s \u2212 \u2212 = = \u22c5 \u2212 C (9) where s1, s2 and s3 are the corresponding transistors\u2019 states (0 or 1) and C2 is the Park ab-dq transformational matrix [1], the predicted current error components at the end of the sampling interval are * 0( 1) ( 1) ( 1) ( )d d d d d t n i n i n v n L \u03b5 \u2206 + = + \u2212 + \u2212 (10) * 0( 1) ( 1) ( 1) ( )q q q q q t n i n i n v n L \u03b5 \u2206 + = + \u2212 + \u2212 . (11) Both predicted errors (7), (8) and (10), (11) have to be calculated and their magnitudes compared. More practical criterion derived form (7-11), when an active voltage vector should be applied, is: ( )2 2 0 0 2 qd DC d q d d q q q d q d LL V K t K t K L K L L L \u03b5 \u03b5 \u2206 + \u2206 < + . (12) An example, shown in Fig. 1, indicates current error if zero voltage vector is applied (\u03b5\u0399 0) or, if active voltage vector is applied (\u03b5I V for v3) . In this particular case, the magnitude of predicted current error for a zero voltage vector (\u03b5\u0399 0) is smaller than the magnitude of predicted error for an active voltage vector (\u03b5I V), therefore a zero voltage vector should be selected. The second DCC approach requires impressing an active voltage vector only for a subinterval ton within the sampling interval \u2206t; a zero voltage vector is impressed for the remaining time of the sampling interval", + " Thus, the predicted current error components at the end of the sampling interval can be written as: * 0( 1) ( 1) ( 1) ( ) on d d d d d t n i n i n v n L \u03b5 + = + \u2212 + \u2212 (13) * 0( 1) ( 1) ( 1) ( ) on q q q q q t n i n i n v n L \u03b5 + = + \u2212 + \u2212 (14) The expression for the duration of the \"active\" subinterval ton is obtained by minimizing this error (13), (14): ( )0 0 2 2 d d q q q d on qd d q DC q d K L K L t LL K K V L L \u03b5 \u03b5+ = + (15) If the calculated ton exceeds the sampling interval \u2206t, then ton is set to \u2206t. For this DCC approach, minimized current error in Fig. 1 is shown with superscript II (\u03b5\u0399\u0399 V), where the active voltage vector v3 is impressed for ton only. The voltage vector drives the final current error vector orthogonally to the applied voltage vector. In comparison with the first DCC variant, a lower current ripple and increased average commutation frequency are to be expected. In case when the drive has no reluctance (Ld = Lq), both terms (12) and (15) can be further simplified and become very similar to those derived in [6, 10]. As it has been already pointed out in [9], the finite calculation time of the algorithm deteriorates the performance of the predictive methods" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000250_eeic.1991.162566-Figure11-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000250_eeic.1991.162566-Figure11-1.png", + "caption": "Figure 11. Continuous Wedge Sensor MountinglWiring.", + "texts": [ + " Maintenance costs must be insignificant and due only to non-installed components such as computers. Turboscanner operating costs must be minimized by simplifying the training of Utility personnel in use and data interpretation. With regards to generator safety, the most important sensor modules arrayed along the rotor to be mounted in cavities in the back the relatively thick field coil wedges (without compromising the structural integrity of any wedge) with the sensors viewing the stator through small apertures in the wedge faces; Figure 11 illustrates this detail for an indirectly-cooled rotor; (2) ( 3 ) (4) ( 3 ) aspect of turboscanner design, it will be necessary for: ( 1) 44 I , interconnecting wiring to be routed through or along the backs of wedges to the retaining ring region at the nondriven end (NDE) of the rotor (segmented-wedge rotors will require suitable, in-line, frictional connectors); Figure 11 illustrates this detail; support electronics to occupy unused space beneath the NDE retaining ring (by hollowing out blocks and spacers); Figure 12 illustrates this detail for a two-pole rotor; electronics power to be derived from field power through the use of a power conditioner connected to the incoming bus through a mutual protection network; Figure 12 illustrates this detail; and Turboscanner on-rotor subsystems to be spatially distributed on the rotor to preserve rotor dynamic balance. (2-Pole Rotor) Turbogenerators are nearly always used for base-load generation and are expected to run for years at a time", + " To illustrate this point, the following describes the turboscanner at the conceptual design level. While it likely that the turboscanners will evolve significantly in reaching commercial status, it is equally likely that the basic approach will be preserved. Figure 12 is a simplified schematic representation of the NDE end of a two-pole turbogenerator rotor showing the generic features of such rotors and the locations of the principal rotormounted subsystems of a turboscanner. As shown in the figure (exaggerated) and in Figure 11, sensors will be mounted behind viewing apertures along four rows of wedges most likely in azimuthally-symmetric orientation with respect to the poles. Four sets of sensors (possibly two) will be provided for reasons of redundancy and mechanical balance. The power conditioner and telemetry transmitter will be mounted in vacant (or static) space beneath the retaining ring. Such space, for example, may be that occupied by centering blocks that maintain axial symmetry of the field coils with respect to the rotor; such blocks are normally either epoxyimpregnated fiber glass or die-cast aluminum" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001313_bfb0004264-Figure3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001313_bfb0004264-Figure3-1.png", + "caption": "Figure 3.", + "texts": [ + " Along the focal surface paths, then: and hence and, of course a_ (%) = - ~ H / ~ s d t a (cos \u00a5) = _I + ,,(w~s 2 + r2)cos%/ _ 2w3s2cos ~/cos dt s w2s 2 - r 2 r(w2s 2 - r 2) (6) (7) 3. N U M E R I C A L RESULTS Units of distance and time have been chosen so that the cage radius as well as Man's speed arc unity. Backwards integration of (6), together with (7) and i = w cos ~, ~ = cos ~, from capture: r / = 1 , s / = 1 for various ~ f between 0 and ~f2, determines the focal surface paths. These are shown in Figure 3 for w = 1.2. The paths terminate (in backward time) either at s = 1 with cos V < 0 or else at cos W < + 1 where x0 vanishes so that the gradient of x becomes continuous. Figure 3 also shows the radial motion tributaries arriving tangentially at the terminations with cos V = + I. Note that if Man starts on the perimeter (s = I) with Lion in a radial position with r > .3 (approximately), Man moves inside the perimeter before being captured on the perimeter. I f 0 < r < .3 (approximately), Man starts by running radially inward towards Lion. It is interesting to note that ~) and V are continuous at termination with cos ~ = + 1 although ~ and ~/ are infinite. A similar situation arose in the backwards construction of a focal surface in a different problem [1], point-capture of two evaders in succession", + " 75 Here a single exceptional focal surface path reaches r = 0 in backward time with s = s 0, V=Tt/2, in fact with cos v=r / (3WSo) for small r , where s o depends on w. For w = 1.2. s o -_- .835, but this cannot be obtained directly from the backward integration which is unstable near r = 0; instead, forward integration must be used starting with small positive r . s = s o and cos V = rl(3wso), s o being adjusted so that r and s reach unity simultaneously. Note that x, and x 0 both vanish at the point r = 0. s = s o and Man's arrival direction there is nonunique. As (barely) indicated by the z-contours included in Figure 3, the value s -- s 0 is the optimal radial distance of Man when Lion is at the center (r = 0). Figure 4 shows the curved paths of both players after Lion passes tluough the center, Man having chosen counterclockwise motion which, as also illustrated, he can reverse at any time. The computation of the time-to-go function x(r, s , O) is straightforward using straight line motions by both players arriving tangentially at various positions on the focal surface. The x-contours for sections r = 1, .75, " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003718_ipec.2010.5544517-Figure3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003718_ipec.2010.5544517-Figure3-1.png", + "caption": "Fig. 3. Cross session diagram of the six-phase PMSG.", + "texts": [ + "00 \u00a92010 IEEE Pulsation of the DC current and electromagnetic torque is produced by harmonic stator currents in the conventional generator system. This demands large smooth inductance and implies that torsional torque may arise in mechanical coupling. We proposed the novel generator system shown in Fig. 2 to solve those problems [4]. 1) Configuration: The configuration of the proposed generator system is the same as that of the conventional generator system except for the stator windings and the number of three-phase diode bridge rectifiers. As shown in Fig. 3, there are two sets of three-phase stator windings spatially shifted by 30 electrical degrees. The neutral points of stators1 (a1, b1, c1) and stator2 (a2, b2, c2) are isolated. The two diode bridge rectifiers (DB1, DB2) are connected in series rather than in parallel because the parallel connection requires interphase reactors to produce balanced currents. In the series connection, no interphase reactors are needed and it is easy to achieve a higher DC voltage. 2) Method for reducing pulsation of the electromagnetic torque: The two three-phase diode bridge rectifiers are connected to two sets of three-phase stator windings so that stator currents become a rectangular waveform if the overlap of currents is ignored" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001293_rob.20099-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001293_rob.20099-Figure2-1.png", + "caption": "Figure 2. A LOS-based sensing module.", + "texts": [ + " For example, while interferometers may have as low as nanometer accuracy, they cannot as easily determine changes in orientation. Triangulation sensors may achieve micron accuracy only in ranges of tens of millimeters. Thus, neither sensor is suitable for even limited workspace vehicle motions in multiple dimensions. The LOS task-space sensing system proposed herein consists of individual sensing modules that can be configured into a multi-LOS system to provide sufficient and accurate data for guidance-based motion planning of autonomous vehicles. One such individual LOS sensing module is depicted in Figure 2: It consists of a detector e.g., PSD attached to the target, a galvanometer mirror, and a laser source. The laser beam is positioned, using the galvanometer mirror, to hit the center of the detector at the target\u2019s ideal desired pose. The laser beam, hence, defines the desired LOS. In practice, after the target is positioned at its desired pose, the LOS would hit the detector at an offset from its center due to systematic errors, regardless of the distance moved. As such, the offset noted could be used as a feedback for the iterative guidance e" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001591_acc.2006.1657414-Figure4-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001591_acc.2006.1657414-Figure4-1.png", + "caption": "Figure 4: Simulated Transmission Line Experimental Testbed", + "texts": [ + " The second abnormal conductor has 9 cuts and 4 semi-cuts on left-hand side about one foot away from the center (Major abnormal). In order to obtain comparative data sets, the EMAT system was mounted and clamped at the end of armor rod during data acquisition. The first 228 feature vectors (Normal: 63, Minor Abnormal: 66, Major Abnormal: 66, and Corrosion: 33) are used to train the classifiers and the other 116 data sets are used for validation of the classifiers (Normal: 32, Minor Abnormal: 34, Major Abnormal: 33, and Corrosion: 17). The experimental testbed for transmission line setup is depicted in Figure 4. The results of classification are summarized in Table 1. As shown in the table, the PCA-ART network showed the best results in classifications. Acceptable results are also obtained using PCA feature extractor and LDA extractor in PNN network. This illustrates that the feature extractors based on such linear transformations as PCA and LDA did work better than those of nonlinear mapping using MLP or kernel PCA extractor for this test. These nonlinear mapping methods (MLP and kernel PCA) performed better by using the MLP classifier, which can find nonlinear boundaries among classes, but the classification was not satisfactory" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000512_810105-Figure9-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000512_810105-Figure9-1.png", + "caption": "Fig. 9 - Original point setting in x-axis", + "texts": [ + " The positioning accuracy of the machine is 1 \u0302 m and the angular accuracy of the rotational table is 2 sec. The enlarged view of the apparatus in operation is shown in Fig. 6, where we see the probe detecting the deviation of the tooth surface in y direction. The probe is held rigidly in the lateral direction (x direction) to prevent it from deflection when placed on the inclined tooth surfaces. It has been confirmed that the lateral force of 50 grw causes only 1 /im deflection as shown in Fig. 7. The setting device of the gear is shown in Fig. 8 and Fig. 9. It has a setting plane and a setting ball which are used to determine the original point, 0, of the measurement. The x and z coordinates of the center of the ball, Xe and Ze, and the y coordinate of the setting plane, Ye, have been measured in advance. At first the y-setting plane of the device is set parallel to the x axis. The x coordinates Xi and X2 are measured from an arbitrary point for two positions of the probe which give the same reading of the deflection of the probe in y direction. As seen in Fig. 9 the x coordinate of the original point, Xs, is obtained as Xs = Xe + Xi + Xs The y and z coordinates of the original point can be determined in a similar manner. Fig. 5 - Measuring apparatus 810105 The variation in the position of setting the original point has proved to be less than 4/im as shown in Table 1. A test gear is fixed by the thin walled cylinder expanded by fluid pressure.so as to: give a very small rotational error of less than.3.#im without clearance. In the case of measurement of the pinion similar method is used to determine the original point except that some modification of the setting device was necessary for the shaft of the pinion" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003165_1.1726106-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003165_1.1726106-Figure1-1.png", + "caption": "FIG. 1. Best simultaneous solution of Eqs. (1-3). (1) Eq. (1), a>O; (2) Eq. (2) a>O; (3) Eq. (3) a\" =7 =\nR \" A ax + c ~ + A + = - - Ox : -OT ~- ax\"- '\nat\" R ~ -a~ c-ax)+\nF C O~U' 8\"(-)i \u00a7 B ax ~ . (6)", + "Here/3 = 0.1-0.4 is the coefficient of dry friction. A special case of Eqs. (6) is the familiar differential equation of motion of an ordinary filament slipping over a pulley [5],\nq /a2u -\\ u fl _ .Ou __a~u\nwhere EF is the longitudinal r ig id i ty of the f i lament ; u(x, t) is the longitudinal s t r a in of the f i lament .\n9 Expres s ions (6) toge ther with the condit ions\nP i ( l i , t) = P e s t \"t- [ai d , M~ (lt. t) = Mes ' + Mtd,\nPl (lt - - Ale, t) = Pist, 1He (It - - AI~, t) = Me st (8)\ncan be used for de te rmin ing the lengths of the s egmen t s A/i . The symbols Pis t , Mis t and P id , Mi d denote the\ns ta t ic and dynamic components , r e spec t ive ly , of the longitudinal s t r e s s e s and to rques in the c r o s s sec t ions of the rope b r anches .\nThe f i r s t expe r imen t s in the field w e r e c a r r i e d out for an o r d i n a r y f i l ament by the au thors of monograph [51.\nThe boundary condit ions for the funct ions U i (x, t) and | t) at the rope ends x = L i a r e the equat ions of mot ion of the end loads Qi. Thus, when the Qi move along r ig id guide t r a c k s we can wr i te\nQlg [rowe-Ji~(Le' t) -t-o-- (t)] + A dU~,..~ + cOOi!L\" = Q\"\nO~(L~, t) = 0. (9)\nIf the end loads move along elastic guide tracks with the pliability coefficients ~i we have\nQt [ O*Ui (Le, t) ] A OUl (Ll, t) 00~ (L,, t) -e- [ - - ~ i ' q: v (t) J + Ox + c Ox = q\"\nO~'~i (Le, t) OOi (L~, t) I~ - - -dP- - + C Ox +\n+ B dO~ (L. t) a,Oe (Li, t) = O, (10) Ox\nw h e r e ! i a r e the m o m e n t s of ine r t i a of Qi-\nThe boundary condit ions in the c r o s s sect ion x = = ~i = li - A/i of the rope b ranches a re\ni au,(i;, 0 ~et, Ui (7.i, t) = O oi + Ox 0 t ~x . + C 00~ (l. t) , , . , . 0i(li, 1) = 0o, j-----~-~-~qut. (Ii)\nI)\nThe in tegra t ion cons tan ts U0i and | can be d e t e r - mined f r o m the init ial condit ions.\nThe lengths /i(t) of the rope b ranches in the undefo rmed s ta te a r e v e r y c lose ly approximated by the exp r e s s i o n s [5]\n!\nh(t) =/o~ + S vlt)dt, t~(t) =1o2-- S v(t) dt. (12) 0 0\nThe in tegra t ion cons tants U0i and | can be de t e r - mined f r o m the init ial condit ions. We have a l ready noted that 101 + 102 = 7rR.\nThe boundary condit ions in the rope c r o s s sec t ions x = li (t) can be wri t ten as [3]\nt r ~2 A,. o'U, q,, 0 o_U, (le, t] i,dt + j ~., OxOt tit. Ue (l . t) = Uo, + o2 ox o\nt ~ O\")i(l. t)\" 9 CA\" O~Oi(l\"t)\nOt (li. l) = 0o, + ~ lidt + _] ~ti ---dxO-t-- dr. (13) I)\nLet us de te rmine t h e re la t ionsh ips between the funct ions Vl(h -- Al 1, 0 ; 01(11 --All, t) and U ~(lt--Al v t);\nO~(/,2--Al~, t) under the condit ion that s l ippage and s l iding of the rope over the pul ley o c c u r s only over the s e g - men t s A/i, while the ve loc i t i es of the r ema in ing rope e l ement s in contac t with the pul ley a r e equal to the v e - loc i t ies of the points of the pul ley r i m .", + "Let the c r o s s sec t ion x = l 1 - Al 1 of the ascending b ranch move to the posi t ion x = 12 - Al~ of the descending b ranch in the t ime At as the r e su l t of the pu l ley ' s ro ta t ion . Since the re is no s l ippage or s l iding of the rope over the pul ley dur ing this t ime, we can wr i te out the following se l f - ev iden t r e la t ionsh ips between the s t r a ins in the ascending and descending rope b ranches :\nOU~ (l.. -- Al~, t) OUt (lx \" All, t -- At) Ox ax '\n002 (/~ - - AI~, t) = - - 00~ (l~ - - Ala, t - - At) ( 1 4 ) Ox Ox\nThe ini t ial condi t ions for the functions Ui(x, t) and | t) a re of the f o r m\nOU~ (x, O) F~ (x), O~ (x; O) = h (x), O-Y-- =\nOO~ (x, O) O/(x, O) = q~ (x) , ~ = r (x), (15)\nwhe re f i , ~i, Fi, and ~i a re known functions of x. Fo r example, i f at the beginning of opera t ion of the hois t the rope is s ta t i ca l ly s t r e t ched by i ts own weight and by the weight of the end loads Qi, then initial condit ions (15) become\na) in the c a s e whe re Qi move along r ig id t r acks ,\nOi (x, 0) =\ng :t: v k~q (x -- ll) (L,~ -- x) - - - - t + g 2~\n| (x, O) -~ - - g ~ ~ kq (x ~ l~) (Lz - - x) g 2t~ '\nOOi (x, O) OUz (x, O) _ O, = O; (16) 9 Ot at\nb) in the c a s e where Qi move along pl iable t r acks ,\ng &\ng:l:Vkg ~ ( x _ _ l , ) [ _ ~ L + q(2Li--x--l,)2 ] - -\ng ::h v kqai (Lt -- I~) 2 (x - - li), g 2~A~\nOUi (x, O) O0~ (x, O) - - - 3 F - - = o, at o. (17)\nHere\nk = C 1 A\nB C ~\" = ~AB - - C2 ' v = A B - - C ~ '\nA l = ~ -- a t (Li -- I0. (18)\nConsidering expressions (3), (6), (8), (9), (!2), (13)-(18), we see that the dynamics of a nonequilibrium\nhoist with a f r ic t ion pulley mus t be analyzed in two s tages :\n1) finding the solution of sy s t ems (3) and (6) for i = 1 which sa t i s f ies conditions (8)-(10) , (12), (13), (15)-(17) for i = 2;\n2) using the above solution for the ascending branch and condition (14) to determine boundary eonditions (13) for the descending branch of the rope; then integrating s y s t e m s (3) and (6) for i = 2 with al lowance for condit ions (8)-(10), (12), (13), (15)-(17) for i = 2.\nInvest igat ion of the above p rob lems of dynamics of a lifting rope of var iab le length by means of d i f fe rential equations involves cons iderab le ma themat ica l diff icul t ies and is not possible at p resen t . In our case it is expedient to conver t f rom t h e differential equations and boundary conditions of the problem to the c o r r e - sponding in tegro-d i f fe ren t ia l equations of motion of a mine lifting rope of va r iab le length.\nLet us r e p r e s e n t Eqs. (3) with al lowance for conditions (9), (i0), (13) in the following i n t eg ro -d i f f e r en - t im f o r m [4, 6]:\nLi\nU i (x, ' ) ~ - - I\" [(,11 (x, s, ' , ) ~ Q ~ ' ) ( s ) d 5 - o ti\nLt ~ .. , , O~Ot (s, t)\n- - t%~ (x, s, ~i) ~ ~ e ) ( s ) d s + ~i (x, t) + Uist\nl i Li OWl (% t) .~ . . ~i (X, t) = - - K~I (X, $, li) ~ O~ ; (Sl d s - -\nl i L t\n, ( 1 9 1\nli\nThe ke rne l s of Eqs. (19) s y m m e t r i c with r e s p e c t to\nx and s, the weight functions p~O(s),- (i, j = 1, 2), and\nthe functions ~,.(x, t), ~l~(x, tt are as follows: a) for mot ion of the end loads Qi along r ig id guide\nt r acks ,\n(S -- [i)[~ (L, - - li) + [,~A (Li - - x)l 5 ~ x , p.A (L, - - l i )\nK u (x, s, 13 = x - - 13 Ix (L, - - li) \"~- k 2A ( L i ~ s)[ s ?~ x;\n~A (Li - - l i )\nt k (Li - - x ) (s - - li) tt (L,. - - 1~) s ~< x,\nK u (x, s, 13 = K21 (s, x, 13 = ] / , (L, - - s) (x - - li)\nI g (L~ - - 13 s >. x;\n(Li ~ x) (s - - li} S -% X,\nK22 (x, s, li) = (Li - - ,Q (X - - li) S > X; (2 0:)\n~t ( L , - - t3" + ] + }, + { + "image_filename": "designv11_69_0002979_978-3-540-70534-5_7-Figure2.1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002979_978-3-540-70534-5_7-Figure2.1-1.png", + "caption": "Figure 2.1: Working like clockwork", + "texts": [ + " Hardware can be described on several different levels, from low-level transistor-level to high-level hardware description languages (HDLs). The socalled register-transfer level is somewhat in-between, describing CPU components and their interaction on a relatively high level. We will use this level in this chapter to introduce gradually more complex components, which we will then use to construct a complete CPU. With the simulation system Retro [Chansavat Br\u00e4unl 1999], [Br\u00e4unl 2000], we will be able to actually program, run, and test our CPUs. One of the best analogies for a CPU, I believe, is a mechanical clockwork (Figure 2.1). A large number of components interact with each other, following the rhythm of one central oscillator, where each part has to move exactly at the right time. Central Processing Unit 18 2 2.1 Logic Gates On the lowest level of digital logic, we have logic gates AND, OR, and NOT (Figure 2.2). The functionality of each of these three basic gates can be fully described by a truth table (Table 2.1), which defines the logic output value for every possible combination of logic input values. Each logic component has a certain delay time (time it takes from a change of input until the corrected output is being produced), which limits its maximum operating frequency" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003552_amr.97-101.3761-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003552_amr.97-101.3761-Figure1-1.png", + "caption": "Fig. 1 Coordinate systems used for hobbing face gears Fig. 2 Cutting face gears with a CNC hobbing machine", + "texts": [ + " But this kind of worm is limited in application because the singularities must be avoided. The authors proved in [4] that it is feasible to cut face-gears with an internal gear spherical hob [5], and the surface error can be ignored. In this paper, the scheme of cutting face gears using a CNC hobbing machine is described in detail. The determination method for the machine parameters is given. At last, the experiment is done to test the method. Litvin [3] described the process of cutting face-gear with a grinding wheel. A shaper is introduced here. Fig. 1 describes the coordinate systems used for hobbing face gears. Mobile systems Sw, Ss and S2 are tied with the hob, shaper and face-gear rigidly, respectively. Sa and Sm are immovable systems. \u03b3ws is the shaft angle between hob and shaper. \u03b3ws=90\u00b0+\u03bb0+\u03b2 (1) Here, \u03c6w and \u03c62 are the rotation angle of hob and face-gear, respectively, and \u03c62=\u03c6w w/ 2\uff0c 2 is number of teeth of face-gear. w is the number of threads of the worm, usually w=1. \u03b3m is shaft angle between face-gear and pinion; L0 is an auxiliary parameter, and it can be taken equal to the inner radius of face-gear. The worm and the face-gear surfaces \u03a3w and \u03a32 are in point contact at every instant, so the generation of the surface \u03a32 by the hob requires application of two parameter enveloping process. The hob moves in the direction of the axis of the shaper by \u2206lw and, simultaneously, the shaper should rotate by angle \u2206\u03c6s (see Fig. 1). The magnitudes of \u2206lw and \u2206\u03c6s are components of screw motion of the hob about the axis of the shaper, and they satisfy equation \u2206\u03c6s=\u2206lw/ps (2) ps is the screw parameter of the shaper, and 0.5 / sins n sp m \u03b2= . The face gear should rotate \u2206\u03c62 All rights reserved. No part of contents of this paper may be reproduced or transmitted in any form or by any means without the written permission of Trans Tech Publications, www.ttp.net. (ID: 72.19.71.81, Univ of Massachusetts Library, Amherst, USA-06/04/15,05:15:01) 2 2 sin 0", + "5w n l m \u03b2\u03d5\u2206 = \u2206 \u22c5 (3) Surface \u03a32 of face-gear generated by the hob is determined as the envelope to the two-parameter family of surfaces of hob thread surface as follows: 2 0 0 w w 2w w w w 0 0( , , , ) ( , ) ( , )l l\u03b8 \u03c8 \u03d5 \u03d5 \u03b8 \u03c8\u2206 = \u2206r M r (4) w(w 2, ) (1) w w w 2 0 0 w w( , , , ) 0f l\u03d5 \u03b8 \u03c8 \u03d5\u22c5 = \u2206 =n v (5) w(w 2, ) (2) w w w 2 0 0 w w( , , , ) 0l f l\u03b8 \u03c8 \u03d5\u2206\u22c5 = \u2206 =n v (6) The parameters in Eq. 4 to Eq. 6 can be found in Ref [3]. Base on aforementioned content, the process of cutting face-gear is more like hobbing involute helical gears [6] with normal hobs, but the sphericity hob\u2019s motion is not parallel to the axis of cylindrical gear. Scheme for applying C C hobbing machine. To realize cutting face gears with a hob, the motions described in Fig. 1 are needed. The ordinary hobbing machines can not satisfy this requirement, because the angle between the hob\u2019s axis and the axis of pinon Z1 can not be adjusted to be \u03b3ws. A scheme of cutting face gears using a CNC hobbing machine is proposed, as shown in Fig. 2. The plot plane is the plane of worktable on hobbing machine. X is radial feed direction and y is tangential feed direction. First, rotate rest of hob to make the axis of hob Zw in horizontal plane. It is suposed that the hob and the imaginary pinion are all right helical" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002098_piee.1967.0321-Figure3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002098_piee.1967.0321-Figure3-1.png", + "caption": "Fig. 3 Relative positions of reference phases of harmonic windings Windings a, b, x and y are reference phases of windings simulating fields of pa, pi,, px and pu pole pairs, respectively", + "texts": [ + "2, the positions of the corresponding phase axes of these hypothetical rotor windings may be conveniently regarded as being coincident. 6 Load angle of a squirrel-cage machine When considering any synchronous effect in a machine, it is convenient and usual to consider one of the variable parameters as being a load angle. In relation to a doubly fed cascade system, this parameter has already been discussed4, and this concept may also be extended to the multiple reactions occurring in a squirrel-cage machine. PROC. IEE, Vol. 114, No. II, NOVEMBER 1967 Fig. 3 shows the relative spatial positions of the reference phases of three harmonics pb, px and py with respect to the reference phase of the fundamental pa. It is convenient to consider that all of these component m.m.f.s rotate in the same direction, although in practice this may not be true. For the reasons previously discussed, the rotor phase windings are shown as coincident. The electrical load angle for the fundamental pair, pa and pb, say, is Kb = (Pa + Pb)\u00a7m + Pb*b (0 Similarly the load angle for the general harmonic pair px and Py is &xy = (Px + Py)&m + Px*x + Py^y \u2022 \u2022 ' \u2022 \u2022 ( 2 ) Resultant torque produced by relative displacement of two components a Angular displacement 3 = 0 b Angular displacement 0 = jt " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002819_6.2007-3813-Figure7-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002819_6.2007-3813-Figure7-1.png", + "caption": "Figure 7. Boundary Layer Elimination Setup.", + "texts": [ + " The second factor was determining at what pressure the blower plenum should be set to provide the optimum mass flow rate. To measure the boundary layer a United Sensor BR-.065-12-F-11-.250 boundary layer probe was mounted to a 2 axis traverse system and used to traverse the inside of the test section in front of the bullet while recording total pressures. Specifically, the probe traversed the test section on both sides of the bullet centerline and from the top plate to 12.7 mm below the plate while the tunnel was running at Mt = 0.72. The probe, shown in Figure 7, was inserted into the test section from slots cut in the bottom plexiglass plate. The total pressures recorded from the boundary layer probe were used to calculate Mach numbers every 0.635 mm for Mt = 0.37 and 0.56 out to 12.7 mm and every 1.27 mm for Mt = 0.72 out to 25.4 mm. The boundary layer height was found to vary with Mt and ranged from approximately 5 mm at Mt = 0.37 to 8 mm at Mt = 0.72. The slot height and mass flow were varied until the Mach numbers at points inside the boundary layer were similar to the unobstructed flow points" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000802_ijmee.32.4.2-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000802_ijmee.32.4.2-Figure2-1.png", + "caption": "Fig. 2 Solid model of a vehicle.", + "texts": [ + " Wheeled bases are provided to the students; the students must design their car bodies such that they can dock with the wheeled bases. Two to three instructors are present to assist the students in the design process. The resulting designs are exported as *.stl files, and transmitted to the RPC. Ideally, this design phase would be placed later in the week, after more theoretical material had been introduced. However, the time limitations and machine availability in the RPC dictate that part production must start as soon as possible. An example of a car design produced by a student participant can be seen in Fig. 2. Students are given an introduction to solid freeform fabrication technology, including an opportunity to tour the RPC and view their car bodies being made. In an effort to stimulate further interest in engineering as applied to automobile design, a half-day program on internal combustion engines is presented. This includes the generation of a torque\u2013speed curve for an engine using dynamometer International Journal of Mechanical Engineering Education 32/4 at UCSF LIBRARY & CKM on March 13, 2015ijj" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000396_2001-01-1445-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000396_2001-01-1445-Figure1-1.png", + "caption": "Fig. 1 Target front suspension system", + "texts": [ + ", in a transmission path of source vibration. In practice, however, this stiffness is rarely changed because of the difficulty in achieving compatibleness among riding comfort, controllability and stability. In this paper, a rubber mount optimization method for road noise reduction corresponding to an improvement in riding comfort (harshness) is demonstrated. The rubber mounts for optimization, which are installed in the front and rear sides of the lower arm on the front suspension system in the target vehicle, are shown in Fig. 1. To change the mount stiffness effectively, it is important to understand the following items in detail. 1. Contribution relationship between suspension system\u2019s vibration and road noise 2. Suspension resonance existing in the problem frequency band 3. Mount location sensitive to vibrations that cause road noise and riding discomfort First, to clarify item 1, Road Noise Contribution Analysis (RNCA) is performed by using noise and vibration data measured in an in-operation test on a road noise evaluation course" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002979_978-3-540-70534-5_7-Figure8.14-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002979_978-3-540-70534-5_7-Figure8.14-1.png", + "caption": "Figure 8.14: Trajectory calculation for Ackermann steering", + "texts": [ + "13) that the vehicle\u2019s overall forward and downward motion (resulting in its rotation) is given by: forward = s \u00b7 cos \u03b1 down = s \u00b7 sin \u03b1 v \u03c9 2\u03c0r 1 2 -- 1 2 -- 1 d -- 1 d -- \u03b8\u00b7 L \u03b8\u00b7 R = s\u00b7 d\u03d5 dt\u2044 \u03d5\u00b7 \u03b8\u00b7 L R, \u03b8\u00b7 L \u03b8\u00b7 R 1 2\u03c0r -------- 1 d 2 -- 1 d 2 -- v \u03c9 = Driving Robots 144 8 If e denotes the distance between front and back wheels, then the overall vehicle rotation angle is \u03d5 = down / e since the front wheels follow the arc of a circle when turning. The calculation for the traveled distance and angle of a vehicle with Ackermann drive vehicle is shown in Figure 8.14, with: \u03b1 steering angle, e distance between front and back wheels, sfront distance driven, measured at front wheels, driving wheel speed in revolutions per second, s total driven distance along arc, \u03d5 total vehicle rotation angle The trigonometric relationship between the vehicle\u2019s steering angle and overall movement is: s = sfront \u03d5 = sfront \u00b7 sin \u03b1 / e Expressing this relationship as velocities, we get: vforward = vmotor = \u03c9 = vmotor \u00b7 sin \u03b1 / e Therefore, the kinematics formula becomes relatively simple: Note that this formula changes if the vehicle is rear-wheel driven and the wheel velocity is measured there" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001701_amc.2006.1631675-Figure4-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001701_amc.2006.1631675-Figure4-1.png", + "caption": "Figure 4. A motion profile ( ) with NCp and NCa free collocation points.", + "texts": [ + " Here, represents a normalized time scale: = t / T. Hereafter, the prime symbol will indicate a derivation with respect to Hence, the problem is transformed to a parametric optimization problem. One of the parameters is the unknown traveling time T. The other parameters are four sets, SPp, SPa, SCp and SCa of free discretisation nodes. The sets SPp and SPa are composed, respectively, of NPp and NPa control points in the robot workspace (Figure 3) while SCp and SCa consists of NCp and NCa collocation points in the ( , ) plane (Figure 4). With SPp and SPa, we can define a path q( ) using parametric functions, such as B-spline, that takes into account constraints (3a) and (3b). Note that ( p) is deduced directly from x( p) and y( p) using the nonholonomic constraint : 0)()()()( pppp CosySinx (6). With SCp and SCa, we can define a motion profile ( ) using, for example, a clamped cubic spline interpolation that takes into account constraints (3b). The key point is that the traveling time T becomes a dependent parameter. Indeed, for any valid trajectory profile q( ) q( ( )), the cost function J can be written explicitly in terms of the single variable T" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000391_2000-01-0920-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000391_2000-01-0920-Figure2-1.png", + "caption": "Figure 2. Abrasive wear", + "texts": [ + " The major problem in lubricated conditions is to decide how much of the load is born by a fluid film and how much is carried by boundary lubricated contact. Because of this fact, the wear coefficient varies in a wide range as function of the lubrication properties. If a fraction of the load is carried by the oil film, there will be less wear than in the case where all of the load is carried by the contacting surfaces. Assuming that the asperities have the shape of a cone, abrasive wear can be calculated as follows: referring to Figure 2 the load on the sliding asperity, respectively the worn volume, which occurs during one sliding act, can be calculated as (2) respectively (3) Integrating over the surface results in V s --- k L H ---\u22c5= \u03b4L 1 2 -- H \u03c0 z2 \u03a6tan 2\u22c5 \u22c5 \u22c5 \u22c5= \u03b4V z2 \u03a6tan\u22c5( )\u03b4s= 3 The development of this Piston Ring Wear Model for piston rings has been divided into seven parts: \u2022 Simplification of the piston ring kinematics for simple calculations in reference to the motion of the piston ring \u2022 Segmentation of the piston ring in order to obtain a three dimensional model for piston ring wear \u2022 Calculation of the force distribution in the axial direction of the piston ring \u2022 Development of a surface model \u2022 Determination of the amount of wear \u2022 Implementation of ring twist and piston tilt \u2022 Implementation of the single models and interface to the piston ring simulation program The following assumptions are valid for the kinematic model and lead to a more simple simulation: \u2022 The oil film thickness is very thin compared to the bore size \u2022 The lubricant is an incompressible fluid \u2022 The surfaces of the ring face and the liner are perfectly smooth and rigid \u2022 The ring may twist and the piston may tilt during an engine operation \u2022 The ring follows the motion of the piston and additionally it twists As long as there is enough information about the contacting circumstances between the piston ring and liner, the problem of a sliding piston ring in a cylinder liner can be simplified to a slider plane model, which is easier to model and to calculate" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001192_s00419-006-0003-2-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001192_s00419-006-0003-2-Figure2-1.png", + "caption": "Fig. 2 Geometry of rotating ring with boundary conditions", + "texts": [ + " The assumption (8) is also a consequence of the small cross-section of a pin and means clearly that the corresponding heat conduction problem can be considered as the one-dimensional. The assumption (9) about the average temperatures equality in the contact region is the standard used one and serves for the definition of the heat distribution coefficient. We shall note, that in the scientific literature also the condition of the maximal temperatures equality in the field of contact, is used. The geometry and boundary conditions for the rotating ring are shown in Fig. 2. It is assumed that the coordinate system (r, \u03b8 ) is fixed to the heat source and the ring rotates with constant speed \u03c9 with regard to this coordinate system. Taking the assumptions (1) \u2013 (7) into account, the governing quasi-stationary heat conductivity problem in dimensionless form is [11]: \u22022T \u2217 \u2202\u03c12 + 1 \u03c1 \u2202T \u2217 \u2202\u03c1 \u2212 \u03c3 T \u2217 = Pe \u2202T \u2217 \u2202\u03b8 , \u03c10 < \u03c1 < 1, 0 < \u03b8 < 2\u03c0, (1) \u2202T \u2217 \u2202\u03c1 \u2223 \u2223 \u2223 \u2223 \u03c1=1 = { 1 \u2212Bi T \u2217 0 \u2264 \u03b8 \u2264 2\u03b80 2\u03b80 < \u03b8 \u2264 2\u03c0 , (2) \u2202T \u2217 \u2202\u03c1 \u2223 \u2223 \u2223 \u2223 \u03c1=\u03c10 = Bi0 T \u2217, (3) using the finite Fourier transform with respect to \u03b8 [12] T \u2217 (\u03c1, n) = \u03b5 2\u03c0 2\u03c0\u222b 0 T \u2217(\u03c1, \u03b8) exp(\u2212in\u03b8)d\u03b8, n = 0, 1, 2, " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001460_s00158-004-0489-6-Figure3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001460_s00158-004-0489-6-Figure3-1.png", + "caption": "Fig. 3 Four-bar linkage demonstration problem", + "texts": [ + " Thus, if the calculated RLD \u2192 0 as \u03b4\u03c9 \u2192 0 we still cannot prove that the solution does not exist. What we can approximate is how much time, or effort, is expected to solve for the true desired MSE solution. By tracking both the trends in the RLDs and the candidate design criteria as \u03b4\u03c9 \u2192 0, we can gain a feeling for how close one is to the performance space boundary. 3.1 General problem description The sample synthesis problem is that of a four-bar linkage with eight linear springs attached to the links of the mechanism (see Fig. 3). The linkage is assumed to be a Grashof mechanism, or Class I linkage, and therefore, the linkage construction is restricted such that the input crank (link l I) can fully rotate 360\u25e6 (Erdman and Sandor 2001). The linear translational springs are the only energy storage elements. Although the springs are linear, the linkage provides the nonlinear motion that enables multiple stable equilibria. Kinematically, this is a one degree of freedom system described by the input crank angle \u03b8. The goal of the synthesis problem is to determine a combination of the following design variables: the spring constants (ki ) and free lengths (loi) of the eight springs, the locations of anchors for the springs fixed to the ground (x1, x2, y1, y2), and link lengths (l I , l II , l III , l IV )" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000210_iros.1996.571043-Figure9-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000210_iros.1996.571043-Figure9-1.png", + "caption": "Figure 9: Mobile manipulator", + "texts": [ + " However, when we set the displacement of the initial position of the object larger or the coefficient of friction smaller , the tracking performance degrades because of the velocity limitation. 6 Pushing Operation with a Mobile Manipulator An approach to realize the pushing operation with a mobile manipulator, that is, a mobile robot with a manipulator, is discussed in this section. We assume the mobile manipulator to be a mobile robot with two independent driving wheels on which a threedimensional manipulator with n degree-of-freedom is mounted (see Figure 9). 6.1 Control of mobile manipulator We adopt a proportional and differential feedback control rule to the arm tip of the manipulator. As for the mobile robot part, we use a feedback control method proposed in [14]. In this method the mobile robot is hauled by the reference point fixed in front of the mobile robot. Using this method, although the orientation of the mobile robot cannot be controlled, the consideration of the non-holonomic constraint can be removed. A unified control method for the mobile manipulator combining these two control rules is described below. As in Figure 9, we denote the position of reference point expressed in the universal frame Cu as \"p,, the position of the arm tip as ' p h , the angular displacement vector of the right and left wheels as qp = [y,, qlIT, and the joint vector of manipulator as q, = [yl, ,y,IT. Since the pushing operation is assumed to be a task on a horizontal plain which is two dimensional space, \"p, is a two-dimensional vector without the axis component, and the component z of \" p h is constant. Then, letting \"r = [\"p:, and q = [q: , q2lT, the relation among the velocities of the reference point and the arm tip, and the joint velocity is given by '+ = J q (15) where and J h m is the Jacobian matrix of \" p h with respect to q,, J,, and J h p are the Jacobian matrixes of \"p, and ' p h with respect to qp , respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003398_ijmr.2009.026577-Figure5-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003398_ijmr.2009.026577-Figure5-1.png", + "caption": "Figure 5 Sections for thermal distributions", + "texts": [ + " The nodal equation for an interface surface or the corner separating two different materials are complex and have been determined by Holman (2002); these equations together with the normal nodal equations are summarised in Appendix A, as they appear in the Microsoft Excel worksheet. Note that the nodal equation for a composite boundary contains a parameter exterior interior/ ,P K K= which compensates for a change in the thermal conductivity. The computations are done twice to determine the thermal distribution in two different sections of the tool insert, Section A-A, approximately normal to the chip flow direction, and Section C-C, approximately parallel to the chip flow as shown in Figure 5. Computations were conducted for cutting speeds of Vw = 50, 100, 150, 200 m/min, depth of cut d = 0.25 mm and a feed of f = 0.2 mm/rev. Computations are done at a constant feed f = 0.2 mm/rev and speed Vw = 100 m/min for different depths of cut of d = 0.25, 0.5, 1.5 mm. The increment size in the nodal network is taken as \u2206x = \u2206y = 0.1 mm = 0.0001 mm resulting in a nodal network of 52500 nodes for Section A-A and 62500 nodes for Section B-B. The geometric configuration and boundary conditions for the problem related to Section A-A and Section B-B are illustrated in Figure 6" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000498_306-Figure3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000498_306-Figure3-1.png", + "caption": "Figure 3. RG-function and flow for increasing manifold size L for the dimensionless renormalized coupling g: (a) in the case \u03b5 > 0, (b) in the case \u03b5 < 0, (c) in the case \u03b5 = 0.", + "texts": [ + " Apart from the trivial solution, g = 0, the flow equation given by (2.12) and (2.13) has a nontrivial fixed point at the zero of the \u03b2-function g\u2217 = 2\u03b5 + O(\u03b52). (2.14) We shall show below that the scaling behaviour is described by the slope of the RG-function at the fixed point, which is universal as a consequence of renormalizabilty. The long-distance behaviour is then governed by the \u03b4-interaction as considered in our model (2.6), which is the most relevant operator at large scales. Let us now discuss possible physical situations (see figure 3): (a) \u03b5 > 0. The RG-flow has an infrared stable fixed point at g\u2217 > 0 and an IR-unstable fixed point at g = 0. The latter describes an unbinding transition whose critical properties are given by the noninteracting system, while the nontrivial IR stable fixed point determines the long-distance properties of the delocalized state, the long-range repulsive force exerted by the fluctuating manifold on the origin\u2014which we recall may be a point, a line or a plane. (b) \u03b5 < 0. Now, the long-distance behaviour is Gaussian, while the unbinding transition occurs at some finite value of the attractive potential, g\u2217 < 0, which corresponds to an infrared unstable fixed point of the \u03b2-function", + " This is the marginal situation, where the transition takes place at g\u2217 = 0; we expect logarithmic corrections to scaling. Note that in the presence of an impenetrable wall constraining the configurational space strictly to half of the embedding space, the above considerations should still apply, when shifting the interaction strength appropriately. We shall discuss that in section 2.3. Since we are mainly interested in the long-distance properties of membranes for which always \u03b5 > 0 (this is (a) in figure 3), let us try to calculate the repulsive force exerted by the membrane on the origin in the case where this point is strictly forbidden. We shall derive a universal expression for this force [17]. We need the (not normalized) membrane density at position r Z( r, g0) := 1 VM \u222b D[ r] \u222b M dDx \u03b4d( r(x) \u2212 r)e\u2212H. (2.15) Since the \u03b4-interaction also appears in H, we can relate the density at the origin to the derivative of the partition function with respect to g0: Z( 0, g0) = \u2212 1 VM \u2202 \u2202g0 \u2223\u2223\u2223\u2223 L Z(g0) = 1 N \u2202(gL\u2212\u03b5) \u2202g0 \u2223\u2223\u2223\u2223 L (2", + "24) Comparing the L-dependence of Z\u221e(r/L\u03bd) and Z( 0, g0), we obtain the exponent identity \u03b8 = \u03b5 + \u03c9(g\u2217) \u03bd . (2.25) Finally, from (2.24) we derive the repulsive force between the origin and the manifold f ( r) = \u2207 r ln Z\u221e(| r|/L\u03bd) = \u03b8 r r2 . (2.26) Note that to derive this result, kBT has been set to 1. Reestablishing the temperature dependence, we find f ( r) = kBT \u03b8 r r2 . (2.27) Also note that this argument gives \u03b8 = 0 at the Gaussian fixed point, which is necessary since for g0 = 0 no force is exerted on the membrane. Let us discuss the physical situation at the UV-stable fixed point in figure 3. The fixed point corresponds to a delocalization transition of the manifold, which is at vanishing coupling g\u2217 = 0 for \u03b5 > 0 and at some finite attractive coupling g\u2217 < 0 for \u03b5 < 0. In the localized phase g < g\u2217, correlation functions such as \u3008[ r(x) \u2212 r(y)]2\u3009 and the associated correlation length \u03be\u2016 (in the D-dimensional internal space) should be finite, as well as the radius of gyration \u03be\u22a5. Approaching the transition point these quantities diverge as [14] \u03be\u2016 \u223c (g\u2217 \u2212 g)\u2212\u03bd\u2016 \u03be\u22a5 \u223c (g\u2217 \u2212 g)\u2212\u03bd\u22a5 . (2" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001498_s10808-006-0105-1-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001498_s10808-006-0105-1-Figure1-1.png", + "caption": "Fig. 1. Layout of the welding bath (section y = 0): 1) solid phase; 2) two-phase zone; 3) liquid phase; 4) vapor\u2013gas channel; 5) laser-beam axis.", + "texts": [ + " The effects of optical breakdown and vapor ionization are ignored. In the domain considered, we introduce a Cartesian coordinate system in which the laser beam incident onto a dense joint of the welded plates is motionless, and the plates move with the welding velocity v. The downward z direction is coaxial with the beam, the x axis is directed along the joint and follows the direction of motion of the plates, and the y axis is perpendicular to the joint. The origin lies at the beam axis on the upper surfaces of the plates (Fig. 1). To protect the metal from oxidation, the welded plates are subjected to a neutral gas flow, which partly entrains metal vapors. 2. Governing Equations and Relations. In the chosen coordinate system, the quasi-steady equations of heat transfer can be written in the form ceiv \u2202T \u2202x = \u03bbi (\u22022T \u2202x2 + \u22022T \u2202y2 + \u22022T \u2202z2 ) ; (1) cei = \u23a7\u23aa\u23aa\u23a8 \u23aa\u23aa\u23aa\u23a9 c1\u03c11, T < Te, c2\u03c12 ( 1 + \u03ba c2 \u2202fl \u2202T ) , Te T Tl0, c3\u03c13, Tl0 < T, (2) where ci, \u03bbi, and \u03c1i are the specific heat, thermal conductivity, and density of the ith phase (the subscripts i = 1, 2, and 3 refer to parameters of the solid, two-phase, and liquid states of the metal, respectively), Tl0 and Te are the temperatures of the beginning and end of metal solidification, fl is the section (fraction) of the liquid phase in the two-phase zone, and \u03ba is the latent heat of melting", + " (28) Here x\u2212 c and x+ c are the coordinates of the points of intersection of the channel boundaries with the surface z = 0; x\u2212 e and x+ e are the coordinates of the points of intersection of the melting-region boundaries with the surface z = 0 and of the solidification-region boundary with the weld bed, respectively; the superscripts minus and plus refer to the left and right boundaries with respect to the z axis, respectively. Equations (16), (18), and (26) with allowance for relations (2), (6), (9)\u2013(15), (17), (19)\u2013(25), (27), and (28) were solved numerically by the finite-difference method. 5. Brief Description of the Numerical Algorithm. In the plane of the section of the welding region considered (see Fig. 1), we choose a computational domain G, whose form at the first stage of the computations (without the vapor channel and the weld bed from the solidified metal) is a rectangle with the sides z = 0, z = h, x = \u2212l1, and x = l2, where l1 and l2 are the distances from the left and right boundaries of the rectangle to the laser-beam axis, respectively. The choice of l1 and l2 depends on the laser power and welding velocity. To solve Eq. (16) numerically on a rectangular difference grid Gh with the nodes {kh1, mh2} (k = 0, 1, " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000354_robot.2001.933231-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000354_robot.2001.933231-Figure1-1.png", + "caption": "Figure 1: A three-link redundant planar manipulator for simulation.", + "texts": [ + " For a two-link planar manipulator, the orientation of the end-effector is defined by the orientation of the second link, and can not be controlled; whereas, we can control the orientation of the end-effector of a three-link planar manipulator through the extra degree of freedom of the redundant joint. Thus, this three-link manipulator is called a \u201credundant manipulator\u201d if we are concerned with only the linear {x, y> coordinates with three joint parameters, {&,&, Os}. The geometry of the redundant manipulator is shown in Figure 1, where three serial links are configured with the reference point at the end-effector at P(x, y), and the robot base frame is at 0 - X Y . We denote the position of the end-effector as x = [x, y] in the Cartesian space and 8 = [el, 02, , & I T are the joint parameters. From Figure 1, the Cartesian position of the end-effector can be written as T x = Llcl + LZC12 + L3c123 (25) y = Llsl f L2s12 $- L3s123 (26) where c1 = cos&, s1 = sin&, c12 = cos (ez + &), etc. Thus, the Jacobian matrix becomes 4.1 Cartesian-based stiffness control The initial orientation of the third link is maintained in horizontal orientation, as shown in Figure 1. After that; the orientation of the third link is specified to align with the radial direction of the trajectory, as shown in the shaded third link in Figure 1, so as to align the end-effector with the orientation of the grasp object for grasping tasks and stable prehension. This gives rise to the advantage of the three-link redundant manipulator over the two-link manipulator. Since we are considering the Cartesian-based stzffness control scheme, the stiffness matrix, K,, and the bias force, f , in Cartesian space should be specified first. In this simulation, the prescribed conservative Cartesian stiffness matrix and the bias force are The matrix K, satisfies the two conservative criteria in equations ( 3 ) and (4)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000913_icnsc.2004.1297033-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000913_icnsc.2004.1297033-Figure1-1.png", + "caption": "Figure 1: Underway Replenishment model (PI [zI,yllT, PZ' [z2,Y2IT, Pd ' [ z d , y d I T )", + "texts": [ + "r) - V2, (11) + G , d e t ) k ~ ) + K'RT(dJ2)d, (1'4 r'2 = -MZ-'Dzv2 + MF1'r2 + A4y'(Gfi(el)8n where r,,f = [z,.r,y,,r.OlT E R3 is the desired vessel separation, yrer is the constant lateral separation, x,,f = 0 is the constant longitudinal separation, and Dvz [ ~ , ~ , Y ~ Z , O ] ~ E W3 is the virtual reference point vector for the tracking vessel. The disturbances Gn(et)Bfi +G,2(et)9,2 E R3 capture interaction forces and moments acting on the tracking vessel. The Underway Replenishment, two-vessel tracking model, is shown in Figure 1. A key advantage in formulating the two-vessel tracking problem by (8), (9), (ll), and (12) is that the interaction disturbances are a function of the states. Furthermore, we can use inertial coordinates for position measurements while employing the relative (vessel) coordinates for tracking. In addition, note that the system has a cascade structure and hence standard backstepping techniques (51 are directly applicable for addressing disturbance decoupling and reference tracking. In our formulation we assume that measurement noise and 'first-order wave disturbances are properly filtered" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003904_6.2009-4887-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003904_6.2009-4887-Figure1-1.png", + "caption": "Figure 1. OH-58 Transmission Schematic (left) and Disassembled (right)", + "texts": [ + " The intermediate shaft from the first stage also contains a 27-tooth sun gear. The sun gear drives three or four 35-planet planetary gears, depending upon aircraft model. The planets mesh with a stationary 99-tooth ring gear splined to the top of the transmission casing. The rotating planet gears drive the carrier, which is attached to the output shaft. The output shaft rests within the intermediate shaft. Various roller and ball bearings hold the shafts and planets in place. The total reduction ratio is 17.44:1. Figure 1 depicts the transmission of the A-model aircraft. III. Transmission Model The model uses a lumped parameter, finite element formulation of the familiar equations of motion. (1) The subcomponents of the model consist of the three shafts, the stationary ring gear, and the planet gears. In other words, these five subcomponents have their own matrix elements for satisfying Eq. 1 above41. The spiral-bevel pinion, spiral-bevel gear, sun gear, and planet carrier are assumed to be point masses acting at the appropriate nodes along their respective shaft" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001513_sice.2006.314990-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001513_sice.2006.314990-Figure1-1.png", + "caption": "Fig. 1 Model of the Furuta pendulum", + "texts": [ + " This method is achieved by reducing the difference between the angular velocity of the actual pendulum and the reference. The new control design method based on the angle is proposed. However, this method doesn't pay attention on the stability of the base link of the pendulum. Therefore, the base link is stabilized by SDRE that pays attention only to the base. The optimal control law to which the base is stabilized is obtained at every control cycle by SDRE[3][7] [8] [9]. These method are evaluated by simulation with Furuta pendulum. 2. FURUTA PENDULUM Fig. 1 shows a schematic model of the pendulum. The M(0)0 + H(O, 0) + G(O) T (1) where 0 and T are the state and the control input variables, respectively defined as 0 02 ' T T 89-950038-5-5 98560/06/$10 C 2006 ICASE 01 and 02 are the rotation angles of the motor and the pendulum (i.e. link 2 as mentioned below). Ti is the control input of the motor. The matrices M(0), H(0, 0), and G(0) are determined respectively by Eq.(7) shows the integration of the right-hand side of Eq.(4). 02 02 0rd0r= sOisn20,+ \u00b1m r2 sinOi) d0r (7) E m11 Mi2 1 iM21 M22 j [0 ]G(0)= Eg Ja + Jp sin2 02 -m21lr2 cos 02 Jpjp -m211r2 sin 02 02 + Jp0lA2 sin 202 + C0l1 -2JPsin2O O2' +CA 0 92 -m2r2g sin 02, whereJa = J1 +mirl1+m2l2,andJp = J2+m2r&2 Eq" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001949_iembs.2006.260557-Figure3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001949_iembs.2006.260557-Figure3-1.png", + "caption": "Fig. 3 The load conditions in Finite Element Model", + "texts": [ + " The following is its second-order function: N k k k N ji ji ij J d IICW 1 2 1 21 )1(1)3()3( Where ijC and kd are material parameters, their typical values can be found in [7]; The definition of W , J , 1I and 2I are the same as them in Mooney-Rivlin model. All the material properties and element types in the FE model could be found in Table.1. In the reposition procedure of orthopedic surgery, the main work is to separate the two broken bones and align them. As the Fig.2 shows, the upper cru was fixed to a bracket using a belt. And the lower cru was fixed to the reposition parallel robot through a Kirschner nail, which drilled though the calcaneus, then was fixed to the parallel robot. As Fig.3 shows that, according to the real conditions in the orthopedic surgery, six degrees of freedom of the points besides the belt in the upper crus were defined as zero. In the two sides of the FE model of Kirschner nail, two forces were applied. Their value was the half of the parallel robot used to separate the two broken bones. Their directions were the same as that of parallel robot moved. . EXPERIMENT To validate the authenticity of the FEM based biomechanical model for HIT-RAOS, a reposition experiment was conducted" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000694_b:tels.0000029042.75697.f0-Figure8-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000694_b:tels.0000029042.75697.f0-Figure8-1.png", + "caption": "Figure 8. In FMCW radar, range is calculated from the difference between transmitted and received signal frequencies.", + "texts": [ + " The result of this mixing is the frequency difference which is known as the intermediate frequency (IF) or beat frequency fb. This is a measure of the range. The range resolution depends on the sweep bandwidth f and the linearity of the sweep. For a sweep bandwidth of f = 600 MHz and sweep duration of Td = 1 ms the theoretical range resolution from equation (17) is obtained as \u03b4Rchrip = 0.25 m. Here c is the velocity of propagation, \u03b4Rchrip = c 2 f . (17) Linearity of the sweep, Lin is defined in equation (18) where S = \u03b4f/\u03b4t is the chirp slope in Hz/s as shown in figure 8. Lin = Smax \u2212 Smin Smin . (18) Nominal linearity of the radar chirp is about 0.1%. From equation (19), the resolution for a target at R = 500 m is obtained as \u03b4Rlin = 0.5 m. At shorter ranges the linearity term becomes less dominant and can often be ignored if the linearity is good \u03b4Rlin = R Lin. (19) The final system range resolution \u03b4R, can be determined from equation (20) which resolves as \u03b4R = 0.56 m: \u03b4R = \u221a \u03b4R2 chrip + \u03b4R2 lin. (20) The beat frequency for the above mentioned target at R = 500 m can be calculated from equation (21) as fb = 2 MHz: fb = Tp \u03b4f \u03b4t \u2248 Tp f Td \u2248 2R c f Td " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003744_demped.2009.5292754-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003744_demped.2009.5292754-Figure2-1.png", + "caption": "Fig. 2 - SHT-induced bar flection.", + "texts": [ + " The absolute amplitude of torsional oscillations, and so the structural stresses and fatigue, essentially depend on the typical mechanical answer of the system ring-stack-ring to external torque solicitations, and in particular on the resonance frequencies related to various resonant modes, besides the damping. SHTs are applied to the rotor magnetic stack (thought as a rigid structure), which rotates at an average speed with superimposed sixth harmonic speed components. Bars and short-rings appear as suspended masses trailed to rotate and oscillate. The bar section between ring and stack is solicited and it bends elastically, Fig. 2; crystallographic analyses of cracked bars (by using penetrating liquids) showed that usually the most stressed point is not the soldering point (since the crystalline structure does not present significant alteration and the soldering site has not defect), but immediately below the ring-bar soldering site itself, where the bending is more accentuate [5]. The soldering (or brazing) process in fact makes the molded material in correspondence of the ring-bar copper junction mechanically stronger, due to the added soldering material; the latter produces a copper-alloy with more robust mechanical properties, thus causing a larger local warp of the contiguous pure-copper regions" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002597_acc.2008.4586601-Figure13-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002597_acc.2008.4586601-Figure13-1.png", + "caption": "Fig. 13. Differential piece of snake.", + "texts": [ + " In this case the flapper angles \u03b81 and \u03b82 have opposite phases to accommodate locomotion. In the next section we will extend the flapper to a simplified snake model. We extend the previous analysis to a simplified model of a snake, Fig. 12. We will derive the friction forces exerted by the environment on the snake. First, consider a small piece with length dr of the snake from either the upper or lower bar, which is located at r(t) with respect to a inertial reference frame with basis {ex, ey}, Fig. 13. According to the differential friction model (1) the friction experienced by this differential slab is FBx (r, \u03b2) = \u2212\u00b5a(\u3008r\u0307, Bx\u3009)\u3008r\u0307, Bx\u3009Bx FBy (r, \u03b2) = \u2212\u00b5T \u3008r\u0307, By\u3009By, (11) where \u3008\u00b7, \u00b7\u3009 denotes the inner product, Bx = [ cos(\u03b2) sin(\u03b2) ]\u2032 is the axial direction and By = [ \u2212 sin(\u03b2) cos(\u03b2) ]\u2032 is the transversal direction, both with respect to the inertial reference frame. The total x and y components of the friction force are Fx(r, \u03b2) = \u3008FBx , ex\u3009 + \u3008FBy , ex\u3009, and Fy(r, \u03b2) = \u3008FBx , ey\u3009 + \u3008FBy , ey\u3009, respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001275_j.optlaseng.2005.01.004-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001275_j.optlaseng.2005.01.004-Figure1-1.png", + "caption": "Fig. 1. Near-field condition and the experimental setup. A collimated laser beam of diameter D falls through a thin cell and the interference between transmitted and scattered wavefronts is recorded at a close distance z far from the sample through magnifying optics. The near-field condition is met when the light falling at each sensor position comes from a region of the sample with dimension D 5D.", + "texts": [ + " An example of the technique for mapping the two-dimensional velocity flow around an obstacle inserted along the fluid flow is presented. In the following sections, we first introduce the experimental apparatus and a brief description of the expected effects depending on the fluid velocity, and then we present the experimental results describing both the spectral analysis and the cross-correlation analysis. The experimental apparatus has been extensively described and discussed elsewere [12,13]. A sketch of the setup used for the experiment is shown in Fig. 1. A spatially filtered, collimated and expanded laser beam is sent through a cell filled with the fluid to be studied (pure water), in which latex colloids of known size are suspended as tracking particles. The speckle intensity distribution at a given distance z from the sample is recorded by a CCD sensor through a magnifying collection optics. The intensity distribution at the observation plane is the result of interference between transmitted and scattered light. If the scattered field eS is negligible with respect to the transmitted field e0, the intensity distribution can then be written as f \u00f0r; t\u00de \u00bc i0 \u00fe df \u00f0r; t\u00de, (1) where i0 \u00bc |e0| 2 is the transmitted intensity and df \u00f0r; t\u00de \u00bc e0e S\u00f0r; t\u00de \u00fe e 0eS\u00f0r; t\u00de is called the heterodyne signal and depends on the scattered field eS(r, t) at position r and time t", + " For any given time t, the heterodyne signal df depends stochastically on the position r over the sensor plane. This speckle-like appearance is due to the sum of many independent interference patterns originating from interference between the transmitted beam and the field scattered by each particle, whose position inside the scattering cell is random. Close enough to the sample, in the so-called near-field condition, the sensor receives light from a region D 2zymax smaller than the illuminated region D (see Fig. 1). This can be due either to the limited scattering angle of the scatterer or to the limited angular aperture of the collection optics. As a consequence, if all scatterers are rigidly displaced transversally to the optical axis, provided that their mutual positions remain unchanged, the speckle field is completely preserved but displaced accordingly. This is true as long as the set of scatterers responsible for a given speckle are subjected to a constant illumination, a condition that is met only when the region D* is well inside the illuminated region D", + " 3, in which the different velocities have been represented as arrows starting from the center of the frame to the maximum value of the cross-correlation function. A key point here deals with the need to guarantee that the local displacement measured within a single frame in the speckle field image really corresponds to the local velocity in the fluid. This is not obvious since the intensity distribution in a point of the observation plane is determined by the light coming from an extended region as shown in Fig. 1. To overcome this problem we used small aperture collection optics and set the observation plane in the center of the cell. Finally, the influence of particle motion on the power spectra is analyzed. The sample was the same as that of Fig. 3, namely 3 mm particles in water, and the analysis was carried out over the same set of data. Fig. 4a shows the static 2D spectrum corresponding to Eq. (9) obtained by averaging 60 independent sample configurations. The corresponding 1D spectrum is reported in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003622_s0026-0657(09)70201-2-Figure6-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003622_s0026-0657(09)70201-2-Figure6-1.png", + "caption": "Figure 6. The freedom to change the cross-section shape enables the DMLS user to implement new solutions. In this case, the preferred design does not require a new channel layout. In this example two ejectors are bypassed in the space between without pushing the remaining wall thickness to a critical limit. The cross section area along the channel has to be kept constant, to avoid flow braking effects.", + "texts": [], + "surrounding_texts": [ + "September 2009 MPR 11metal-powder.net\nOne way to get around this, he says, is to combine DMLS with traditional technology and use it to complete manufactured tools by adding the complex exterior parts.\nLBC updated their technology 2006 when the company bought a batch of newly updated machines from EOS. So how have things improved in the last few years? \u201cSince 2004 we\u2019ve found that building accuracy and surface quality has improved and it is now possible to use a greater range of metals,\u201d says Mayer.\u201d We generally used steel 1.2709 or stainless steel for inserts but it is likely that aluminium, titanium and other steel grades will be used in future, provided that their carbon content is relatively low.\u201d\n\u201cRecently we have been able to use new materials matching properties and materials already known, such as cobalt-chrome, titanium, and maraging steel,\u201d adds Jordan. \u201cThere have also been software updates. The development of parts for the dental industry is a new innovation.\u201d\nBut how will the technology fare in an economic slowdown? \u201cOne economic benefit is that customer demands can be quickly adapted to. In the case of insecure demand forecasts, tools can be omitted and plastic parts can be made directly with plastics laser-sintering. In this way, the labour cost share of total costs is less and material can be saved,\u201d says Jordan.\n\u201cIn the future we anticipate interest from all markets that combine complex product designs with an urge to adapt to individual customer demands (such as aerospace, medical implants, special purpose equipment, tooling, etc.) We plan to continue to develop the technology for dedicated applications.\u201d\nWorldwide expansion could also be on the cards. Due to disputes with 3D Systems, EOS has only been able to distribute DMLS technology in the US for the last four or five years. The first purchaser of the technology was Morris Technologies Inc (MTI) is a rapid prototyping company. In May 2009, the company bought its ninth machine, an EOSINT M 270 system for laser sintering titanium. \u201cWe were the first US firm to install DMLS equipment,\u201d says Greg Morris, CEO/ COO of MTI, \u201cand yearly demand for laser-sintering services has increased. We expect interest in titanium parts to follow the same strong demand curve.\u201d\nEnd users are sometimes initially wary of the product, however. \u201cWe did have some trouble persuading our cus-", + "12 MPR September 2009 metal-powder.net\ntomers of the benefits of the technology at first. It sounds very strange at first \u2013 like something out of science fiction \u2013 and some found it hard to believe it could actually work. Another customer issue was the consistency of the finished metal. Many didn\u2019t think that metal powder could create the same consistency as ordinary solid metal, such as rolling mill steel,\u201d Mayer says. \u201cIn fact, using metal powder ensures that the material is finally more consistent and has the same specifications throughout.\u201d\nJordan agrees. \u201cA lot of clients were suspicious of the technology in the beginning because at that point in time they haven\u2019t heard about the technology before and as such did\nnot know the advantages,\u201d she says. \u201cWe needed to convince them. We did this by showing off the scope of this technology through illustrative benchmarks. Customers use these benchmarks in their manufacturing processes and experience the advantages hands-on.\u201d\nCase study \u2013 golf ball In this example, DMLS technology was used to make a give-away golf ball to be produced in large quantities at low costs, Blow moulding extruded polypropylene (PP) combined with the injection of an elastomer was required. Usually the main challenge with this kind of form is venting the\ntool which can, in the worst case, lead to deformed golf balls. The solution was to integrate venting channels with almost invisible openings near the cavity. Due to the choice of the process parameters, the pressure could escape and but the channels did not clog up. The volume of the cavities could be minimised, which helped reduce building time and thus the costs of the DMLS tool. Eight cavities were combined, making up for a four cavity tool, producing more than 20 million golf balls. Building the tool took only 50 hours, and conformal cooling channels increased productivity by 20%.\nCase study \u2013 PE bottles In this case study small DirectTool inserts with conformal cooling channels were built and integrated into a conventionally manufactured tool in order to extract the heat from these parts more quickly. The inserts reduced cycle times from 15 down to 9 seconds. This enables a 75 % increase in productivity for a 4 bottle blow mould without sacrificing on quality. The cycle-time and productivity of this type of tool were limited by the time it took to cool down the bottle necks as wall thickness reaches a maximum there.\nFigure 5 shows three examples using conformal cooling channels. Figure 5a shows a tool for blow-moulding PE bottles, while figure 5b shows a cooling pin for cooling an injection point \u2013 a classical hot spot. Conformal cooling in this case reduces cycle times by two thirds. Figure 5c depicts a core with spiral conformal cooling channels inside the dome. By conventionally milling the lower part of the tool and limiting DMLS to the part with conformal cooling, costs were reduced. A 0.3mm machining allowance was added for finishmachining of the outer surface." + ] + }, + { + "image_filename": "designv11_69_0002347_15397730701404684-Figure6-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002347_15397730701404684-Figure6-1.png", + "caption": "Figure 6. The whole computational mesh of impeller.", + "texts": [ + " It is especially suitable for the cyclic symmetric structure to adopt the substructure technique to make an analysis and meshes only need to be divided for one substructure mode to create the global stiffness matrix, computed results can be taken to other super elements being of the same construction. This gains higher computational efficiency of the substructure mode and leads to a significant reduction of computer cost. The more the repeated structures are, the higher computational efficiency will be achieved. Table 1 gives the overall dimensions of the impeller taken for analysis. The computational model of the impeller is shown in Fig. 4. The mesh of substructure is shown in Fig. 5. Figure 6a shows the side elevation of the finite element mesh of the impeller. Figure 6b shows the front elevation of the finite element mesh of the impeller. In order to truly represent the impeller and improve the accuracy in analysis, the computational model used a four-node tetrahedron solid element with three degrees of freedom at each node. The effects of temperature and pressure loading were not considered, as they are small in comparison with the centrifugal force due to the high speed of rotation (Ramamurti et al., 1995). The centrifugal stiffening effect due to the high speed of rotation is taken into account for the analysis" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002467_bf02133099-Figure3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002467_bf02133099-Figure3-1.png", + "caption": "Fig. 3.", + "texts": [], + "surrounding_texts": [ + "Similar formulae and results can be obtained if system 2 is considered, with the following changes\nH(0, t) = H(L, t) = - rt-2,s + o(/-21s) OH OH 2 . (-~-t)=3_0 = (-~--t)=3.L = - ~ t - s / ' + o ( t - 5 / 3 ) \u2022\nFormulae (4.2), (4.5), (4.6), (4.7), (4.8), (4.11), (4.12) are still valid, whereas formulae (4.13) are substituted by\nRb%IL3 ( 2 ~ 6 r ) )\nl (2e) 3m = 2m.aEH ~ - - 1\nc cosh ln -T- 1 c A = 2 s inh/n + In 2 w l ( r ) ~ 3 w 2 c o l ( r ) - to\n5. Whir ls of l imi ted ampl i tude .\nSelf-excited whirls are possible even if ~u< 2coW); we obtain here the asymptotic formulae which describe them.\nLet us consider first system 1 to prove that, with an appropriate choice of the constant ? and of the functions H(x,), f (xs) and G(x3 ,t), formulae\n1 \"1 = [H(x3) + G(xs, t)] cos (y +f(x , ) t -4) t\nus = [H(x,) + G(xs, t)] sin(? +f(xo)t-a)t (5.1)\nsatisfy Eqs. (2.1), (2.2), (2.3), (2.4) for t - ~ o o . We will suppose, as before, that\nH(O) = c, H(L) = + \u00a2, (5.2)\nwith the same convention regarding symmetric and antisymmetric solutions, and we take\nG(x. , t) = -- ced(x.)/-2 (5.3)\nwith E positive constant and\nd(x.) = 1\nt a(=.) = 1 - - - - 2x3 L\nfor symmetric solution\nfor antisymmetric solution.\nFirst of all we notice that Eqs. (2.3), (2.4) together with (4.10) force the choice\nto\nY = 2\nIntroducing (5.1) in (2.1) and accepting the usual approximation for t - + oo one obtains the equations\nwhere\nH ' V(xs) - - f l4H(xs) = 0\nh ' ~ ( ~ s ) - ~ h ( x s ) = - - ~ S d ( x , )\nyr.4 to2 4 L4(~r ' ) 2 ' S = - - 02\n(5.4)\n(5.5)\nand\nRemembering that\nh(~.) = H ( = . ) f ( x . ) .\ntO tO l F. sin ~ - t X~ = F, cos - ~ - t - -\nt o o) 111 = F. sin - ~ - t + F . cos- - f f t\nto OJ\n1 -\u00a52---- + F, c o s - ~ - t T - F , sin - ~ - t tO to Y2 = + F, s i n - ~ - t + _~,cos-~-- t\nF~ = -- 12nR\"b,tc-u(2r) -1/2\nI F. : 24nRSb~c-2(2~)-l/~ f(O)t -3 F, = 24nRab~c-2(2~)-lt~ f (L)t-s\nEqs. (2.2), (2.3), (2.4) give\nI HH(O) = O,\nEIHH'(O) = F,, H\"(L) = 0 EIHm(L) = -T- F~\n(5.6)\nand\nh, , (o) = o , h,1(L) = 0\nI h-,(0) = ~Mh(0) h , , , ( L ) = - - ~.~h(L)\n(5.7\nwith\n1 ~ c o s h 3 L \u00a5 1 M=-T ( ~ Z\ncos flL ~ 1 sin/~L\nwhere the upper (lower) sign applies to symmetric (antisymmetric) solutions.\nEqs. (5.4) with the boundary conditions (5.2), (5.6) determine the function H(xs) and the constant r.\nc l c o s f L -T- 1 H(xs) = - ~ - cos Bxs + cosh Bxs sin 3 L sin B x s - -\ncosh flL ~ I sinh pxs I (5.8) sinh ilL,\nDECEMBER 1969 339", + "(ilL) s ~cosh flL ~ 1 cos flL ~ 1~ 2 ( s inhflL sin pL )\n~ L = 7t, ~ / co 2~o([)\"\n12#RBb~L 8 c3Ei(2e)z/2 '\n(5.9)\nFurther, from (5.5), (5.7)\n4~ a(xs) = {2cg(xs)-- H(xs)} 02\nhence\nf ( xa )= 4e 12c d(xa) I co H(x3) 1 . (5.10)\nFormula (5.9) shows that, during the special motions studied here, the lubricant forces act as though they were elastic forces with rigidity\nk = 12#RSb)~ (5.11) c3(2~)1/2\nWe will come back to this question in the next paragraph; we explore here some consequences of Eq. (5.9).\nFor the existence of solutions of the type (5.1) the left hand side F(~o/~z(r)) of Eq. (5.9) must be positive..A look at Fig. 1 (where graphs of F(~/e.),(r)) are traced: the dotted lines refer to symmetric solutions and the solid lines to antisymmetric ones) show that a symmetric solution exists when ~o< 2~oa(r), where an antisymmetric solution exists when co< 8coz(r) = 2o~a(r).\nFurther properties of solution (5.1) can be derived from graphs of the function H(xs) #versus xs/L. Graphs are given\n6,10 ~\n10 ~ 5.10 )\n.iO) 5\";OZ . - ' \"\ntO ~ [ '\u00b0iI s l\nI I Q\n{l! 0 )0 ~s 20 25 3o ~\nFig. 1.\nin Figs. 2 and 3 for some values of ~flol(r): Fig. 2 (3) shows the symmetric (antisymmetric) case. It appears that when ~ approaches 20~1(r) the graphs for the symmetric case approach that of the lowest restrained mode; similarly in the antisymmetric case, when ~--> 2to~ (r), the graphs approximate the second mode.\nInstead, at low speeds of rotation, the symmetric (antisymmetric) mode appears as that of a parallel (conical) whirl; then the behaviour of our shaft approximates that of a rigid rotor (See point 2 of Sec. 3 above).\n340 MEGGANICA", + "Results of a similar nature are obtained when system 2 is considered and formulae (5.1), are substituted by\nju, = [H(xs) + C ( x s , t)] cos (y +f (x s ) t - s / s ) t tuu = [H(xa) + G ( x s , t)] sin g, +f(xs ) t - s t~ ) t (5A2)\nwith\nG(\u00d7s , t) = - - \u00a2~d(xs) t -~.\nOne gets\n\u00a2.0\nY = 2\n\u00a2 l cos #L ~ 1 H ( x ~ ) = - ~ - cos t~xa+cosh /~xs - - s i n 3 L s i n ~ x s +\ncosh ~L -T- 1 1 sinh/~L sinh ~x~ (5.13)\nico i cos ; Ii 2 { sinh f lL sin f lL\nr~RbS~L ~\n~/i ~o (5.14) ?~L = ~ 2 o ~\n1\n\u00a2 . (5.15) f ( x z ) = 3 w H(xz ) 5\nAlso in this second case, during the special motions we consider, the force due to the lubricant is equal to an elastic force of rigidity\nk = nRbSo (5.16) \u00a2s(Z~)S~\n6. S o m e remarks.\nWe have already noted that, for t sufficiently large, the forces due to the lubricant act as though they were elastic forces during the special motions considered here. Precisely, when the displacement of the journal follows the law\n\u00a2.O { u~ = \u00a2(1 - - ~t -9 cos - ~ - t\nO) uz = \u00a2(1 - - ~t-*) sin ~ t\n(6.1)\nfor long bearings or\n(0 ul = c(1 - - ~t-,ts) cos T t\n\u00a20 u, = c(1 - - et-2t3) sin ~ t\n(6.2)\nfor short bearings.\nWe can reverse this remark and look for the most general trajectories along which the property now mentioned holds. It must be\nF~ = - - k e , F n = O\nand these equations are satisfied if and only if\n2 (6.3)\nand\nkc 3 c 12nRSb ~ t + 23 = (c 2 ~ e~),/2 e\nfor long bearings or\nlog \u00a2+(cS--e~)tt~ (6.4)\nkc3 23 -- \u00a2(2c2 - - eu) log ~\" \"q- (f2__ g2)1/2 (6.5) =Rb% t + * - - (c2-- e2)3/--------q e\nfor short bearings; here E , 231 are constants. When t is large these formulae give as an approximation\n(6.6) c ( 1 2 u R 3 b ~ 2 1\ne = c--~- \\ ~s / t\u00a5 ~'\ne = e-- -~- t -~\" (6.7)\nBut these relations together (6.3) and (5.11), (5.16) are equivalent to (6.1), (6.2).\nWe want to determine now, how the energy is distributed asymptotically between the kinetic and potential components T and V.\nLet us call tV the work done per unit time by the force due to the lubricant on one of the journals\nIV = F.~ + F.e;;;\nfor t large, as a consequence of (4.7), one gets\n~ / ~ F n c o l ( r ) c .\nSimilarly (see, for instance, [4])\ndT 1 d f L /~Ac J dt 2 dt o g\n2 g\nw 23z d t\nD i n\n2 - - F . = Fno~l'~ c.\nDECEMBER 1969 341" + ] + }, + { + "image_filename": "designv11_69_0000687_eej.4390950213-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000687_eej.4390950213-Figure2-1.png", + "caption": "Fig. 2. Model.", + "texts": [ + " The Electromagnetic Equations and Their Solutions Figure 1 shows the arrangement of attractive electromagnets and ground rails discussed in this paper. The theory developed in this paper holds equally for the flat ground rail. Although the exciting current varies in practice to keep the air gap length constant [3], we assume in this paper that both the exciting current and the air gap length are kept constant; problems of controlling the exciting current and air gap length are beyond the scope of this paper. To facilitate analysis, we consider an analytical model as shown in Fig. 2, in which the electromagnet is assumed to be at a fixed position and the ground rail moves in the direction of x-axis. The electromagnetic 5eld of this model can be determined by solving the following Maxwell equations : VxE=O (2) B P J = V X (-) where V is the rail velocity and J is the current density vector. We solve the above equations under the following assumptions: ground rail is in the direction of the y-axis and the fringing effect in the direction of z-axis is negligible. (ii) The permeability of iron is much larger than that of air" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001384_j.ijmachtools.2006.02.011-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001384_j.ijmachtools.2006.02.011-Figure2-1.png", + "caption": "Fig. 2. Schematics and prototype of the RHA apparatus: (a) schematics of the RHA appararus; (b) prototype of the RHA apparatus", + "texts": [ + " The conventional CNC machining process consists of rough cutting, fine cutting and pencil cutting. Most of the cutting time is consumed during the rough cutting process, because the tool cuts a given workpiece step by step at a constant depth to eliminate the remaining material well and reduce the cutting resistant force. However, the use of a hot tool enables ablation of the workpiece from a raw shape to a part shape at once because of its heat ablation characteristics. The scanning toolpath as rough cutting can be applied to the RHA process for rapid shaping. Fig. 2(b) illustrates the prototype of the RHA apparatus, and Table 1 shows the specifications of the RHA apparatus. The RHA apparatus comprises control software, a CNC controlled motor at each axis, an indexing table, and a hot tool. In the present work, the RHA process sequence was as follows. Table 1 Specifications of rapid heat ablation system 1. Range of X, Y, Z axis X 200mm After the material block was set up on the indexing table, the origin of the tool related to material position was regulated with control software" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001164_1-84628-559-3_20-Figure9-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001164_1-84628-559-3_20-Figure9-1.png", + "caption": "Fig. 9. Photograph of the overall experimental setup", + "texts": [ + " A 16-bit DA/AD converter was used to transform the pulse driving waveform to the power amplifier and then to the PZT actuator. Due to the number limitation of measuring probes, two kinds of gap sensors were used to detect the simultaneous motion behaviors of the positioning stage along three rotation axes. They were the capacitive gap sensor (ADE 5300-5504) and the fiber optic displacement sensors (Fotonic MTI-2100). Their characteristics are, separately, with the bandwidths of 20 kHz and 100 kHz, the measuring ranges of 25/250 m and 10/36 m, and the corresponding resolutions of 5/50 nm and 1/10 nm. Figure 9 shows the overall experimental setup for examining the motion behaviors of the positioning stage. Since the rotation displacement caused by one single actuation of the PZT actuator is expected to be very small, the rotational motion are assumed as the case of linear motion. As shown in the photograph, one capacitive gap sensor was used to measure the rotational motion along zaxis, and two fiber optic displacement sensors were used to measure the rotational motions along x- and y-axis. The works mentioned above focused on examining the motion behavior of the positioning stage actuated by only one single excitation for the PZT actuator" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002541_memsys.2007.4432973-Figure5-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002541_memsys.2007.4432973-Figure5-1.png", + "caption": "Figure 5. Experimental setup for the characterization of the intelligent textile. The left stage moves laterally while the right stage generates rotational movement.", + "texts": [ + " A low-end household sewing machine was used in our case. This machine has one roll of yarns stitched to the fabric from front side and the other roll of yarns stitched from bottom sides. The resulted intelligent textile is shown in Fig. 4. The front side yarns were chosen to be conductive whereas the backside ones were conventional yarns. The conductive yarns were fastened to the surface by the conventional yarns stitched from backside of the fabric. The assembled intelligent textile was tested using a homemade mechanical apparatus as shown in Fig. 5. The fabric was fixed between two clamps mounted on two moving stages. The left stage moves laterally while the right stage generates rotational movement. The assembled intelligent textile was twisted and stretched by this apparatus. The preliminary reliability test was conducted by applying cyclic twisting loading. No signs of failure were observed even after 5000 twisting cycles. The magnitude of the twist was ~30\u00b0/cm. We had to apply excessive forces to break the metal interconnect. The breakage of metal wires at the edge of the silicon islands was found to be the major failure mode" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002819_6.2007-3813-Figure17-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002819_6.2007-3813-Figure17-1.png", + "caption": "Figure 17. Color Coded Grid Areas.", + "texts": [ + " This information was obtained by using the measured pressures to calculate forces and moments generated by the different devices about the projectile. The results were then compared for each device as to its best location and the different actuators could be compared with each other to determine which device develops the most steering forces and holds the most potential for controlling a guided subsonic projectile. Pressures were measured at 60 locations on the surface of the bullet, and each of these pressure taps was assigned an area. These areas are highlighted in gold on the model shown in Figure 17a. Using a SolidWorks model of the projectile, a small area around each tap was constructed using grid lines on the surface of bullet. SolidWorks calculated the distance of each tap from the nose of the projectile and the change in each coordinate direction, (dx, dy, and dz). Using this information, the area surrounding each pressure tap could be calculated along with a vector distance of the centroid of that area to the center of gravity of the projectile. A portion of the 60 pressure taps and their associated areas were also mirrored across the centerline under the assumption of symmetric flow. These mirrored points and areas are shown in green on Figure 17a. While the original 60 areas and the mirrored areas accounted for much of the surface of the model, there were still some gaps. Pressures and areas were also assigned to some of these gaps by averaging the pressures adjacent to the gaps and by creating the additional areas shaded blue in Figure 17b. Page 10 With a pressure assigned to each area, the force and moment generated at each gold tap on the bullet were calculated by Eqs. (1) and (2) respectively ))/(sin(arctan** dzdyApFtap \u2206= (1) CGtap xFM *= (2) where Ftap is the force generated, \u2206p is the pressure at a specific tap, A is the grid area enclosing the tap, Mtap is the moment generated, and xcg is the distance in the coordinate x direction between the tap and the center of gravity. These values were also used as the values for the corresponding mirrored green taps on the bullet", + " For each test case, data collected from the pressure taps was run through a customized Visual Basic program to mirror all of the tap locations that were not on the centerline of the bullet. Also, the Visual Basic program generated x,y,z coordinates for each tap and delta pressures for each tap by subtracting the test data from data obtained for the clean model. Contour plots of the change in surface pressure were then generated using Tecplot which was allowed to interpolate between the original and mirrored points (as Page 11 shown in Figure 17a). These contours clearly show the changes resulting from various devices. Figure 18 shows an example contour plot and illustrates this process of generating the \u2206P contour plots. VIII. Results The reaction control jets generated a significant amount of forces and moments about the projectile. However, due to the location of the jets tested in the vicinity of the relatively complicated geometry of the round in these locations it was difficult to find consistent trends in the data. Spaid and Cassel in Ref" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002762_isie.2007.4374783-Figure5-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002762_isie.2007.4374783-Figure5-1.png", + "caption": "Fig. 5 \u2013 Active circuit during a phase excitation.", + "texts": [ + " Using the half-bridge converter topology to drive the SRG, the excitation period of each phase begins when the controlled switches starts to conduct, the inductance is increasing, the diodes are not conducting, and the phase winding generates a positive counter EMF. The generating period begins when the controlled switches stops to conduct, the inductance is decreasing, the diodes are conducting, and the phase winding generates a negative counter EMF. The voltage over the load is obtained from equation (2). Using the alternative topology, the only remarkable change is that the voltage over the load now is e plus the output voltage of the rectifier bridge. Fig. 5 and Fig. 6 show the active circuit during the excitation and the generation periods of the alternative converter topology. The SRG mathematical model is evaluated for both converters using a computing program which inputs are the phase voltages and the mechanical torque. The outputs are the currents at the phases, the angular speed and the rotor position. Each new set of values for the phase voltages and for the torque is used to feedback the program in order to evaluate the next state. For the 6/4 poles prototype, Fig", + " 8 also shows that there is not voltage over the load from the end of a phase power transfer to the beginning of the next phase power transfer. Therefore, a SRG transfers power in pulses, suggesting the need of an end capacitor to control the load voltage. Then a capacitor is used to continuously supply the load. It can be noticed that the energy stored in the magnetic field of that phase flew to the end capacitor and to the load when the correspondent diode is conducting (Fig. 6). The voltage (Fig. 9) that excites a phase winding is the rectified voltage (VE) that supplies the converter as shown in Fig. 5. During the excitation process the phase switch is conducting and its diode is not conducting because the less impedance path for the current crosses the switch. The diode begins to conduct when the switch ends its conduction window. As a result, the voltage at the winding terminals is a negative back EMF which, added to the rectifier voltage, supplies the capacitor and the load. Fig. 10 shows the variables involved in the electro-magnetic torque production in a single phase. With the alternative converter topology, running at 900 rpm under stable conditions, the system absorbs 81" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003807_amr.139-141.1079-Figure4-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003807_amr.139-141.1079-Figure4-1.png", + "caption": "Fig. 4 3D statics model of the Tapered roller bearing 3811/750/", + "texts": [ + " The geometry of the roller busbar of the ARC convex busbar tapered roller and parameters as in Fig. 3. In this paper, we try to make analysis with different radius of convexity of roller. We will consider the conditions R=4000mm, 5000mm, 6000mm, 7000mm, and 8000mm, where R represents the radius of convexity of roller. Establishing the computational model According to the loading characteristics of the tapered roller bearing 3811/750/HC, for simplicity, we take partial of the model to analyze. The part of the model we chose in this paper is as in Fig. 4. Choosing the Type of the Unit. In using ANSYS to analyze bearing contact stress, choose SOLID45 as Structural entities unit, CONTA174 as interface unit, and TARG170 as Target surface unit. SOLID45 unit is suitable for the analysis of elasticity, creep deformation, dilatancy, stress hardening, large deformation, and Great strain. With degeneration function, it can degenerate as pentahedron or tetrahedron, which is better for making complicated structural unit grid. For our case, under the loading condition, inner ferrule of rolling bearing, outer ferrule, and roller will take elastic deformation" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002819_6.2007-3813-Figure31-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002819_6.2007-3813-Figure31-1.png", + "caption": "Figure 31. The Effect of Mt.", + "texts": [ + " Elliptic Cam The flow fence device, which simulated an elliptic cam, generally generated larger forces and moment than the other devices because its frontal area was significantly larger than the other devices. Two locations for the flow fence were tested as shown in Figure 16 where the first fence location, taps 15, 20, and 28, is in the middle of the cavity region and the second fence location, taps 19, 21, and 31, is next to the downstream cavity wall. For this aft location, the fence shielded the forward facing step of the cavity which resulted in a large region of negative pressure behind the fence especially at higher Mt values like 0.72 as shown in Figure 31. Nevertheless the high pressure region forward of the fence still dominated the overall change in pressure and a positive force was developed on the projectile which increased with tunnel speed. It can also be seen in Figure 31 that the forward high pressure region wrapped around the body further and extended further upstream than for the pins or jets. Again this is likely due to the larger frontal area. When the flow fence was moved forward, the fence was not somewhat shielded by the cavity and thus did not create as large of a flow disturbance. While the basic pressure distribution pattern around the fence was the same for both locations as shown in Figure 32, the aft location created a greater change in surface pressures" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001240_gt2006-90435-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001240_gt2006-90435-Figure2-1.png", + "caption": "Fig. 2 Typical rotor configuration and coordinates", + "texts": [ + "org/ on 01/30/2016 sd \u03b1\u03b1\u03b1 ,, = the torsional displacement of the disk and the shaft sd \u0392\u0392\u0392 ,, = angle rotations of the disk and the shaft about Y axes sd \u0393\u0393\u0393 ,, = angle rotations of the disk and the shaft about Z axes 21 ,, rrr \u03b4\u03b4\u03b4 = magnitude of residual bows of the shaft 1and the shaft 2 ][],[],[ \u03a8\u03a6\u0398 = matrix of the mode shape function iii \u03d5\u03c6\u03b8 ,, = the corresponding shape functions of the shaft element \u03ba \u2032 = shear factor for circular cross-section of the shaft \u03c1 = mass density per unit volume of the shaft 21 ,, rrr \u03c6\u03c6\u03c6 = phase angle between disk eccentricity and residual bow p\u03c6 = pressure angle of gear pair \u03c9 , \u03a9 = whirl speed and spin speed n\u03a9 = the excitation frequencies 21,\u03a9\u03a9 = spin speed of the shaft 1 and shaft 2 t\u03a9 = the tooth passing frequency EQUATION FORMULATION The configuration of the geared rotor-bearing system is shown in Fig. 1. The two uniform flexible shafts are of the same length L with two residual bows of magnitude 1r\u03b4 , 2r\u03b4 and phase angles 1r\u03c6 , 2r\u03c6 . The gear pair is modeled as two rigid disks mounted at a distance 1x . The bearings are modeled as flexible elements with damping and stiffness coefficients denoted as bC and bK , respectively. A fixed reference frame, X-Y-Z, is used to describe the system motion. For a single shaft system with a rigid disk in the Y-Z plane is shown in Fig. 2, and five degrees of freedom V , W , \u03b1 ,\u0392 , \u0393 are considered at each nodal point of shaft. The torsional displacement is denoted by \u03b1 and axial translational vibration is neglected. The components of the system include disks, gear pairs, bearing supports and rotor shafts with residual bow. The linearized equations of motion are derived for each component as follows: 1. Disk and Gear Pair 2 Copyright \u00a9 #### by ASME 2 Copyright \u00a9 2006 by ASME Terms of Use: http://www.asme.org/about-asme/terms-of-use The kinetic energy of a disk for lateral motion given by Shiau and Hwang (1993) is modified to include the torsional kinetic energy" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002792_gt2007-28200-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002792_gt2007-28200-Figure2-1.png", + "caption": "Figure 2. Compressor drive.", + "texts": [ + " The changes in the pressure ratios and efficiencies of all compressors were recorded and analyzed. The test stand included measurements of the inlet and outlet flow, the pressures and temperatures before and after the compressor, the cooling and leakage flows and temperatures, as well as the electrical power. All the six different compressors were tested with the same drive unit, which is capable to reach a power of 200 kW at the speed of 60 000 rpm. The construction of a typical compressor assembly is shown in Fig. 2. loaded From: http://proceedings.asmedigitalcollection.asme.org/pdfaccess.ashx?u The main parts of the compressor are the impeller with the volute, a high-speed electric motor and active magnetic bearings. The active magnetic bearings enable axial movement of the rotor shaft within 0.01 mm accuracy, which can be done while the unit is in operation. The position of the rotor shaft was measured at both ends of the machine and in 5 directions. The axial direction is the most important from the impeller clearance point of view" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002979_978-3-540-70534-5_7-Figure13.5-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002979_978-3-540-70534-5_7-Figure13.5-1.png", + "caption": "Figure 13.5: Self-righting moment of AUV", + "texts": [ + " AUV Design Mako 199 Propulsion is provided by four modified 12V, 7A trolling motors that allow horizontal and vertical movement of the vehicle. These motors were chosen for their small size and the fact that they are intended for underwater use; a feature that minimized construction complexity substantially and provided watertight integrity. Autonomous Vessels and Underwater Vehicles 200 13 The starboard and port motors provide both forward and reverse movement while the stern and bow motors provide depth control in both downward and upward directions. Roll is passively controlled by the vehicle\u2019s innate righting moment (Figure 13.5). The top hull contains mostly air besides light electronics equipment, the bottom hull contains heavy batteries. Therefore mainly a buoyancy force pulls the top cylinder up and gravity pulls the bottom cylinder down. If for whatever reason, the AUV rolls as in Figure 13.5, right, these two forces ensure that the AUV will right itself. Overall, this provides the vehicle with 4DOF that can be actively controlled. These 4DOF provide an ample range of motion suited to accomplishing a wide range of tasks. Controllers The control system of the Mako is separated into two controllers; an EyeBot microcontroller and a mini-PC. The EyeBot\u2019s purpose is controlling the AUV\u2019s movement through its four thrusters and its sensors. It can run a completely autonomous mission without the secondary controller" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001509_detc2005-84462-Figure3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001509_detc2005-84462-Figure3-1.png", + "caption": "Figure 3: Spherical quadrangle", + "texts": [ + " 2(b) are derivable, 0 || || )( )( )( )( )( )( )( = = = i z r i s i yi y r i s i xi x B OA RAB OA RAB i=1,2,3 (6) where denotes the spherical radius of the partial spherical PM mentioned in the Mechanism Description; || denotes the distance between A sR )( r i OA (i) and rotation center Or. Orientation of the Moving Platform In order to calculate the direction cosines of the unit vector O(i)C and O(i)D, the Duffy\u2019s spherical analytic theory is applied here, including both notations and formulas [17]. In the coordinate frame local on the ith limb, O(i)-x(i)y(i)z(i) shown in Fig. 3, the vector )()( ii CO and )()( ii DO are \u239f\u239f \u239f \u239f \u23a0 \u239e \u239c\u239c \u239c \u239c \u239d \u239b \u2212\u2212 \u2212\u2212= \u239f \u239f \u239f \u239f \u23a0 \u239e \u239c \u239c \u239c \u239c \u239d \u239b = 323122312 23122312 323 )()( )()( )()( )()( 3 csscc cscss ss CO CO CO CO z ii y ii x ii ii (7) where )()( ii CO denotes the vector directs from the O(i) to the C(i); x ii CO )()( denotes the projection of the vector )()( ii CO on the x-coordinate; and denote the sine and cosine of angle mns mnc mn\u03b1 , respectively; mn\u03b1 denotes the angle shown in Fig.3; and denotes the sine and cosine of angle js jc j\u03b8 , 3 Copyright \u00a9 2005 by ASME Terms of Use: http://www.asme.org/about-asme/terms-of-use Dow respectively; j\u03b8 denotes the jth exterior angle of spherical quadrangle. \u239f \u239f \u239f \u239f \u23a0 \u239e \u239c \u239c \u239c \u239c \u239d \u239b \u2212+ \u2212+ \u2212 = \u239f \u239f \u239f \u239f \u23a0 \u239e \u239c \u239c \u239c \u239c \u239d \u239b = ZcYcXss ZsYcXsc YsXc DO DO DO DO z ii y ii x ii ii 123312 123312 33 )()( )()( )()( )()( )( )( (8) where 434233423 434233423 434 )( cssccZ csccsY ssX \u2212= +\u2212= = (9) Specially, for the 3-RCRR, \u03b112=\u03c0/2. According to the Eqs. (7), (8) and (9), variable expressions of the coordinates of the point C, D for three limbs are available" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002979_978-3-540-70534-5_7-Figure8.12-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002979_978-3-540-70534-5_7-Figure8.12-1.png", + "caption": "Figure 8.12: Trajectory calculation for differential drive", + "texts": [ + " This routine allows accurate setting of the steering angle between the values \u2013100 and +100. However, most cheap model cars cannot position the steering that accurately, probably because of substandard potentiometers. In this case, a much reduced steering setting with only five or three values (left, straight, right) is sufficient. Drive Kinematics 141 8.6 Drive Kinematics In order to obtain the vehicle\u2019s current trajectory, we need to constantly monitor both shaft encoders (for example for a vehicle with differential drive). Figure 8.12 shows the distance traveled by a robot with differential drive. We know: \u2022 r wheel radius \u2022 d distance between driven wheels \u2022 ticks_per_rev number of encoder ticks for one full wheel revolution \u2022 ticksL number of ticks during measurement in left encoder \u2022 ticksR number of ticks during measurement in right encoder First we determine the values of sL and sR in meters, which are the distances traveled by the left and right wheel, respectively. Dividing the measured ticks Program 8.1: Model car steering control 1 #include \"eyebot" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001870_icima.2004.1384209-Figure3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001870_icima.2004.1384209-Figure3-1.png", + "caption": "Fig. 3 The structure of PRPU limb", + "texts": [ + " Revolute pair(R) 8. Prismatic pair(P) 9. Universal pair(U) Fig.2 The structure of 5UPS/PRPU PMT This work is supported by NSF of Hebei Province, China (Grant #503287) 0-7803-8748-1104/$20.00 02004 IEEE. 304 The moving platform and the stationary platform are connected by five UPS limbs and the PRPU limb. The UPS limbs are driving limbs and the pose of moving platform can be changed through changing the length of every UPS limb respectively. The PRPU limb is a passive limb, which structure is showed in Fig.3. It includes two prismatic pairs, a revolute pair and a universal pair. The prominence characteristic of the limb is the form of the kinematic forward and inverse solutions are very simple. 111. THE KINEMATIC ANALYSIS OF THE PRPu LIMB The PRPPU kinematic limb coordinate system showed in Fig.4, is set by using the D-H method[I3]. The basis coordinate system {O} is same to the stationary platform coordinate system { A } , and { B } is the manipulator coordinate system. The corresponding D-H parameters are listed in Table I" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002364_detc2007-35166-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002364_detc2007-35166-Figure1-1.png", + "caption": "Fig. 1 Cutting tool mechanical model: a) without cutting process, b) with cutting process.", + "texts": [ + " Once the cutting process proceeds, the cutting forces include the normal support force and friction force on the rake surface, and the work-piece supporting force should be considered as well. The supporting force for cutting are modeled through a damper of 2d and a spring of 2k . The mass is subject to an external excitation force ( ) cospF t A t= \u2126 with an angle of \u03b7 off the vertical direction, and A and \u2126 are excitation amplitude and frequency, respectively. Such a mechanical model is shown in Fig.1. Once the cutting process occurs, two additional external forces 1( )tF and 2 ( )tF exerts the mass. The two external forces act on the two surfaces of the 2 Downloaded From: http://proceedings.asmedigitalcollection.asme.org/pdfaccess.ashx?u cutting tool. On the left surface, in addition to the normal force, the friction force between the cutting metal and the cutting tool will be considered. On the bottom surface, only the normal force is modeled. To model the cutting process, the global system ( , )X Y is introduced with the origin at point O, as shown in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000286_20.877817-Figure3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000286_20.877817-Figure3-1.png", + "caption": "Fig. 3. A magnetic flux plot of the analyzed model.", + "texts": [ + " Torque pulsation due to current ripple and cogging torque are calculated by (12) where is the number of incremental integration path on the rotor surface, is the rotor radius, is stator core length, and are the normal and the shear component of average flux density on an incremental integration path [7]. The unbalanced magnetic force acting on the rotor surface is computed by (13) This unbalanced magnetic force has a serious effect on the dynamics of the hard disk drive [8], [9]. A magnetic flux plot obtained from the finite element analysis is shown in Fig. 3. The air region out of the rotor steel shell is included in the analysis because the rotor steel is likely to saturate magnetically. Fig. 4 shows the distribution of magnetic radial force acting on the rotor magnet inner surface. It is observed that the unsymmetric magnetization distribution makes an irregular magnetic force distribution wave form, that is, the force differences in the opposite direction at the magnet inner surface is greater than that of the symmetric magnetization. Therefore the unbalanced magnetic force defined by (13) is also increased" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001643_sicon.2005.257870-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001643_sicon.2005.257870-Figure1-1.png", + "caption": "Fig. 1. Physical Implementation of the Glucose Sensor and Controller System", + "texts": [ + " The key limitation to the successful development of an artificial pancreas is the implantable glucose sensing technology and the electronic support needed to control the instrumentation. This paper describes the design and development of a controller for an implantable glucose sensor that provides continuous and reagent-free optical analysis of interstitial fluid (ISF). The controller, aptly called an electronic support unit (ESU), performs the crucial task of the physical interface between optical sensing elements and glucose related data. The physical implementation of the glucose sensor is illustrated in Fig. 1. The sensor will be implanted in the subcutaneous tissues of the human body and the ESU will enable the sensor to operate for months with minimal user intervention. This new technology relies on the unique optical characteristics of glucose in the near infrared spectrum [1], Copyright 2005 by ISA - The Instrumentation, Systems and Automation Society. Presented at Sicon/05, 8-10 February 2005, Houston, Texas; http://www.isa.org 59 [2] and will be used as the sensing element in a feedback controlled insulin delivery system for the in situ treatment of diabetes" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002879_acemp.2007.4510534-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002879_acemp.2007.4510534-Figure1-1.png", + "caption": "Fig. 1. Equivalent circuits of the IG (magnetizing inductance m is constant).", + "texts": [ + " In the referential axis linked to stator, the following equations are deduced [1]: ++= ++= \u03b2 \u03b2 \u03b2\u03b2 \u03b1 \u03b1 \u03b1\u03b1 e dt di LiRu e dt diLiRu s eqseqs s eqseqs (1) += + \u2212= + \u03b1\u03b2\u03b2 \u03b2\u03b1\u03b1 \u03d5\u03c9\u03d5 \u03d5\u03c9\u03d5 rms r rr r r rms r rr r r i l mrp l r i l mrp l r (2) +\u2212= \u2212\u2212= \u03b2\u03b1\u03b2 \u03b2\u03b1\u03b1 \u03d5\u03d5\u03c9 \u03d5\u03c9\u03d5 r r rrm r rm r r r r l mr l me l m l mre 2 2 (3) Where p is the Laplace operator, and 2 += r rseq l mrrR , seq lL \u03c3= , .1 2 rs ll m \u2212=\u03c3 The mechanical part of the machine is given by: \u2212\u2212= pp DTT J pp dt d m me m \u03c9\u03c9 . (4) With: ( )\u03b1\u03b2\u03b2\u03b1 \u03d5\u03d5 srsr r e ii l mpp T \u2212= . . (5) Where pp is the pole pair number. If the magnetizing inductance m is constant, according to the model of equations (1), (2) and (3), the induction machine can be seen as a two generators with internal sources ae , \u03b2e respectively, and an internal impedance \u03c9eqeqeq jLRZ += (Figure 1). Due to the saturation of the main magnetic circuits, the magnetizing inductance m varies with the phase voltage Since eqR and eqL are depending on the magnetizing inductance m , the induction IG can be considered as a two generators with a variable internal impedance eqZ (Figure 3). wind turbine, r is the turbine rotor radius, and \u2126 is the turbine rotational angular velocity. Figure 4 shows a typical set of wind turbine outputpower/shaft-speed characteristics for a 0.75 kW machine. IV. INPUT-OUTPUT FEEDBACK LINEARISATION The aim of the proposed method is to achieve a simultaneous control of the considered outputs that are the active and reactive powers to be injected in the dc-bus" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003354_optim.2010.5510527-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003354_optim.2010.5510527-Figure1-1.png", + "caption": "Fig. 1. example flux dispersion of electromagnetic machines with rotating magnetic field, here asynchronous machine", + "texts": [ + " In addition the settling time is important in many applications. The control concept presented in this section gives a very fast transient behaviour by adjusting the output voltage of the machine inverter at changes of the desired torque as fast as phyically possible to its new values. The limitation remaining is mainly given by the rating and the switching frequency of the converter used. The priniciple used is similar to the Indirect Stator-Quantities Control (ISC) already known for asynchronous machines (cp. fig. 1). The main difference is the air-gap rotor flux: In case of the ISC the air-gap rotor flux results from the air-gap stator flux and rotor currents resulting from the asynchronous rotation of the rotor with respect to the air-gap stator flux. In contrast the air-gap rotor flux of the permanent magnet synchronous machine is defined by the rotor permanent magnet. The direction of this flux equals the actual orientation angle of the rotor. The principle of flux dispersion of all electromagnetic machines with rotating magnetic field can be displayed based on the example of an asynchronous machine (cp. fig. 1). 383978-1-4244-7020-4/10/$26.00 '2010 IEEE This physical similarity gives the basis for a common control principle for such machines. The difference between the airgap flux of the rotor \u03a8\u03b4,rotor and the air-gap flux of the stator \u03a8\u03b4,stator defines the actual torque as well as the field weakening of the machine. In the special case of a permanent magnet synchronous machine (PMSM) the direction of the main component of the rotor flux is equal to the actual orientation angle of the rotor. The idea of the flux based control of the PMSM is to directly compute the necessary stator flux of the machine to realise the desired torque and (if needed) field weakening" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000428_6.1979-2033-Figure8-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000428_6.1979-2033-Figure8-1.png", + "caption": "FIGURE 8. THE AN-2E: AN ANTONOV AN-2W MODIFIED BY CLST(~)", + "texts": [], + "surrounding_texts": [ + "FIGURE 1. T. KAARIO'S SURFACE-EFFECT VEHICLE \"AEROSANI\" NO. 8(1)\nKey: 1 -forward wings; 2-articulated controlling wings; 3-hull with driver cockpit; 4-side stabilizers; 5-rudder; 6-tail stabilizing beams with planes; 7-flap; 8-main lifting wing; 9-skis\nFIGURE 2. A SWEDISH WATER-BORNE WIG VEHICLE (\"AEROBOAT\") DEVELOPED BY TROENG\nIN THE LATE 193dl)\nKey: 1-hull with crew cabin; 2-floats; 3-stabilizer with controlling surfaces; 4-outboard engines; !&lifting wing; 6-propeller a ~ e ~ i b ! y\nD ow\nnl oa\nde d\nby F\nre ie\nU ni\nve rs\nita et\nB er\nlin o\nn D\nec em\nbe r\n12 , 2\n01 6\n| h ttp\n:// ar\nc. ai\naa .o\nrg |\nD O\nI: 1\n0. 25\n14 /6\n.1 97\n9- 20\n33", + "FIGURE 5. GENERAL ARRANGEMENT OF THE RFB X-113 AM AEROFOIL BOAT(^^)\nFIGURE 6. THE OllMF-2 SINGLE-SEAT WING-IN-GROUND EFFECT RESEARCH CRAFT(^)\nD ow\nnl oa\nde d\nby F\nre ie\nU ni\nve rs\nita et\nB er\nlin o\nn D\nec em\nbe r\n12 , 2\n01 6\n| h ttp\n:// ar\nc. ai\naa .o\nrg |\nD O\nI: 1\n0. 25\n14 /6\n.1 97\n9- 20\n33", + "D ow\nnl oa\nde d\nby F\nre ie\nU ni\nve rs\nita et\nB er\nlin o\nn D\nec em\nbe r\n12 , 2\n01 6\n| h ttp\n:// ar\nc. ai\naa .o\nrg |\nD O\nI: 1\n0. 25\n14 /6\n.1 97\n9- 20\n33" + ] + }, + { + "image_filename": "designv11_69_0001346_1.1947198-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001346_1.1947198-Figure2-1.png", + "caption": "Fig. 2. a Video frame showing a steel ball on the edge of the table; b a schematic showing the ramp, ball, and table.", + "texts": [ + " Then we determine when Ff /FN= s and note the cor- responding values of and \u0307. These are used as the initial values for and \u0307 in Eq. 9 , which is solved numerically. The result for and \u0307 is substituted in Eq. 5 to determine FN /m. When FN /m=0, the ball leaves the table. To calculate the corresponding final rotational velocity of the ball about the center of mass, Ff /m must be calculated from Eq. 6 and then substituted into Eq. 7 , which must also be solved numerically. One half of a steel ball of diameter 2.5 cm was painted white see Fig. 2 a to provide a reference line that was used to determine the angle of rotation of the ball as it rolls on the table and undergoes free fall while continuing to rotate. The launching ramp, ball, and table are shown schematically in Fig. 2 b . To avoid the effect of wood grain on the motion of the ball, a slab of plastic was used as a table top. The ball was initially positioned on the ramp with the white boundary line cross wise to the length of the ramp. The ball was then released and the subsequent motion was recorded using a digital video camera. The video was then transferred to a personal computer and captured using DVVCAP.5 The angle of rotation as a function of time was measured using the video analysis software VIDEOPOINT" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003662_iciea.2010.5516940-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003662_iciea.2010.5516940-Figure1-1.png", + "caption": "Figure 1. Phasor diagram explaining indirect vector control", + "texts": [ + " Once the decouple control has been achieved, the relation between r\u03c8 and isd can be shown as (3) and the calculation of sl is shown in (4). r r r d dt m sdT L i\u03c8 \u03c8+ = . (3) m r r sl sq L i T \u03c9 \u03c8 = (4) where Tr=Lr/Rr is the rotor time constant, 21 / ( )m r sL L L\u03c3 = \u2212 \u2217 is the total leakage factor. It can be seen in (3) that with correct alignment of ids, rotor flux r\u03c8 is only decided by ids, and in steady state, (3) will become r m sdL i\u03c8 = . The rotor flux is directly proportional to current ids. Figure 1 shows the phasor diagram of the rotor flux orientation IVC. The \u03b1 \u03b2\u2212 axis are fixed on the stator; the axis of rotor is moving at the speed of r\u03c9 , the synchronously rotating d-q axes are rotating ahead of the rotor axis by the slip angle sl\u03b8 corresponding to the slip frequency sl e r\u03c9 \u03c9 \u03c9= \u2212 . So we can write: ( )e e r sl r sldt dt\u03b8 \u03c9 \u03c9 \u03c9 \u03b8 \u03b8= = + = + (5) Based on the above analysis, the block diagram of indirect vector control is shown in Figure 2. III. THE EFFECTS OF INACCURATE FIELD ORIENTAITON Generally speaking, in the indirect vector control, as is shown in Figure 2, the feedback current (the observed value, written as sdi\u0302 and sqi\u0302 ) which are converted from ai , bi , ci will always equal to the command currents * sdi and * sqi due to the existence of current regulator, but they may not equal to the real flux component current sdi and torque component current isq unless sdi\u0302 is oriented in the direction of real rotor flux r\u03c8 and sqi\u0302 is established perpendicular to it [5]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001340_biorob.2006.1639242-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001340_biorob.2006.1639242-Figure1-1.png", + "caption": "Fig. 1 Analyses of PARM and ACA mechanisms.", + "texts": [ + " Analysis of PARM mechanism Subheading Each air pressure inside PARM and ACA is maintained constant by means of a servo valve, even with changes in the displacement D. For that reason, appearance of passive stiffness on PARM ought to produce a change in the output F of the respective gum artificial muscles against a change in D. This feature can be structurally explained, by modeling PARM as shown in Figure 2. In the model, the arrows in the figure indicate stretching of surfacing, and the generated force F is produced by the stretching concentrating on end parts. Figure 1 shows the analyses of PARM and ACA mechanisms serving for reference. In PARM, the generated force F = F1 \u2013 F2 P1 \u2013 P2, and passive stiffness F/ D P1*P2. Here, passive stiffness means the stiffness produced without any change in two air pressures P1, P2 on a servo valve. In ACA on the contrary, the generated force F = F1 \u2013 F2 P1 \u2013 P2, but passive stiffness F/D = 0. From the figure, you can see that a problem is that \u201cpassive stiffness\u201d becomes zero with a double-acting type ACA in a general state of pneumatic control" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003349_s00216-009-2859-9-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003349_s00216-009-2859-9-Figure1-1.png", + "caption": "Fig. 1 The scanning electro chemical microscope time of flight experimental arrangement", + "texts": [ + "1 M sodium hydroxide solution or pH 7.4 phosphate (S\u00f6rensen) buffer solution was used as the background electrolyte. All the solutions were prepared with deionized water (specific conductivity less than 0.5 \u03bcS cm\u22121). SECM time of flight (SECM-TOF) diffusion coefficient measurements were made in 0.5% agarose gel prepared with buffer solution and in different sand slurries made with buffer or sodium hydroxide solution. Measuring with the SECM-TOF method The schematic outline of the measurement setup is shown in Fig. 1 It can be seen that the detector, the micro-sized working electrode (electrocatalytic copper microdisc or amperometric glucose enzyme electrode), and the sample source (micropipette filled with the glucose solution) were set close to each other. A pressure-delivery Tygon tube connects the output side of the loop injector to the dropletdelivery micropipette. Nitrogen gas from a cylinder produced the overpressure pulse. The loop injector was operated manually. The \u201cflight distance\u201d between the source and the detector is represented by d" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003512_s0012500810050010-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003512_s0012500810050010-Figure1-1.png", + "caption": "Fig. 1. Experimental setup for the synthesis of PANI in sc CO2: (1) autoclave, (2) liquid thermostat, (3) reaction mixture, (4) valves, (5) capillaries, (6) pressure gages, (7) hand pressure generator, and (8) CO2 cylinder.", + "texts": [ + " Traditional methods of production of PANI involve the oxidative polymerization of the monomer with the use of ammonium persulfate (NH4)2S2O8 as the oxi dant or electrochemical synthesis, which is currently more popular since it offers significant advantages over the chemical synthesis [7]. In the present work, we developed for the first time the method of synthesis of PANI through oxidative polymerization of aniline in sc CO2 and characterized the products by spectroscopy. The setup for synthesis by means of sc CO2 is shown in Fig. 1. It is equipped with a generator creat ing a pressure of up to 35 MPa, which is connected through a system of capillaries with a stainless steel autoclave with a working volume of 30 mL. The setup is equipped with pressure gauges for controlling the sc CO2 parameters. Carbon dioxide CO2 is fed to the cell through a system of valves. The desired tempera ture is maintained by means of a thermostat in which the autoclave is placed. Carbon dioxide with a purity of 99.995% (State Standard GOST 8050 85) was used in the synthesis of PANI" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001217_20060920-3-fr-2912.00027-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001217_20060920-3-fr-2912.00027-Figure2-1.png", + "caption": "Fig. 2. Instrumented needle", + "texts": [ + " We suppose that the only informations at disposal are the position or the velocity of the needle tip and the interaction forces measured by a force sensor. These can be obtained if the needle is instrumented or hold by a robotic assistant. To estimate needle insertion models we use a PHANToM haptic device from Sensable Technologies as an instrumented passive needle holder. The PHANToM end effector is equipped with an ATI Nano17 6 axis force sensor. A needle holder is mounted on the force sensor, so that needles of different sizes can be attached (see figure 2). The PHANToM encoders are used to measure the motions of the needle, with a precision of 30 \u00b5m. During a manual insertion, the velocity of the needle tip is generally very low. Since it is derived from position encoders, it is corrupted by an important quantization noise. To reduce its effect we estimate the velocity with a standard Kalman filter. Measurements are acquired at a frequency rate of 1 kHz, under real-time constraints imposed by the software implemented on Linux RTAI operating system" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000370_iros.1991.174444-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000370_iros.1991.174444-Figure1-1.png", + "caption": "Fig. 1. Geometry of a three-link manipulator and a task.", + "texts": [ + " We note that the optimality constraint (5 ) consists of local minima and local maxima. Therefore. one must determine either local maxima or local minima depending on the objective function whether it is to be maximized or minimized. The effects of local minima and maxima on the performance of a manipulator is not well defined yet in optimality constraint-based schemes so that we will discuss this property in the next section. 2.2. Manipulability Constraint Locus Consider a three-link planar manipulator in a horizontal plane shown in Fig. 1. We choose the position of the end-effector in 2-D space described in Cartesian coordinates, accordingly x E R 2 so that the degree of redundancy at nonsingular points is equal to one. In Fig. 1, a circle is the task used in this paper. The manipulator has link lengths, 1 , = 3, l 2 = 2.5, and l 3 = 2 units. If we denote s 1 = c 1 = cos(O1), sI2 = sin(8,+62), and c Iz = cos(61+62), the kinematic equations are Using (I I), the MCL for the three-link planar manipulator is obtained as Fig. 2. For the case of a three-link planar manipulator, the configuration space is a product space formed by the individual joint manifolds such that a 3-torus as T3. However, the MCL and the basis of the null space of J are", + " In the numerical examples, the cyclic tasks are described as -r cos(2 U 1 ) + c, where r is the radius of the circle to be camed out and c, is the x-axis position of the center of the circle. The task is to rotate the circle of r unit radius, centered at (e, , 0). in unit time, in counterclockwise, thus the initial position at to = 0 is ( c , - r , 0). And the farthest position from the base of the manipulator and the initial position is at t = 112, therefore, we denote that as mid-position of a cyclic task, which is (c,+r, 0). Consider a task denoted as Task 1 where r = 1 and c, = 3 units. As shown in Fig. 1, Task 1 is placed between W-sheer 2 and W-sheet 4(a). There is one self-motion manifold comsponding to the initial position, i.e., (2. 0) in W-sheet 2, and one to the mid-position of the task, i.e., (4, 0) in W-sheer 4(a). Fig. 4 shows the MCL and self-motion manifolds for Task I . In this example, there exist twelve optimal configurations as shown in Fig. 4. But, the manipulator posture for the three-link manipulator is symmetric with respect to the origin of the O r e 3 plane, so that we may only consider the case of e3 2 0 that is an upper-arm posture with respect to link 2" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001817_6.1971-436-FigureI-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001817_6.1971-436-FigureI-1.png", + "caption": "Fig. I. TEMPERaTURE FIELD", + "texts": [ + " The r e s u l c s of this a n a l y s i s can be used t o w r i t e t h e f o l l o w i n s expression: R~ - (TTA,)-~ I n C Z ( I + c o s e l ) / s i n e l ~ + (nh3)- l I\" C Z ( I + cosa2) i s ine2 I ( 4 ) f o r t h e t o t a l c o n s t r i c t i o n r e s i s t a n c e across t h e c y l i n d e r wheil t h e h e a t e n t e r s through t h e c o n t m t area suhtending an angle 0, et t h e center of t h e c y l i n d e r , and leaves t h r a u i h a secund c o n t a c t a w a @ v3 E Fig.4. SCHEMATIC OF CONTACT AREAS Subtendine an angle a2 Fig. i. The rcmlrindei- of ilie c y l i n k r boiindsry , impervicus t o h:at f lzm. The rhernnl cn-ductiv! . o f the cyl inder is i.. Tie t ~ t a 1 c o n s r r i c t : o n r e s i s t a n c e of R t y r i c a i .w hea t chance l i s , t h e r e f o r e , R = R + R + R (5) 1 2 3 .J-- E l a s t i c Contact Between Cyl inde r and- E l a s t i c i t y theo ry (3 ) shows t h a t t h e h a l f - w i d t h of t h e c o n t a c t area between t h e plane s o l i d 1 and t h e c y l i n d e r i s g i v e n by a1 = i- (61 n h e r e f is the force per u n i t l eng th of c y l i n d e r , D i r t h e c y l i n d e r d i a n e t e r , and kl a" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000031_s0307-904x(81)80025-x-Figure6-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000031_s0307-904x(81)80025-x-Figure6-1.png", + "caption": "Figure 6", + "texts": [ + " Because p(co) w2/ku(O) is usually negligible in practice, only a slight drop in radial stiffness could be expected. However, the dynamic stiffness o f the rubber and nonuniformity effects o f a real tyre increase with w: . To take account of this and to avoid the difficult problem of determining the stiffness coefficient ku(w), we simply put k u ( w ) = 1.5p(w) co: + ku(O) and ku(co) = 1.Sp(co) co2 + ku(0) and Ep(co) = Ep(O)(1 + co:1400) On the other hand, the belt tension increments produce distinct changes o f the lateral and torsional stiffnesses. Tiffs is demonstrated for the 175/70 R 13 tyre in Figure 6. O n t y r e g e o m e t r y Let the equator radius o f the 185 SR 14 tyre vary and let the belt width be leg = 2Lx (Figure 1). The corresponding courses o f the lateral stiffness C L = [dF(w)/dw]w=o and the average radial stiffness C R = Flu, where F = 4.9 kN and u is the corresponding radial deflection, are shown in Fig~tre 7. Both functions CL and C R attain their maxima at aL ~ 310 mm and aR ~ 295 mm. The optimum geometry (with respect to these two stiffnesses only!) is characterized by the compromise radius ao, aR < ao < aL, that may be constructed, for example, by means o f game theory3 a R e f e r e n c e s 1 Frank, F" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003238_20100802-3-za-2014.00006-Figure4-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003238_20100802-3-za-2014.00006-Figure4-1.png", + "caption": "Fig. 4. Hardware Simulator.", + "texts": [ + " First, let G(s) := 1 s a0 sv\u22121 + av\u22121sv\u22122 + \u00b7 \u00b7 \u00b7 + a1 (33) Then, choose a small \u03c1 > 0 such that a disk D(\u22121/\u03c1l\u2212 ,\u22121/\u03c1l+) is disjoint from the Nyquist plot of G(s) and the plot does not encircle the disk (Back and Shim, 2008). Add to simple NDOB, saturation function and deadzone function are used: q\u0307 = Aa\u03c4q + a0 \u03c4v By, (34) p\u0307 = Aa\u03c4p + a0 \u03c4v B(u + (1 \u2212 \u03c1)d\u0304(\u03c6)), (35) w = f\u0304(s\u0304x(q)) + g\u0304(s\u0304x(q))ur, (36) \u03c6 = p1 \u2212 \u03c1q\u0307v, (37) u = s\u0304(\u03c6) + \u03c1w, (38) where d\u0304(x) = x \u2212 s\u0304(x) From Back and Shim(2008), s\u0304x\u2019 level and s\u0304(x)\u2019 level satisfy s\u0304x(x) = x,\u2200x \u2208 Ux, and \u2223 \u2223 \u2223 \u2223 \u2202s\u0304x \u2202x (x) \u2223 \u2223 \u2223 \u2223 \u2264 k0,\u2200x \u2208 Rv, (39) s\u0304(s) = s,\u2200s \u2208 S\u03c6, and0 \u2264 \u02d9\u0304s(s) \u2264 1,\u2200s \u2208 R. (40) To simulate our method, we used hardware simulator (Fig. 4). To control hardware simulator, we used MITSUBISHI PLC Q02H. MITSUBISHI support the ethernet communication library and software, MX component. We used Microsoft visual C++ MFC, Using a MFC and MX component, MITSUBISHI GX developer Version 8.03D, and MITSUBISHI MX component Version 3.12N. We can receive data from PLC to computer. After receive data, computer calculate next input value and send to PLC. It takes almost 28\u223c34ms, from read a data to give a input to PLC. Therefore, we set 1 cycle \u2206t as 40ms", + "9478 \u00d7 105. When all strip go to lower roll, hardware simulator change direction. Then, lower roll become upper roll and upper roll become lower roll. The initial values are measured value at this moment. To demonstrate that NDOB can recover the time trajectory of the nominal closed-loop system, we simulate with \u03c4 = 0.0001, \u03c4 = 0.004, and \u03c4 = 0.04. In MATLAB simulation we added two disturbances. They are d = 0.4 sin(2t) and Rlow = R + (\u03c9op/2/\u03c0 \u2217 h)t. Rlow is radius of low roll and h is strip depth. As Fig.4, The radius of lower roll increase as low roll move. A solid line is result of nominal response(Fig.5). As \u03c4 is smaller, NDOB is faster to recover the time trajectory of the nominal response. We simulated when \u03c4 = 0.0001, but in this case simulation result covers nominal result. Therefore, we remove the result when \u03c4 = 0.0001. In the hardware simulator, there exist time delay. As we mentioned in previous chapter, it is 28\u223c34ms. Therefore, we compared simulation result and experiment result, when \u03c4 = 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003204_amm.44-47.1482-Figure3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003204_amm.44-47.1482-Figure3-1.png", + "caption": "Fig. 3 Three types of overlapping mechanism", + "texts": [ + "5hf< < In this case, tracks in 1n \u2212 layer and 1n + layer respectively do not overlap. To produce continuous layers, the adjacent tracks in n layer must overlap, then: 2 1/ 2 ( ) 4 2 2 h H h D D S f H \u2212 = \u2265 (1) According to geometric relationship, when powder thickness is h , then: s h S D h H + \u2206 = + \u2206 = (2) Substituting Eq. 2 to Eq. 1, relationship between sf and hf can be obtained: 1/ 2 1s hf f+ \u2265 (3) The continuity of layers in this case depends on sf (or s\u2206 ). It is called intra-layer overlapping mechanism in this paper, as Fig. 3 shown. Case 2: 0.5 1hf< < In this case, tracks in 1n \u2212 layer and 1n + layer respectively overlap, then: 2 2 1/ 2 1/ 2 ( ) ( 2 ) (2 1) 4 4 2 2 2 h h H h D H h D D D S f f H H \u2212 \u2212 + = + \u2212 \u2265 (4) Substituting Eq. 2 to Eq. 4, relationship between sf and hf can be obtained: 1/ 2 1/ 2(2 1) 1s h hf f f+ + \u2212 \u2265 (5) When sf (or s\u2206 ) is small, the continuity of layers in this case depends on hf (or h\u2206 ). It is called inter-layer overlapping mechanism in this paper, as Fig. 3 shown. Note that when sf (or s\u2206 ) is large enough, it may coexist with intra-layer overlapping mechanism, which is called mixed overlapping mechanism in this paper. The depositing efficiency in SLM process is defined as the actual volume of formed part to the whole volume of tracks. If the actual volume is constant, the amount of tracks fluctuates according to intra-layer overlapping rate and inter-layer overlapping rate. Note that once the intra-layer overlapping rate and the inter-layer overlapping rate vary, the overlapping mechanism may change", + " The overlapping topography was observed by optical microscope; the intra-layer overlapping rate and the inter-layer overlapping rate were calculated when the width and the depth of tracks were measured. Fig. 5 showed the overlapping topography of Sample One. It could be seen that Track 1 and Track 2, in the same layer, overlapped while Track 2 and Track 3 also overlapped. Measurement and calculation results were as: the intra-layer overlapping rate 0.43sf \u2248 and the inter-layer overlapping rate 0.42hf \u2248 , which satisfied Eq. 3 and lay in the region \u201cabc\u201d of Fig. 3. Fig. 6 showed the overlapping topography of Sample Two. As shown, Track 4 and Track 5 were in the same layer and hardly overlapped, but lapped over Track 6 in the previous layer respectively. Measurement and calculation results were as: the intra-layer overlapping rate 0.04sf \u2248 and the inter-layer overlapping rate 0.55hf \u2248 , which satisfy Eq. 5 and lay in the region \u201cbgd\u201d of Fig. 2. By comparing the samples, it could be seen that the continuity of layers of Sample One depended on sf (or s\u2206 ), the type of intra-layer overlapping mechanism, and the continuity of layers of Sample Two depended on hf (or h\u2206 ), the type of inter-layer overlapping mechanism" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003275_detc2009-86970-Figure5-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003275_detc2009-86970-Figure5-1.png", + "caption": "Figure 5. Geometry of testing gear tooth.", + "texts": [ + "org/about-asme/terms-of-use 3 Copyright \u00a9 2009 by ASME ML: \u23a5 \u23a5 \u23a6 \u23a4 \u23a2 \u23a2 \u23a3 \u23a1 + \u2212\u2212 =\u23a5 \u23a6 \u23a4 \u23a2 \u23a3 \u23a1 = LMRMbMLMbM LMLMbMLMbM ML M ML MML M rr rr y x ___ ___ )( )( )( sincos )cossin( \u03b8\u03be\u03b8 \u03b8\u03be\u03b8 R , (2) \u23a5 \u23a5 \u23a6 \u23a4 \u23a2 \u23a2 \u23a3 \u23a1 \u2212 \u2212 = \u23a5 \u23a5 \u23a6 \u23a4 \u23a2 \u23a2 \u23a3 \u23a1 = LM LM ML yM ML xMML M n n _ _ )( )( )( sin cos \u03b8 \u03b8 n , (3) where iNq MMLMLM )2(__ \u03c0\u03be\u03b8 \u2212+= . (4) MR: \u23a5 \u23a5 \u23a6 \u23a4 \u23a2 \u23a2 \u23a3 \u23a1 + \u2212 =\u23a5 \u23a6 \u23a4 \u23a2 \u23a3 \u23a1 = RMRMbMRMbM RMRMbMRMbM MR M MR MMR M rr rr y x ___ ___ )( )( )( sincos cossin \u03b8\u03be\u03b8 \u03b8\u03be\u03b8 R , (5) \u23a5 \u23a5 \u23a6 \u23a4 \u23a2 \u23a2 \u23a3 \u23a1 \u2212= \u23a5 \u23a5 \u23a6 \u23a4 \u23a2 \u23a2 \u23a3 \u23a1 = RM RM MR yM MR xMMR M n n _ _ )( )( )( sin cos \u03b8 \u03b8 n , (6) where iNq MMRMRM )2(__ \u03c0\u03be\u03b8 ++= . (7) 2.2 Mathematical Model of Testing Gear T\u03a3 Figure 5 shows a single tooth of the testing gear T\u03a3 which comprises the left and right side tooth flanks TL and the TR in coordinate system ),( TiTiTi YXS . Herein, pTr and tTr are the radii of the pitch circle and addendum circle, respectively. In order to simulate the error of pressure angle on different sides of tooth flanks, the parameters of the left and the right side tooth flanks are assigned individually. As shown in Fig.5, LT _\u03b1 and RT _\u03b1 represent the pressure angles, while LbTr _ and RbTr _ denote the base radii of the left and the right side tooth flanks, respectively. Otherwise, LT _\u03be and RT _\u03be are the involute profile parameters, which determines the position of point on involute curves TL and TR, and TN denotes the tooth numbers of the testing gear T\u03a3 . origin TO of coordinate system TS . The offset EOO eT = indicates the radial eccentricity of the testing gear. According to the Figs.5 and 6, the tooth flanks and their unit normals can be obtained in coordinate system eS by TL: \u23a5 \u23a5 \u23a6 \u23a4 \u23a2 \u23a2 \u23a3 \u23a1 ++ \u2212\u2212 =\u23a5 \u23a6 \u23a4 \u23a2 \u23a3 \u23a1 = Err rr y x LTLTLbTLTLbT LTLTLbTLTLbT TL e TL eTL e _____ _____ )( )( )( sincos )cossin( \u03b8\u03be\u03b8 \u03b8\u03be\u03b8 R , (8) \u23a5 \u23a5 \u23a6 \u23a4 \u23a2 \u23a2 \u23a3 \u23a1 \u2212 \u2212 = \u23a5 \u23a5 \u23a6 \u23a4 \u23a2 \u23a2 \u23a3 \u23a1 = LT LT TL ye TL xeTL e n n _ _ )( )( )( sin cos \u03b8 \u03b8 n (9) where iNq TLTLTLT )2(___ \u03c0\u03be\u03b8 \u2212\u2212= , (10) and ( ) LTTLT Nq __ inv2 \u03b1\u03c0 += " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003344_20100901-3-it-2016.00278-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003344_20100901-3-it-2016.00278-Figure1-1.png", + "caption": "Fig. 1. Pendubot with inverted outer link.", + "texts": [], + "surrounding_texts": [ + "The pendubot is an underactuated double pendulum system, introduced by Block and Spong (1995), for which the (stationary) pivot of the inner pendulum (arm) is fully actuated and the outer pendulum is completely free, controlled only through the motion of its pivot. In a sense, the pendubot is an inverted pendulum on a cart system where the cart moves in a vertical circle rather than on a horizontal track. With this in mind, we will refer to the inner and outer links as the inner arm and the (inverted) pendulum, respectively. Relative to horizontal motion of the pendulum pivot, there is a kinematic singularity when the inner arm is horizontal that results in a loss of linear controllability at the corresponding inverted pendulum equilibrium point(s). That is, the linearization of the system about each of the corresponding stationary (equilibrium point) trajectories is not controllable. This linear controllability loss makes sense intuitively: without the ability to move the pivot point to \u22c6 This work was supported in part by AFOSR under grant FA955009-1-0470. the other side of the (vertical projection of the) center of mass, we have lost the obvious way to affect the angular acceleration of the pendulum. An innocent enough question is then: do there exist (short-time) trajectories between any two states in a (small) neighborhood of such a special equilibrium point; in other words, is the pendubot nonlinearly controllable at a kinematic singularity point? Despite great efforts over the years to develop necessary and sufficient conditions, and appropriate algorithms for testing them, the search continues. Furthermore, many of the conditions that do exist are specialized on systems with special properties, e.g., driftless, homogeneous, etc.\u2014 properties which the pendubot system does not seem to possess. Having little success (and less skill) working with many of the Lie bracket strategies, we decided to approach it as a question about trajectories. With some intuition (and lucky guesses), a strategy emerged for showing controllability. In this paper, we explore this question with an emphasis on trajectory exploration. In particular, we will show how one can find families of linearly controllable trajectory loops based at the the kinematic singularity. Coron (2001) refers to this approach as the return method. It should be noted that specialists in the field of controllability are quite skilled at using expansions involving brackets to determine controllability in many cases. These approaches typically make use of nilpotent approximations Hermes (1991) to the system vector fields, see Kawski (2001) for a (more) accessible introduction. While initially, the pendubot appeared to quickly yield to these tools, further investigation suggests that managing the good and 978-3-902661-80-7/10/$20.00 \u00a9 2010 IFAC 114 10.3182/20100901-3-IT-2016.00278 bad brackets Sussmann (1987) may require some delicacy Kawski (2010). It is likely that the strategies developed in Kawski (1987, 1990) can be adapted to this example." + ] + }, + { + "image_filename": "designv11_69_0000477_s1474-6670(17)37108-2-Figure2.1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000477_s1474-6670(17)37108-2-Figure2.1-1.png", + "caption": "Figure 2.1 Yacht Construction & Coordinate System", + "texts": [ + ", 1979) since they were initially investigated by Astrom and Kallstrom, (1967). Autopilot research for smaller motor-powered marine vessels has been undertaken more recently by Vaneck, (1997), Kose and Gosine, (1995) and Polkinghome, et al., (1994), but it appears that adaptive control of yacht motion has so far received little attention. The hydrodynamics of a yacht are determined by its special structure, comprising its main forcing elements: sails, hull, keel and rudder. A typical yacht is shown in Figure 2.1. Figure 2.2 presents the forces acting on a yacht that is beating into the wind. Here Flirt is the lifting force on the yacht due to the wind on the sail, Fdrag is the drag force of sails passing through the wind. and Fs is the sum of Flirt and Fdrag . The coordinate system is defined by having x lying along the heading direction. The symbols used are defined as follows: J:F;,. J:Fy. I:Fb: total driving force, side force and venical buoyancy force; I.Mx. I.My. I.Mz: total heeling moment, pitching moment and yawing moment; M~: damping moment determined by heeling and yawing motion; u, v" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003086_2009-01-2121-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003086_2009-01-2121-Figure1-1.png", + "caption": "Figure 1. The inertia force of a single cylinder engine", + "texts": [ + " SAE Customer Service: Tel: 877-606-7323 (inside USA and Canada) Tel: 724-776-4970 (outside USA) Fax: 724-776-0790 Email: CustomerService@sae.org SAE Web Address: http://www.sae.org Printed in USA Fs = mr\u03c92(cos\u03b8 + r/l*cos2\u03b8) Fs = 4mr2\u03c92/l*cos2\u03b8 Generally, there are a lot of excitational sources for a powertrain, and they can be divided into three category: torque fluctuation due to gas explosion inside a cylinder, inertia force and moment caused by the movement of internal unbalance masses. In case of a single cylinder engine, as Figure 1 shows a simplified model, the inertia force due to the vertical movement of a piston can be expressed as , and this can be extended into for an in-line 4 cylinder engine by summing the inertia forces of all cylinders, since the phase-alignment of a crank shaft between cylinders can cancel out the 1st order component leaving only the 2nd-order component [2, 3]. As we can see, the vertical inertia force is proportional to the square of rotational speed, and it becomes much larger and dominates the 2nd-order motion of a powertrain at higher engine speed" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000801_j.jmatprotec.2004.07.013-Figure4-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000801_j.jmatprotec.2004.07.013-Figure4-1.png", + "caption": "Fig. 4. Specific power loss and flux density distributions under PWM excitation at core back flux density of 1.3 T.", + "texts": [ + " he error for localised flux and loss measurements were esimated to be within \u00b15 and \u00b18% of measured value. . Results and discussion Localised flux and loss distribution under sine and PWM oltage excitation in locations behind an arbitrary tooth and lot were found using the measurement system described bove. The tangential and radial flux density variation with ime at locations behind a slot and a tooth under sine and WM voltage excitation were measured at peak core back ux density of 1.3 T. Fig. 3a, b and Fig. 4a, b illustrate flux ensity and power loss at positions S1\u2013S4 under sine wave nd PWM excitations, respectively. The modulation index nd the switching frequency remained constant at 0.7 and .5 kHz, respectively. The flux density distribution under PWM voltage excitaion is similar to that under sinusoidal excitation behind both lots and teeth as shown in Figs. 3a and 4a. The peak value f resultant flux density in both S and T locations decreased towards the outer region of the stator core. This is due to core geometry, which influences the flux and loss distribution [7]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000677_ps.2780060511-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000677_ps.2780060511-Figure2-1.png", + "caption": "Figure 2. Speed of response of the chromogenic strips (strips exposed to 0.3 part/l06 PH3). Symbols as for Figure 1.", + "texts": [ + " In all the concentrations of phosphine tested the intensity of colour developed by the strips impregnated with a mixture of mercuric chloride, cresol red and dimethyl yellow was far more than the strips impregnated with either mercuric chloride plus cresol red or mercuric chloride Indicator for phosphine 513 plus dimethyl yellow. At the lowest dosage of phosphine tested (i.e. 0.05 part/l06) and at concentrations up to 0.3 part/l06 strip (c) was three times as sensitive as the other strips. The development of colour was also much quicker in the case of strip (c) than with the strips containing single indicators (Figure 2). The speed of colour development was more or less directly proportional to time of exposure in the case of strip (c), whereas there was no correlation between the duration of treatment and intensity of colour developed with cresol red or dimethyl yellow impregnated strips. When the strips were exposed to 0.3 part/106 phosphine the intensity of the colour developed after 10 min exposure with strip (c) was twice and after 30 min was about three times that of strips (a) and (b). When tested for a shelf-life of 85 days, strip (c) maintained its sensitivity more or less at the same level during the test period (Figure 3)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003879_ac60206a006-Figure3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003879_ac60206a006-Figure3-1.png", + "caption": "Figure 3. Logarithmic plot of percentage conversion of DPPH to A and R at 10 minutes after addition of BHPO vs. concentration of BHPO (initial concentration of DPPH in benzene 2 mmoles per liter)", + "texts": [ + " A or R,B,4: R concentration after addition of BHPO (4 rnrnoles per liter) spectrum obtained when the reaction is still taking place is a hybrid of the quintet and the triplet. The slight difference in g-factor between the triplet and the quintet makes this hybrid spectrum asymmetric. The concentration of the transformed free radical R (giving rise to the triplet) can be estimated by an analysis of this spectrum. The relation between the concentration of added BHPO and the percentage of decomposed and transformed DPPH, is shown in Figure 3. It can be seen that the rate of decomposition of DPPH shows a greater dependence on the concentration of BHPO than the rate of the transformation of DPPH. The intensities of the ESR spectra of the solutions of DPPH, a t various concentrations and before and after t.he addition of BHPO, are compared when the concentration of BHPO is the same in each case-i.e.; 2.5 molea per liter. I t is shown in Figure 4. ..I solution of DPPH exhibits ail eschange interaction phenomenon when the concentrat,ion is greater than some value between 2 and 9 mnioles per liter ( 1 ) ", + " On the basis of these results, the reaction between DPPH and hydroperoxides will be formulated as: DPPH + 0.2 HPOZ + X -+ A + 0.2 HP02:kl (1) (2) DPPH $- 1.5 HPOz + X:kz where, HPOs denotes the hydroperoxide, X denotes some decomposition intermediate, and -4 denotes the final nonparamagnetic species, and kl and IC2 are rate const'ant,s, k1 being the overall rate const,ant for the reaction, DPPH -+ A (3) The proportions of HPOz appearing in Equation. 1 and 2 were determined from the slopes of the curves in Figure 3. The values of kl for various hydroperoxiden were estimated from the initial slopes of the curves in Figure 1 and Figure 2. The initial velocities for the reaction 1 and 2 will be, (DPPH) = -kl (UPPII), (HP02),0.* dt ( ' L ' ) (DPPH! = -kz (DppH)01.5 (2') dt If the ratio of initial rates of disappearance of D P P H in reaction 1 and 2 is approximated by the ratio, I/?, of the amount of DPPH which disappeared in reactions 1 and 2 after the reaction for the first 10 minutes, then (4) Thus Ls can be estimated from kl and T" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003273_msf.628-629.679-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003273_msf.628-629.679-Figure1-1.png", + "caption": "Fig. 1 Schematic diagram of squash presetting method Fig. 2 Instrumentation plan of Sj and Sf", + "texts": [ + " Especially, it has excellent advantages in developing novel cladding alloy systems [1-2]. The presetting method of cladding powders plays an important role in improving qualities of presetting laser cladding coatings, which directly influences the service performance of the coatings. Aiming at the shortcomings of current presetting methods, a novel presetting method adopting squash technique is presented in this paper, which uses the principle of powder metallurgy for reference. The schematic diagram of squash presetting method is shown in Fig.1. The basic principle of the presetting method is as follows. First of all, place powders awaiting cladding into a squash mould, and then press powders into piece with pressure equipment, and at last attach powders piece to the surface of substrate. The powders can be directly pressed onto and adhered to the surface of substrate, or would be also changed into a piece in the mould at first and then fixed onto the substrate surface with fixturing unit just as shown in Fig.1. It is demonstrated by lots of experiments that the squash presetting method may effectively avoid many disadvantages of current presetting methods. It has many good characteristics, such as high powder utilization ratio, good adaptability to powders size, avoiding importing impurity into the coating, etc. Moreover, the qualities of coatings which prepared by using the presetting method is quite excellent. Therefore, the squash presetting method is worth popularizing in the engineering practice [3]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001637_iros.2004.1390027-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001637_iros.2004.1390027-Figure1-1.png", + "caption": "Fig. 1 Planar hyper-redundant modularized maoipulator", + "texts": [ + " It is based on the concept variable structure regular polygon and subsystem. Based on neural networks controller, the approach presented in this paper is completely capable of solving the control problem of a planar hyper-redundant manipulator with any number of links following any desired path. Simulation of a six-link modularized manipulator\u2019s inspection work in a honle-like concave has demonstrated that this control technique is available and effective. II. REPRESENTATION OF PLANAR HYPERREDUNDANT MODULARIZED MNIPLILATORUSING As shown in Fig.1, the hyper-redundant manipulator has 0-7803-8463-61041$20.00 @ZOO4 IEEE 3924 a lot of links, so it is often repeat and modularized. The manipulator is composed of a serial chain of links with the same length where the intemal variables 6, ( i = 1 - n) are the relative angles between adjacent links. The end effector's oosition is ziven bv where y, is the absolute link angle given by To control the hyper-redundant manipulator's end effector's position to realize the task as point to point, we have to calculate any joint's relative angle, that is, to solve the inverse problem basically" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003743_acc.2009.5159984-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003743_acc.2009.5159984-Figure1-1.png", + "caption": "Fig. 1. Multi-RTAC testbed.", + "texts": [ + " Furthermore, assume that the k-transversality condition (15) holds for the continuous-time dynamics of the closed-loop system (2)\u2013(6) with Xi(x) = d dt Vci(qci, q\u0307ci, yqi ), i = 1, . . . , s. Then the zero solution x(t) \u2261 0 to G is asymptotically stable. Finally, if Dq = R n and the total energy function V (x) is radially unbounded, then the zero solution x(t) \u2261 0 to G is globally asymptotically stable. Proof. The proof is omitted due to page limitation. In this section, we describe the multi-RTAC nonlinear system and design decentralized energy-based hybrid controllers to stabilize the zero equilibrium state. The multiRTAC system shown in Figure 1 consists of three identical translational oscillating carts connected by linear springs along with three identical eccentric rotational inertias which act as proof-mass actuators mounted on each cart. Rotational motion of each proof-mass is nonlinearly coupled with the translational motion of the corresponding cart that the proof-mass is mounted on which provides the mechanism for control. The oscillator carts, each with mass M , are connected to each other as well as fixed supports via linear springs of stiffness k", + " Then decentralized state-dependent hybrid subcontroller has the form mcq\u0308ci + kc(qci \u2212 \u03b8i) = 0, (qci, q\u0307ci, \u03b8i, \u03b8\u0307i) 6\u2208 Zi, (23) [ \u2206qci \u2206q\u0307ci ] = [ \u03b8i \u2212 qci \u2212q\u0307ci ] , (qci, q\u0307ci, \u03b8i, \u03b8\u0307i) \u2208 Zi, (24) ui = kc(qci \u2212 \u03b8i), (25) with the resetting set (13) taking the form Zi = { (qci, q\u0307ci, \u03b8i, \u03b8\u0307i) \u2208 R 4 : kc\u03b8\u0307i(qci \u2212 \u03b8i) = 0 and [ \u03b8i \u2212 qci \u2212q\u0307ci ] 6= 0 } . (26) It was shown in [14] that the closed-loop system (17)\u2013 (22) and (23)\u2013(26) satisfies k-transversality condition given in Definition 2.1, and hence, by Theorem 2.1, is globally asymptotically stable. In the next section, we implement the decentralized energy-based hybrid control framework on the multi-RTAC testbed and present the experimental results. The experimental testbed constructed to implement the decentralized energy-based hybrid control technique is shown in Figure 1. It consists of an aluminum base with two rails that air bushings float on providing translational motion for the carts with very low friction. Rotary actuators affixed with eccentric arms and masses are fixed to the carts providing the control torques. The actuation is provided by DC motors driven by a set of linear motor controllers, and the measurements of the eccentric arm angles and cart positions are performed with a quadrature encoder on each motor and a laser displacement sensor for each cart, respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002861_9780470289303-Figure2.10-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002861_9780470289303-Figure2.10-1.png", + "caption": "Figure 2.10 sensor nodes. Hierarchical data dissemination among three disjoint sets of", + "texts": [ + " Channel includes all nodes within the n-hop neighborhood of the node hosting the task instance neighborhood-hops:n neighborhooddistance:d the task instance Channel includes all nodes within a distance d of the node hosting k-nearest-nodes:k Channel includes the k nearest nodes of the node hosting the task instance The input (output) channel includes the the set of nodes that host the k nearest producers (consumers) of the data item associated with this channel k-nearest-pc:k all Channel includes all nodes in the system domain ~~ Channel includes all nodes that are owned by the task instance. This value is used in conjunction with the nodes-per-instance or area-perinstance values of the Instantiation annotation of the abstract task (see Fig. 2.10 for an example) parent Channel applies to the parent of the node hosting the task instance; in the virtual tree topology imposed on the network by the runtime system. Channel applies to all children of the node hosting the task instance; in the virtual tree topology imposed on the network by the runtime system. children domain. Section 2.5 illustrates the application of these annotations through a set of ATaG programming examples. In the following sections, we discuss in more detail the task and channel annotations listed in Tables 2", + " Annotate the input channel between data item D and task T2 as all-nodes . Annotate taskT2 with the any-data firingrule and one-on-node-ID : 8 placement annotation. Whenever the task TI is fired and produces data item D, it will be sent by the runtime to its nearest node in the network and then routed to the supervisor node. This example illustrates how sophisticated behaviors can be modeled using the basic set of annotations. Naturally, support for interpreting the annotations must exist in the compiler and in the runtime system. Figure 2.10 illustrates the effect of using the domain channel annotation in conjunction with the partitioning annotations for task placement. Note that the domain abstraction is valid only if the task associated with the channel has a placement annotationsnodes-per-instance : /k or area-per-instance : /k. As mentioned earlier, the partitioning of the set of nodes or the area of deployment is left to the compiler. The use of domain as the channel annotation in this case means that the scope of the dissemination (collection) of the output (input) data for an instance of the associated abstract task is defined by the partition that is \u2018assigned\u2019 to that task by the compiler" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001393_kem.291-292.483-Figure3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001393_kem.291-292.483-Figure3-1.png", + "caption": "Fig. 3 Coordinate system of meshing roller", + "texts": [ + " The transformation matrix of these two coordinate systems is given by M3 3\u2032 . The relationship between the planet worm-gear rotary coordinate system 2\u2032S and the meshing roller coordinate system ),,( kjiS is also illustrated in Fig.2 and the transformation matrix of these two Fig. 2 Meshing coordiante system coordinate systems is given by M . The tooth profile of the internal gear is generated by the compound movement of the meshing roller and the surface of the meshing roller is called generating surface. Fig. 3 illustrates the meshing roller in its coordinate system S . In meshing roller coordinate system, the generating surface can be given as, ( )T 1sinsincossincos vuvuu \u03c1\u03c1\u03c1=r . (1) Where, u and v represent the meshing parameters on the generating surface, \u03c1 represents the radius of the meshing roller. Via coordinate transformation matrix M , the generating surface can be illustrated in planet worm-gear rotary coordinate system 2\u2032S as ( )T 22 1sinsincossincosM vuvuur \u03c1\u03c1\u03c1+==\u2032 rr . (2) Where, 2r represent the radius of reference circle in planet worm gear" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003552_amr.97-101.3761-Figure4-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003552_amr.97-101.3761-Figure4-1.png", + "caption": "Fig. 4 Surface error of both side of face gear", + "texts": [ + " So, when a face gear with spur teeth is cut on this hobbing machine, the surface error will appear due to the immanent compensation rotate of the worktable. For this case, the bigger the inputting module is, the smaller the error is. An cutting experiment is done on YKS3140 CNC hobbing machine. The parameters of face gear are: the number of teeth is 77; the inner diameter is 224 mm; the outer diameter is 266 mm; the width is 21 mm; the module is 3 mm; the nominal press angle is 20\u00b0; the shaft angle is 90\u00b0. The CNC hobbing machine and the face gear are shown in Fig. 3.The surface errors of both sides of face gear are shown in Fig. 4. We can see that the max surface error of both sides is less than 100 \u00b5m. The tooth surfaces of face gear are symmetric on both sides, but the surface errors show the asymmetry. Many reasons result in the asymmetry, such as the manufacture error of hob cutter, installation error and so on. These questions are under improvement. The scheme of cutting face gears using a CNC hobbing machine is described in detail. According to the research in this paper, a funny phenomenon is found that the helix angle and the module need to be modified respectively when face gears with helix teeh are cut on two kinds of CNC hobbing machine" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003928_s10409-010-0364-1-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003928_s10409-010-0364-1-Figure2-1.png", + "caption": "Fig. 2 Conventions used in this paper: a helicity, b rotation, c torsion", + "texts": [ + "8\u00c5, is responsible for the helix formation of flagellar filaments. The details of this part can be found in Refs. [2,3]. The objective of this paper is to perform a mechanical analysis on Hotani\u2019s experiment which was conducted on a single bacteiral flagellar filament during the mechanical force-induced phase transition [16]. Before the analysis, we shall first give a brief description of the experiments. The conventions used in the analysis are the same as those used in the paper of Macnab [17], as summarized in Fig. 2. In Fig. 2a, helical sense is defined as left-handed (LH) or right-handed (RH) by the usual convention that the phase vector rotates CCW or CW, respectively, as the locus moves away from an observer looking along the helical axis. In Fig. 2b, sense of flagellar rotation is defined as CCW or CW from the viewpoint of an observer looking along the helical axis of the flagellum into its point of attachment to the cell. In Fig. 2c, the torque applied by the motor to the flagellum combines with viscous resistance to rotation to give a twisting moment, or torsion, which is LH if the motor is rotating CCW and RH if the motor is rotating CW. 2.1 Experiment of bacterial filament in vivo The experiment was to investigate the natural flagellar filament attached to an energized cell body [17]. During cell swimming, Macnab observed normal to curly form transition on both wide and mutant bacteria filament. Figure 3 shows the motion of a filament of radius \u03b7 and wavelength (pitch) \u03bb moving at a constant velocity through a viscous medium" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003847_cefc.2010.5481064-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003847_cefc.2010.5481064-Figure1-1.png", + "caption": "Fig. 1 8-pole magnetized PCMM", + "texts": [ + " The operation of a memory motor is based on its ability to change the magnetization of its magnets with a low amount of stator current. It is illustrated how the magnetization of rotor magnets can be continually varied by applying a short pulse of stator current [1], [2]. If the rotor of a memory motor is built following the same sandwich principle shown in [1], but with more than one magnet per pole one can group equally magnetized magnets in various manners. As a consequence, the number of rotor poles changes. This is the basic principle of operation of a pole-changing memory motor, as illustrated in Figs. 1 and 2. In Fig. 1, the cross-sectional view of a pole-changing memory motor with 32 tangentially magnetized magnets is shown. On the rotor side there are four magnets per pole, all of them being magnetized in the same direction. PM along with iron segments build the rotor wreath which is mechanically fixed to a nonmagnetic shaft. After the stator winding is reconnected into six-pole configuration, a short pulse of stator current changes the rotor eight-pole magnetization into a six-pole one, as shown in Fig. 2. Since the number of magnets per pole is not any more an integer (32/6=" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000270_iemdc.1999.769039-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000270_iemdc.1999.769039-Figure2-1.png", + "caption": "Fig. 2. Phasor diagram.", + "texts": [], + "surrounding_texts": [ + "I. INTRODUCTION\nPrediction of motor parameters when the motor is still in the stage of construction is one of the most important tasks of every motor designer. Generally, motor parameters are not constant but vary depending on load conditions and the level of saturation of flux paths in the motor [I].\nMagnetic conditions at different load conditions can be easily simulated by the 2D Finite Element Method.\nConsequently, the authors proposed an efficient procedure for calculation of the parameters of the two-axis model of a synchronous motor with permanent magnets by postprocessing the static magnetic field calculation results by 2D Finite Element Method (FEM).\n0-7803-5293-9/99 $10.00 0 1999 IEEE 98\n11. METHOD OF ANALYSIS The magnetic conditions in the motor were computed by\n(1) where v denotes the reluctivity, A is the magnetic vector potential, J, is the current density, and M is the magnetization.\nIn the case of sinusoidal balanced supply, the current of the first phase is given by\n2D FEM using the basic equation\nrot( v rot(A)) = J,, + rot Y M.\ni, = J?:-I-cos(ot+n. (2)\nwhere I is R M S value of stator current, p initial phase angle and o electrical angular velocity. In the remaining two phases the phase shift of currents is f120\" in the relation to the first phase.\nRotor starting position wt=O is chosen in such a way that the magnet axis of phase a is in alignment with the rotor direct axis. Then the initial phase angle p of the stator current is also the electrical angle between the stator MMF and rotor direct axis. For steady state conditions at constant load, constant voltage and frequency, electrical angle p is constant. Different operating modes can be simulated just by shifting the initial angle p. For each point of load the discrete time forms of phase voltages were calculated from the average values of the vector magnetic potential in stator slots for different discrete time moments by moving the rotor body and simultaneously changing the stator excitation over a half of electrical cycle. The end winding contribution was included in the modelling with constant value of end winding inductance Le. The instantaneous value of the phase voltage in the winding of phase a is given by\nwhere R is phase resistance,y/, is the instantaneous flux linkage of the a phase winding per pole, p is the number of pole pairs and c is the number of parallel circuit of the phase winding. The RMS values of phase voltages were calculated from their calculated time forms.\nThe load angle 6and the phase angle p were calculated from the calculated time form of the phase voltage and known time form of supply current. In the calculation of the phase winding flux linkage per pole the skewing of the stator slots is taken into account by calculation of average value of flux linkage of three slices of the stator package along z axis.", + "From the phasor diagram the components of the induced voltage in quadrature and direct axis Ek and Ea are defined by equations (4) and (5):\nEiq = Ei .COSSi= Eo +I\u2019cosp.xd\nEid = Ei -s insi= I - s inp -X, .\n(4)\n( 5 ) where Xd and Xq are direct and quadrature synchronous reactance, EO induced EMF due to the magnets, 6i internal load angle between the phasor of induced voltage Ei and quadrature axis and p torque angle between the phasor of stator current and direct axis.\nThe components of induced voltage Eh and Eid are determined as\n(6)\n(7)\nThe parameter X, can be obtained from (5), but & cannot be determined from (4) without the assumption of constant value of Eo. The uniqueness of the separate determination of EO and Xd depends on the superposition which cannot be applied under saturated conditions. Because of this condition the variation of the parameter Eo must be taken into account in determination of Xd.\nFor this reason an additional expression to expression (4) is needed. On the supposition of a small change of the stator current the magnetic conditions in the motor are the same and the parameters Xd and Eo remain constant.\nFrom measured or calculated data of load conditions the approximation cumes Lj(p), I v ) and E @ ) with use of orthogonal polinoms approximation were calculated.\nFor p\u2019=p +Ap ( A p =O. 1 degree) curves S;\u2019(p \u2019), I\u2019(p \u2019) and Ei\u2019(p\u2019) were made and considered in (8)\nEin = V .cos8 - I. R .cos(g, - 8)\nE,d = V - s ins+ I . R .sin((p - s).\nE; \u2018 C O S 8 := Eo + I\u2019 * C O S P \u2019 . xd. (8)\nParameter Xd and parameter Eo are then calculated from (4) and (8).\n014- 0\u20193 -. 0 12 -. -0 I 1 -. &IO-.\n3007 5006 %oos 3004 -om\ng: 1:\n002 001 000 i\n111. UNBALANCED STANDSTILL WORK TEST\nThe exact value of end winding inductance can be determined by combining unbalanced standstill work test and its simulation by FEM. The experimental arrangement of unbalanced standstill work test is presented in Fig. 3 [2].\n- n\n+tolal w e mjwtmx+ f r ~ m t leakage .Q- e d m l e a l c d g e - L e ( H ) s ~ t r l l ~ r a k test by XJ -cl+ Cnd e d m\n-. - - -. -. -. -. - -\nI LR - V b +\nFig. 3. W Unbalanced standstill work test arrangement. Fig. 4. Variation of end winding inductance L (H)." + ] + }, + { + "image_filename": "designv11_69_0001704_wcica.2006.1713035-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001704_wcica.2006.1713035-Figure2-1.png", + "caption": "Fig. 2. Model of 2 DOF planar space robot system.", + "texts": [ + " Collect the minimum the fitness function value. Step 10. Let k = k + 1, generate a new population popk using reproduction, crossover and mutation operators. Go to Step 4, until k = ng Step 11. Obtain the minimum fitness function value, thus, the parameters in this situation is optimal values. Get the optimum trajectory Q\u2217 ij . In order to verify the performance of the proposed optimal algorithm. Let\u2019s consider an example to better understand the optimum algorithms. A model of a planar 2 DOF freeflying space robot is shown in Fig. 2. The detailed derivation of dynamic equations of space robots can be obtained from [12],[13]. The parameters of the space robot are shown in Table I. For a real space robot system, the joint angle, angular velocity, acceleration, and torque of the manipulator should have constraint values. We can define the constraint conditions of the model of space robot as follows. \u2212pi \u2264 qj \u2264 pi, j = 1, ..., 2 vjmax = 2.5rad/s, j = 1, ..., 2 ajmax = 5rad/s, j = 1, ..., 2 \u03c41max = 25Nm, \u03c42max = 10Nm . In the simulation study, we plan a point to point trajectory in joint space" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001103_etfa.2005.1612520-Figure6-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001103_etfa.2005.1612520-Figure6-1.png", + "caption": "Figure 6. Field of Predictive potential", + "texts": [ + " If the robot is predicted to pass the crossing point earlier than the obstacle, the predictive potential is attractive, and vice versa. If the situation is opposite, the potential is repulsive. When the approaching times of the both objects are almost the same, the power of the potential should be strong to strength the acceleration or the deceleration. Besides, the closer the robot is to the crossing point, the stronger the predictive potential becomes. Under the assumption of uniform linear movement, the predicted arriving point of the obstacle is predicted when the robot reaches the crossing point (Fig.6). The notations in this figure are defined as follows. Xe: the distance between the obstacle's current position and the predicted arriving point xc: the distance between the obstacle's current position and the predicted crossing point Xr: the distance between the robot's current position and the predicted crossing point Xa the influential area of the prediction potential ke a constant value The prediction potential is calculated by Eq.(3) using the above notations. Xc Xa An example of the potential field is shown in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003209_pime_conf_1964_179_275_02-Figure24.2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003209_pime_conf_1964_179_275_02-Figure24.2-1.png", + "caption": "Fig. 24.2. Schematic diagram of apparatus showing the rolling elements and the loading cantilever", + "texts": [ + " Optical arrangement used in the measurement of creep of the ball degree of geometric conformity and the direction of motion. It is therefore conceivable that by a suitable choice of these parameters the slip due to a combination of these two conditions may be minimized and this should result in a minimizing of the ensuing wear. The following describes a series of experiments in which these possibilities have been investigated. APPARATUS The apparatus used in the creep and associated wear tests is shown in Figs 24.2 and 24.3. Fig. 24.2 indicates the general arrangement of the apparatus in which the balltype specimen was fixed on a spindle supported horizontally in ball-bearings. These bearings were located in a frame which was fixed to the \u2018free\u2019 end of a loading cantilever constructed from four spring-steel strips. The ball specimen rests on a track which is fixed to the driven shaft, a normal load being applied by a screw at the centre of the propped cantilever. The deflection of the cantilever is registered by a dial gauge (1/10 000 in/division) which had previously been calibrated in terms of the normal load at the ball-track contact" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002034_2007-01-2860-Figure16-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002034_2007-01-2860-Figure16-1.png", + "caption": "Figure 16 \u2013 Crossbar twist movement and moment applied by the trailing arm", + "texts": [ + " 2 \u2013 C-profile borders asymmetry: The absence of symmetry in relation to XY-axis can cause a flexibility difference in superior and inferior borders, generating different displacements in Y-axis, finally dragging the trailing arm and forming toe angles. Profiles with this asymmetry can present, for example, greater flexibility of the inferior border, producing great displacements in comparison to the superior border. In the asymmetrical movement, the arm applies a moment around the X-axis in the profile. The compression movement creates a binary, as in Figure 16. The twist movement causes the crosspiece to bend around point Cc, if the inertia of the inferior profile border relative to point Cc is very low when compared to the inertia of the superior border, as in Figure 17. This flexibility difference generates bigger Ydisplacement of the inferior border, forcing positive toe during compression. Therefore, it can be concluded that the greater the proportion between border inertias is, the greater will be the trend to produce toe during compression. 7 3 \u2013 Relation between C-profile border size: As illustrated in Figure 18, the size difference between the borders generates a moment that will direct the trailing arm to toe or divergence movement during compression" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000451_b978-0-12-555160-1.50017-4-Figure13-1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000451_b978-0-12-555160-1.50017-4-Figure13-1-1.png", + "caption": "Fig . 13-1 . Micro antimony pH-sensitive electrode, (a) Glass capillary substrate; (b) metal or glass rod substrate; (c) magnification of electrode tip. (Transidyne General Corporation, Ann Arbor, Michigan.)", + "texts": [ + "2 pH units), it suffers from a salt error, calibration is needed for each specific application, oxidizing and reducing agents will interfere, the electrode is poisoned by traces of metals such as copper, silver, and other metals below antimony in the electromotive series, and interference is found by certain complexing agents if present in the solution. The electrode is useful only in the range pH 1 to 10 because of its amphoteric properties. A recent development is the modification of the antimony electrode into a microelectrode. The tip configuration which involves a vacuum deposition of a very pure antimony film is shown in Fig. 13-1. Such a miniaturized electrode has been suggested to have improved properties over the conventional antimony electrode. Providing the limitations are recognized, the electrode in its micro form should be extremely useful for registering pH information in studies in microbiology and physiology. For example, in vivo measurement of pH level in blood is possible. It can be shown that for both the quinhydrone and antimony electrode the cell potential is given by Eq. (13-6). In practice, therefore, these two electrodes must also be calibrated by standard buffers", + " Calcium ion has been determined in beer, boiler water, soil, feedstuffs, flour, minerals, milk, sea water, serum and biological fluids, sugar, pulping liquor, and wine with the calcium electrode. Not all of these are direct measuring procedures. Some involve titration techniques. One last application is shown in Fig. 13-13 where the pH or cation activities of luminal fluid in kidney tubules are detected in situ. The tubules are of microscopic size and contains the preurine. In this application micro-tipped electrodes such as the one shown in Fig. 13-1 were used. Many of the ionselective electrodes are available as microelectrodes and are routinely used in a variety of applications. Figure 13-14 summarizes many applications of these microelectrodes. M E A S U R E M E N T S WITH A pH M E T E R One of the limitations of the glass, solid state, and liquid-liquid membrane electrodes is that they have very large internal resistances (megaohm range). For this reason, the simple potentiometric circuit (see Chapter 10) must be modified before it can be used with these electrodes" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001979_50037-5-Figure7.78-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001979_50037-5-Figure7.78-1.png", + "caption": "Figure 7.78 Contact Pair created by Contact Wizard.", + "texts": [ + " Contact Wizard is the facility offered by ANSYS. From ANSYS Main Menu select Preprocessor \u2192 Modelling \u2192 Create \u2192 Contact Pair. As a result of this selection, a frame shown in Figure 7.71 appears. Contact Wizard button is located in the upper left-hand corner of the frame. By clicking [A] on this button a Contact Wizard frame, as shown in Figure 7.72, is produced. Pressing [A] OK button brings back Contact Wizard frame (see Figure 7.76) where the [D] Create button should be pressed. Created contact pair is shown in Figure 7.78. Finally, Contact Wizard frame should be closed by pressing Finish button. Also, Contact Manager summary information frame should be closed. Before the solution process can be attempted, solution criteria have to be specified. As a first step in that process, symmetry constraints are applied on the half-symmetry model. From ANSYS Main Menu select Solution \u2192 Define Loads \u2192 Apply \u2192 Structural \u2192 Displacement \u2192 Symmetry BC \u2192 On Areas. The frame shown in Figure 7.79 appears. Three horizontal surfaces should be selected by picking them and then clicking [A] OK" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000396_2001-01-1445-Figure10-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000396_2001-01-1445-Figure10-1.png", + "caption": "Fig. 10 Resonant mode (at 22Hz)", + "texts": [ + ", one of the two resonant frequencies, the shock absorber generates a type of bending resonance in the lateral direction, the tire and wheel vibrate in the lateral direction in phases opposite to each other, and the lower arm moves laterally. Therefore, it is clear that the bending mode of the shock absorber is the cause of the lateral vibration of the cross member described in the previous section. Note that the mode at 148Hz is also a bending mode of the shock absorber in the lateral direction. From 15-30 Hz, which is related to the riding comfort (harshness) characteristics, the fore-and-aft resonance of an un-sprung mass exists as shown in Fig. 10. In this section, Sensitivity Analysis using Measured FRF data (SAMF) is applied to the front suspension system, and the stiffness combination of the front and rear side mounts to effectively reduce resonant peaks of FRFs is examined. As an example of a structural modification, the system shown in Fig. 11 is considered. In this example, the point a means the FRF (Inertance) evaluation point at the center of the cross member, and the point b is the excitation point at the wheel axle. The points c and d are stiffness modification points at which rubber mounts are installed" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001979_50037-5-Figure7.41-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001979_50037-5-Figure7.41-1.png", + "caption": "Figure 7.41 Contour plot of nodal solution (von Mises stress).", + "texts": [ + " The frame shown in Figure 7.39 is produced. The selection [A] Load step number = 1 is shown in Figure 7.39. By clicking [B] OK button the selection is implemented. From ANSYS Main Menu select General Postproc \u2192 Plot Results \u2192 Contour Plot \u2192 Nodal Solu. In the resulting frame, see Figure 7.40, the following selections are made: [A] Item to be contoured = Stress and [B] Item to be contoured = von Mises (SEQV). Pressing [C] OK button implements selections. Contour plot of von Mises stress (nodal solution) is shown in Figure 7.41. Figure 7.41 shows stress contour plot for the assembly of the pin in the hole. In order to observe contact pressure on the pin resulting from the interference fit, it is required to read results by time/frequency. From ANSYS Main Menu select General Postproc \u2192 Read Results \u2192 By Time/Freq. In the resulting frame, shown in Figure 7.42, the selection to be made is: [A] Value of time or freq. = 120. Pressing [B] OK implements the selection. From Utility Menu choose, Select \u2192 Entities. The frame shown in Figure 7" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001120_bfb0035241-Figure3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001120_bfb0035241-Figure3-1.png", + "caption": "Figure 3. The Transmission Model", + "texts": [ + " (4) The projection of re, guarantees that force-control torques only operate in valid force-control directions. The equation for multiple-arm motions is similar, and is described in [7]. 3 . T r a n s m i s s i o n D y n a m i c s To model the actuator/transmission system, we will use the equations developed by Pfeffer et al. in [5]. However, since our experimental results [7] indicate a significant amount of damping at the motor, we will retain the viscous damping term at the motor to account for its effect. With this change, the actuator/transmission model fbr each joint, shown in Figure 3, is described by the equations: ImiIm + flmqm + kt(qm - Nql ) = 7\"m -- T I (5) l lq i + fllOl + N k t ( ~ r q l - - qru) .... 0 \u2022 - N k t ( N q ~ - q~,~) = T~ In this model, I ~ is the inert ia of the motor, and I~ is the inert ia of link i, all ou tboard links, and the load, about the joint 's axis of rotat ion. Current to the motor causes a torque of ~-,~ which is t ransmi t ted through a shaft of stiffness kt and a gear reduction with gear ratio N to the link. The final torque applied at the link has a value of 7z" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003301_amc.2010.5464114-Figure9-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003301_amc.2010.5464114-Figure9-1.png", + "caption": "Fig. 9. The modeling of the systems. (a) Master system. (b) Slave system", + "texts": [ + " In here, it is assumed that the rings are contact with the center of the middle phalanx and the distal phalanx. The master system represents the exoskeleton system. In the slave system, the end effector of x-y table reproduces the trajectory and the applied force in the finger tip point. This system represents the endoskeleton system. The information obtained by each actuator is transformed into finger modal space using the equations written in following section. The modeling of the master system is shown in Fig. 9 (a). In the master system, only the position and force information about y axis is obtained. Hence the information about x axis is unknown. The position of the center point in link 2 is obtained by [ yM1 yM2 ] = \u23a1 \u23a2\u23a3 1 2 l1 cos(\u03b8M1 + \u03b8fix) l1 cos(\u03b8M1 + \u03b8fix) + 1 2 l2 cos(\u03b8M1 + \u03b8M2 + \u03b8fix) \u23a4 \u23a5\u23a6 . (10) The Joint angle \u03b8M1, \u03b8M2 are obtained by \u03b8M1 = cos\u22121 2yM1 l1 \u2212 \u03b8fix (11) \u03b8M2 = cos\u22121 2(yM2 \u2212 l1 cos(\u03b8M1 + \u03b8fix)) l2 \u2212\u03b8M1\u2212\u03b8fix. (12) The velocity of the contact points are shown by[ y\u0307M1 y\u0307M2 ] = [ J11 M J12 M J21 M J22 M ] [ \u03b8\u0307M1 \u03b8\u0307M2 ] = Jaco M [ \u03b8\u0307M1 \u03b8\u0307M2 ] (13)\u23a7\u23aa\u23aa\u23a8 \u23aa\u23aa\u23a9 J11 M = \u2212 1 2 l1 sin(\u03b8M1 + \u03b8fix) J12 M = 0 J21 M = \u2212l1 sin(\u03b8M1 + \u03b8fix) \u2212 1 2 l2 sin(\u03b8M1 + \u03b8M2 + \u03b8fix) J22 M = \u2212 1 2 l2 sin(\u03b8M1 + \u03b8M2 + \u03b8fix) (14) where Jaco M stands for Jacobian matrix of the master system", + " The joint torque TM1, TM2 are obtained by[ TM1 TM2 ] = JT aco M [ FyM1 FyM2 ] . (15) The acceleration reference of the actuator y\u0308ref M1, y\u0308ref M2 are obtained by [ y\u0308ref M1 y\u0308ref M2 ] = Jaco M [ \u03b8\u0308ref M1 \u03b8\u0308ref M2 ] (16) The information in the actuator space is transformed into the finger modal space by equations (11), (12), and (15). Then the angular acceleration reference which is obtained by bilateral control is transformed into the actuator space by equation (16). The modeling of the slave system is shown in Fig. 9 (b). Tip position is obtained by[ xS yS ] = [ l1 sin(\u03b8S1 + \u03b8fix) + l2 sin(\u03b8S1 + \u03b8S2 + \u03b8fix) l1 cos(\u03b8S1 + \u03b8fix) + l2 cos(\u03b8S1 + \u03b8S2 + \u03b8fix) ] . (17) From equation (17), joint angle \u03b8S1, \u03b8S2 are obtained by \u03b8S1 = tan\u22121 \u2223\u2223\u2223\u2223xS yS \u2223\u2223\u2223\u2223 \u2212 tan\u22121 \u2223\u2223\u2223\u2223 l2 sin \u03b8S2 l1 + l2 cos \u03b8S2 \u2223\u2223\u2223\u2223 \u2212 \u03b8fix (when x > 0, and y > 0) \u03b8S1 = \u03c0 \u2212 tan\u22121 \u2223\u2223\u2223\u2223xS yS \u2223\u2223\u2223\u2223 \u2212 tan\u22121 \u2223\u2223\u2223\u2223 l2 sin \u03b8S2 l1 + l2 cos \u03b8S2 \u2223\u2223\u2223\u2223 \u2212 \u03b8fix (when x > 0, and y < 0) \u23ab\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23ac \u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23ad (18) The velocity of finger tip is shown by[ x\u0307S y\u0307S ] = [ J11 S J12 S J21 S J22 S ] [ \u03b8\u0307S1 \u03b8\u0307S2 ] = Jaco S [ \u03b8\u0307S1 \u03b8\u0307S2 ] (20) \u23a7\u23aa\u23aa\u23a8 \u23aa\u23aa\u23a9 J11 S = l1 cos(\u03b8S1 + \u03b8fix) + l2 cos(\u03b8S1 + \u03b8S2 + \u03b8fix) J12 S = l2 cos(\u03b8S1 + \u03b8S2 + \u03b8fix) J21 S = \u2212l1 sin(\u03b8S1 + \u03b8fix) \u2212 l2 sin(\u03b8S1 + \u03b8S2 + \u03b8fix) J22 S = \u2212l2 sin(\u03b8S1 + \u03b8S2 + \u03b8fix) (21) where Jaco S stands for Jacobian matrix of the slave system", + " In this experiment, the position information was measured by the position encoder, and force response was estimated by RFOB without using force sensors. Control program was implemented in RTAI 3.7. The parameters used in these experiments is shown in Table I, and the parameters about finger model is set as Table II. The experimental system is constructed by the master and slave systems, and these are shown in Fig. 10. In the master system shown in (a), two linear actuator systems were used, and the rings were connected to the actuators by the universal joints. The operator puts on the rings as shown Fig. 9 (a). Both linear motors are actuated only y-axial direction. The experimental system used as the slave system is shown in (b). This is an x-y table composed by two linear actuator systems. The end effector of x-y table tracks the finger tip position. The information acquired by the actuators used in the master and slave systems was transformed into finger modal space. The bilateral control was implemented in modal space. In this experiment, the human operator\u2019s motion included both the constrained motion, that the slave system is contact with an environment, and unconstrained motion, that slave system is not contact with an environment" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000169_iros.1999.812986-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000169_iros.1999.812986-Figure2-1.png", + "caption": "Figure 2 - Turning Maneuver Kinematics of Vehicle", + "texts": [ + " \u2018The common notations used throughout this paper are as listed below: , k T Xglobal , Yglobal %eh 9 Yveh Discrete-time index Sampling time of Kalman filter x- and y-axes of global reference coordinate frame x- and y-axes of vehicle local reference coordinate frame 2-D position,of vehicle with respect to global reference frame Average velocity of vehicle Front-centre velocity Actual front-centre velocity Rear-centre velocity Actual rear-centre velocity Heading or orientation of vehicle with respect to global reference frame Heading rate or angular velocity of vehicle Front\u2019and rear side-slip angles Velocities of respective wheels Steering angles of front respective wheels Average steering angle 3 Sensor Fusion Algorithm A kinematic representation of the vehicle is shown in figure 2. The model for the vehicle is simplified with the \u201cbicycle model\u201d [2]. Two side-slip angles, a and P , are used to reflect the actual directions that the front and rear virtual wheels are translating respectively [4]. 3.1 Formulation of Process Model A piecewise constant white acceleration model is used to describe the motion of the vehicle [ 11. In addition, the side-slip angles are incorporated into the process model to be estimated on-line by the filter. An assumption is made that the input to the system, 4 , is approximately constant over the sample time interval T " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003176_cca.2010.5611314-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003176_cca.2010.5611314-Figure1-1.png", + "caption": "Fig. 1: Schematic overview of the table.", + "texts": [ + " This mask is also removed from the images, such that the ball remains the only visible object. In this paper first the hardware design is discussed. In Section III the Unified Modeling Language is adopted to create a structural software architecture, after which in Section IV the practical implementations are discussed. Finally the results are presented in Section V. A professional football table is acquired on which an overhead camera is mounted. The automated rod can translate or perform a kicking movement, see Fig. 1. The Prosilica GC640c ethernet camera sends images to a HP xw4600 workstation where the images are processed. The workstation also communicates with the Beckhoff data acquisition equipment that allows to control the motors. The initial strategy is to control a single rod to intercept the ball and kick it back towards the human goal. To successfully execute this strategy, an activity diagram [5] is developed, see Fig. 2. In the activity diagram several components are directly related to motion control of the rod" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002034_2007-01-2860-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002034_2007-01-2860-Figure2-1.png", + "caption": "Figure 2: Crossbar profile Twist-beam Type", + "texts": [ + " At the same time, the crossbar has a critical structural aspect. Since it is subject to high rotation, which leads to great strain, fatigue dimensioning becomes a complex stage to define geometry, mainly in the connection with the trailing arm, where variation in profile section generally occurs. Structural reinforcements are added to this region in many cases because of high concentration of tension. Many crosssection solutions have been used along time, U and C profiles are commonly used now, but great variations from vehicle to vehicle still exist - Figure 2 illustrates a crossbar profile used in the market. Twist-beam suspensions offer the advantages of using parts simple to manufacture and a limited number of components, which have a low cost of production. Good dynamic performance is also noticed, compared to independent-type suspensions, although there is some dependence between the wheels. Another advantage is the small transversal profile, allowing great versatility in suspension models and body construction. A perceptible disadvantage of this solution is the impossibility of rear transmission" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001646_icma.2006.257749-Figure5-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001646_icma.2006.257749-Figure5-1.png", + "caption": "Fig. 5 Location of the strain gauge on the dynamometric ring in its standing position.", + "texts": [ + " The electronic operation is to translate the ring deformation (in millimeters) into the force exerted on the ring (in Newtons). Each sensor use an extensiometric bridge. This bridge is a high sensibility Wheaston bridge in which one strain gauge is used instead of one resistance element. The strain gauge is said to be set in a quarter of bridge. The strain gauge (resistance of 350 Ohms) is pasted on the external face and placed so that it is at the equator of the ring when the ring is standing on its plane (Fig. 5). The entry voltage in the bridge (Vin) is known and the exit one (Vout) is measured. The variation of this tension is due to the variation of the strain gauge resistance which translates its stretching (l) as written in (1). )1024/(10/ 63 \u2212\u2212 \u00d7+\u00d7= KlKlVV outin (1) with K: strain gauge coefficient. On the other hand, the gauge deformation during the ring compression may be known by the ring parameters (2) as written by [11] taken from [12]. 2/3)/21( EatFRl \u00d7\u03a0\u2212= (2) with \u2022 a: ring width \u2022 E: ring Young modulus (elasticity) \u2022 F: ring submitted force \u2022 R: ring external radius \u2022 t: ring thickness Thus, knowing all these parametres allow us to obtain the force (F) exerted on the ring by (3): )/21)(2(3/4 2 \u03a0\u2212\u2212= outinout VVRKEatVF (3) A dual PCMCIA card expansion jacket allows the connection of two PCMCIA cards on the PDA iPAQ HP 5500" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002628_cdc.2007.4434082-Figure4-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002628_cdc.2007.4434082-Figure4-1.png", + "caption": "Fig. 4. A general differential drive UGV. The darker rectangles are the tracks or wheels. z = (z1, z2) is the center of the robot while x = (x1, x2) is the camera position.", + "texts": [ + " Finally, Section V discuss future investigations and experiments using this approach and conclusions are drawn in Section VI. In this section we will first review two ordinary differential equation (ODE) models capturing the kinematic and dynamic behavior of the robots depicted in Figures 1, 2 and 3 above. We then describe what a standard teleoperation interface looks like, and contrast this to the computer game interfaces used in FPS-games, [8]. We present kinematic and dynamic models for both cases. Consider the general UGV model in Figure 4. If we identify the two tracks of the Packbot in Figure 1, with the main wheels of the Scout in Figure 2, and the two pairs of wheels on each side of the ATRV in Figure 3, we find that the following model is applicable to all three robots. z\u03071 = v1 + v2 2 cos \u03b8 (1) z\u03072 = v1 + v2 2 sin \u03b8 \u03b8\u0307 = v1 \u2212 v2 d \u03c6\u0307 = k, where z = (z1, z2) and \u03b8 are the position and orientation of the vehicle, \u03c6 is the orientation if the camera relative to the vehicle, v1, v2 are velocities of the wheels/tracks and d is the width of the vehicle" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002892_ichr.2008.4755933-Figure3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002892_ichr.2008.4755933-Figure3-1.png", + "caption": "Fig. 3. Desired postures and switching condition.", + "texts": [ + " To realize this interaction, we determined the above parameters according to following procedure: 1) configure the initial posture 2) configure the final posture and the feedback gain that can be assumed to maintain a standing posture 3) configure the intermediate postures that can be assumed to lead to the final posture from the initial posture 4) configure the timing for switching of each desired posture that can make the robot stand while receiving physical help. The switching times were defined as conditional expressions using previously realized posture vectors and the currently desired posture vector. This indicates that the robot motion is adaptively generated by the interaction because the robot's posture keep to be influenced from the helper in PHRI. Figure 3 shows the frame format of the desired posture vectors and the switching times that realize the rising-up interaction. As indicated by the figure, we assume that the robot can realize the rising-up interaction by switching the desired posture vector only two times. However, it needs a trial-and-error process using the robot to determine the appropriate switching times because simulating this kind of PHRI is nearly impossible. Figure 4 shows sequential photographs of an accomplished rising-up interaction using the parameters shown in Fig. 3. De spite the fact that there are drastic changes in the desired vector between postures, the realized motion still looks smooth. It is nearly impossible to realize a similar interaction using traditional robots and control systems. Therefore, the proposed system seems better suited for such physical interaction. v. EXPERIMENTAL SETUP An evaluation of physical interactions must be made before we can investigate a learning system operating through such physical interactions. In order to clarify the difference between successful and unsuccesful interaction, we conducted the ex periment explained in this section" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002988_978-0-387-22459-6_3-Figure3.12-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002988_978-0-387-22459-6_3-Figure3.12-1.png", + "caption": "Figure 3.12 (A) Six-state and (B) five-state simplification of the eight-state diagram for the Na+/glucose cotransporter.", + "texts": [ + "3: The Na+/Glucose Cotransporter 69 A D B C E F 1 2 3 4 5 6 7 8 1 2 3 4 5 6 7 8 1 2 3 4 5 6 7 8 1 2 3 4 5 6 7 8 1 2 3 4 5 6 7 8 1 2 3 4 5 6 7 8 Figure 3.11 Six possible diagrams for the Na+/glucose transporter with transport steps included. Only (A), (E), and (F) are compatible with experiment. mechanisms diagram (A), which is fully connected, and diagrams (E) and (F), each of which is missing Na+ transport steps. All three of these diagrams are compatible with the experimental evidence, and all three can be \u201creduced\u201d to a diagram with the 6-state skeleton givenin Figure 3.12A. This method of reducing diagrams uses the rapid equilibrium approximation that applies to steps for which the forward and reverse rates are rapid with respect to other steps in the diagram. The details of how this method works are explained in Chapter 4, although the basic idea can be seen by comparing Figure 3.12A and Figure 3.12B. The experimental values of rate constants for the six\u2013state model have been assigned by Parent and colleagues (Parent et al. 1992b). Step S4 to S5 in the six\u2013state model is the dissociation of glucose inside the cell, and this step is extremely fast. This permits the two states to be approximated as a single combined state (state S4,5 in Figure 3.12B and reduces the diagram to five states as shown. We must be careful in doing so to readjust the rates to account for the reduction. The details of the process for doing this is given in Chapter 4, but it is not difficult. In short, only a portion of the combined state S4,5 reacts to the other states. The portion to be used in each reaction is determined as a result of the reduction using simple algebra. It is possible to write diagrammatic expressions for the transport rate for either the five-state or six-state models in Figure 3.12. However, the number of directed diagrams and cyclic diagrams increases quickly with the complexity of cycles in the complete diagram. For example, for the five-state model there are 6 pairs of cyclic diagrams and 55 directed diagrams. Nonetheless, the general expressions in (3.19) and (3.20) remain valid and can be used to obtain explicit expressions for the steady\u2013state fluxes. The diagrammatic method does not, however, provide information about the transient timedependence of the fluxes. This is most conveniently obtained by numerical integration of the equations", + " Use the rate constants k12 k43 2.4 mM\u22121min\u22121, k21 k34 42 min\u22121, k14 k14 k41 k23 k32 1000 min\u22121. The transport rate is given by R c \u00b7 Jss 34, where c 2 mM is the concentration of GLUT transporters per unit volume of cells. [Hints: Because of the size of the rate constant k14, etc. you will need to use a small step size (try 0.0001min). For the same reason you only will need to integrate for about 0.3min.] 10. Write down the 6 pairs of cyclic diagrams and the 11 directed diagrams for state 1 for the five-state diagram in Figure 3.12B. What are the directed diagrams for state 4,5? 11. The cardiac form of the Na+/Ca2+ exchanger is electrogenic with a stoichiometry of 3Na+:1Ca2+. Assuming that the 3 Na+ bind sequentially to sites of decreasing affinity, how do you anticipate that the transport rate will depend on [Na+]out? Prove your answer." + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001504_s0005117906090049-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001504_s0005117906090049-Figure1-1.png", + "caption": "Fig. 1. General view of the wheel system.", + "texts": [ + " The following section describes the WS dynamics. Section 3 describes the problem of WS control and formulates the task of the present paper. The control law stabilizing the WS motion without regard for the state measurement error is constructed in Section 4. Sections 5 through 7 are devoted to the analysis of the impact of errors. The results of modeling and some generalizations are described, respectively, in Sections 8 and 9. 2. WHEEL SYSTEM AND THE CONTROLLED PLANT The general diagram of the wheel system under study is depicted in Fig. 1. The WS has body, back drive axle, and controlled front axle. The state of the body is characterized by the angle a and the coordinates x and y of some point p, and v is the absolute magnitude of its speed. The state of the front axle is characterized by the controlled angle b. With regard for the notation introduced, the WS motion obeys the following equation system: x\u0307 = v cos a, y\u0307 = v sin a, a\u0307 = v tan(b)/L, b\u0307 = F (b, u, t) . (2.1) The three first equations of system (2.1) describe the linear and angular motion of the WS. The last equation describes the motion of the drive of the controlled front axle, u is control, and v, L = const > 0. The relation p \u2208 S defines the starting aim of WS control where the system point p lies on the given curve S. The two first equations of (2.1) describe the mechanical relations of the WS shown in Fig. 1. They reflect the assumption that the back wheels do not skid in the direction of wheel axles. A similar assumption about the front wheels allows one to construct the third equation of the system. These equations represent a kinematic model of the mechanical wheel system. Models of this kind are studied actively [1]. For example, some publications take into consideration the inertial characteristics of the mechanical wheel system, which enables one to study, for example, the impact of the external forces on the WS [8, 11]", + "2) \u2200b, \u2200t \u2265 0, |u| \u2264 h, |\u2202F/\u2202u| \u2264 L, H,h,L = const > 0. Therefore, the angular speed b\u0307 of the controlled WS front axle is bounded, and the control u actually can change only the sign of speed b\u0307. In one or another form, (2.2) is in essence the necessary property of any physical control device. We note that the WS drive dynamics can be taken into account in the general form where u additionally satisfies equation, for example, of the form u\u0307 = \u03c8 (b, u, U, t), where U is a new control (Section 9). The smooth curves on the plane X,Y (Fig. 1) \u2223 \u2223A\u2032\u2223\u2223 \u2264 A \u2032, \u2223 \u2223A\u2032\u2032\u2223\u2223 \u2264 A \u2032\u2032, A \u2032, A \u2032\u2032 = const \u2265 0, A\u2032 = dA(s)/ds (2.3) will be considered as the given trajectory S of WS motion. Here, A(s) denotes the angle of the tangent to S at the point s. The curve S is assumed to be given parametrically S = \u3008(x, y) : x = \u03a6x(s), y = \u03a6y(s)\u3009, (\u03a6\u2032 x) 2 + ( \u03a6\u2032 y )2 = 0, where x and y are the coordinates of a point on the curve S in the system {X,Y } corresponding to the parameter s which has the sense of the length of the corresponding arc of the curve S" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001823_1.3547940-Figure13-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001823_1.3547940-Figure13-1.png", + "caption": "FIG. 13. \u2014 Proposed asymmetric test specimen.", + "texts": [ + " Whether such predictions are correct can only be answered by experiments. But our efforts here have led to the suggestion of a direct test of the various failure theories. It is noteworthy that the stress distributions at the bond become relatively insensitive to specimen height when h/a > 2 because we can begin to visualize the specimen as a 3-part cylinder. If this is so, then there is no need to keep tall specimen symmetrical. An asymmetrical specimen with a conical substrate at one end and a spherical substrate at the other (Figure 13) has the advantage that the reaction force at each end must be identical. So, in a single test, we have a direct comparison of two specimen geometries. The cone angle and the spherical radius can be varied independently. Making the specimen out of a transparent rubber will permit the observation of cracks that initiate in the interior. Indeed, if we characterize our materials and have crack growth characteristics in the form (15) it is possible to make quantitative predictions of n, the number of cycles to failure from (16) where c0 and c1 are the initial and final crack lengths, respectively", + " The flat substrate is the limiting case of the spherical substrate when the radius of curvature is infinitely large. Figure 12 shows that the tearing energy for an external crack at a spherical substrate with r = 3.304a is lower than that for a flat substrate. Thus, reducing the radius of curvature has the effect of suppressing the tearing energy for crack growth at the perimeter. This is shown in Figure 17. Therefore, reducing the radius of curvature at the spherical end of the asymmetric test specimen in Figure 13 is a useful technique to ensure that failure at the other (coni- cal) end is favored. This is a potentially useful experimental arrangement. In a fatigue machine, typically one end is held fixed while the other end is attached to a moving member. It is quite difficult to observe small cracks at the moving end without stopping the machine. It might be convenient to arrange for a spherical substrate with a small radius of curvature at the moving end so that failure at that end is unlikely. Although the specimens in ASTM D 429 look quite simple at first sight, they have proven to be well worth revisiting" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001979_50037-5-Figure7.50-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001979_50037-5-Figure7.50-1.png", + "caption": "Figure 7.50 A half-symmetry model.", + "texts": [ + " The objective of the analysis is to observe the stresses in the cylinder when the initial gap between two blocks is decreased by 0.05 cm. The dimensions of the model are as follows: cylinder radius = 0.5 cm; cylinder length = 1 cm; block length = 2 cm; block width = 1 cm; and block thickness = 0.75 cm. Both blocks are geometrically identical. All elements are made of steel with Young\u2019s modulus = 2.1 \u00d7 109 N/m2, Poisson\u2019s ratio = 0.3 and are assumed elastic. Friction coefficient at the interface between cylinder and the block is 0.2. For the intended analysis a half-symmetry model is appropriate. It is shown in Figure 7.50. In order to create a model shown in Figure 7.50, the use of two 3D primitives, namely block and cylinder, is made. The model is constructed using GUI facilities only. When carrying out Boolean operations on volumes it is quite convenient to have them numbered. This is done by selecting from the Utility Menu \u2192 PlotCtrls \u2192 Numbering and checking appropriate box to activate VOLU (volume numbers) option. From ANSYS Main Menu select, Preprocessor \u2192 Modelling \u2192 Create \u2192 Volumes \u2192 Block \u2192 By Dimensions. In response, a frame shown in Figure 7.51 appears", + "58 shows that the cylinder was moved by [A] 0.05 cm downward, i.e., toward the block, after clicking [B] OK. From Utility Menu select Plot \u2192 Replot to view the cylinder positioned in required location. Finally, from Utility Menu select PlotCtrls \u2192 View Settings \u2192 Viewing Direction. The frame shown in Figure 7.59 appears. By selecting [A] X, Y, and Z, coordinates as shown in Figure 7.59, clicking [B] OK button, and activating Plot \u2192 Replot command (Utility Menu), a half-symmetry model, shown in Figure 7.50, is finally created. Before any analysis is attempted, it is necessary to define properties of the material to be used. From ANSYS Main Menu select Preferences. The frame in Figure 7.60 is produced. From the Preferences list [A] Structural option was selected as shown in Figure 7.60. From ANSYS Main Menu select Preprocessor \u2192 Material Props \u2192 Material Models. Double click Structural \u2192 Linear \u2192 Elastic \u2192 Isotropic. The frame shown in Figure 7.61 appears. Enter [A] EX = 2.1 \u00d7 109 for Young\u2019s modulus and [B] PRXY = 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002098_piee.1967.0321-Figure18-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002098_piee.1967.0321-Figure18-1.png", + "caption": "Fig. 18", + "texts": [ + " 33 and 36 combine to give Px*rx + Py*ry = \u00b0 (37) The significance of this result in relation to synchronous crawling is made clear by considering the load angle of a doubly fed cascade system formed by a pair of harmonics of 2px and 2py poles. These harmonics produced by the stator are assumed to rotate in the same direction, so that the preceding analysis is directly applicable to the 2px- and 2/ypole fields which are together induced in the rotor by both stator harmonics, as shown, for example, by the fundamental and 13th harmonics in Table 2. With reference to Fig. 18, the load angle of this doubly fed cascade system is hy = Using the relation established in eqn. 37, this expression for the load angle reduces to = (Px Px\"s (39) It is seen from this equation that, owing to the relationship that exists between the spatial positions of the rotor harmonics, the origin of the electrical load angle is independent of these positions. This same result may be derived for the other pairs of locking harmonics shown in Table 2. It is of interest to note that, owing to the absence of the displacements arx and ary in eqn" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002287_s11548-008-0159-z-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002287_s11548-008-0159-z-Figure2-1.png", + "caption": "Fig. 2 Different regions excluded from the robots workspace: 1 a box surrounding the patient, 2 a box representing the patient couch, 3 the robot base, and 4 regions where the beams path through the CT is not fully covered", + "texts": [ + ", such that the robot can move from point to point without colliding with the patient or any other obstacle. As robotic radiosurgery is an image guided procedure, the set of admissible positions may be further restricted to avoid occlusion of the imaging subsystem. In our simulation environment, an initial set of beam source points can be established from a given treatment plan. Subsequently, obstacles can be defined and beam source points can be generated automatically, or manually added or removed. Figure 2 illustrates how the workspace is restricted. Simple shapes surrounding the patient, the patient couch, and the robot base are used for basic collision detection. Furthermore, only beams which enter the patient through the skin surface within the CT data-set can be considered in order to allow accurate dose calculation. The environment is linked to an in-house developed treatment planing system. In the present study we adopt the two-phase planning approach proposed by Schweikard [6]. The first phase consists of generating a large set of candidate beams by a randomized selection of beam source point and beam orientation" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003410_2013.26820-Figure6-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003410_2013.26820-Figure6-1.png", + "caption": "Figure 6. Mesh stiffness of gear.", + "texts": [ + " Equations 11 and 12 give the torques acting on the two gears under the drive and free conditions: ][ ][ 211 2 2 1 2 ,2 211 2 2 1 1 ,1 ggg gg g driveg ggg gg g driveg DNDT JNJ JN T DNDT JNJ J T \u22c5\u2212\u2212 \u22c5+ \u22c5 = \u22c5\u2212\u2212 \u22c5+ = (11) 2,2 11,1 gfreeg ggfreeg DT DTT \u2212= \u2212= (12) The backlash was characterized by two parameters, back\u2010 lash (2 ) and mesh stiffness (kg), as shown in figure 5 (Wang et al., 2001). The mesh stiffness is defined as a non\u2010linear spring constant along the line of action of two mating gears, as shown in figure 6. Let xGP be the relative space between two mating gears of pitch circle radii of Rg1 and Rg2. The tooth deflection by the impact force between the two gears may be expressed as: \u239f \u23a9 \u239f \u23a8 \u23a7 \u03b5\u2212<\u0394\u03b5+\u0394 \u03b5\u2264\u0394\u2264\u03b5\u2212 \u0394<\u03b5\u03b5\u2212\u0394 =\u03b5\u0394 GPGP GP GPGP GP xifx xif xifx xg , ,0 , ),( (13) The mesh stiffness is then given by: ),( \u03b5\u0394 = GP GP g xg F k (14) The impact force was calculated from equation 2 by divid\u2010 ing the torque transmitted from the engine by the pitch circle radius of the driving gear. When a gear drives its mating gear, the mesh stiffness changes continuously because the relative space between the two gears varies constantly within the backlash" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003892_appeec.2010.5448977-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003892_appeec.2010.5448977-Figure2-1.png", + "caption": "Figure 2. Power supply for the rover unit", + "texts": [ + " The positioning information from the DGPS receivers is sent to a computer through wireless devices. The computer saves the positioning information in a hard disk for postprocessing software to draw the tracks of transmission conductor wave and wind galloping. For a practical conductor wave and wind galloping measurement instrument, there is the matter of instrument power supply. The power requirements at the base station are derived from conventional sources. The rover is mounted on the transmission conductor and power must be derived from the conductor itself [14]. Fig.2 shows a configuration based on a current transformer (CT) design. Fig.3 shows the signal communication between the base and the rover. The base sends the differential GPS signals to the rover through a wireless network device. The real-time DGPS positioning signals that obtained by the rover is send to the IBM server through another wireless network device. IV. MEASUREMENT SYSTEM SOFTWARE DESIGN The measurement software includes four major parts: 1. The DGPS signals receiving and displaying area which shows the information obtained from the rovers; 2" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003238_20100802-3-za-2014.00006-Figure5-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003238_20100802-3-za-2014.00006-Figure5-1.png", + "caption": "Fig. 5. Simulation result of NDOB", + "texts": [ + " Then, lower roll become upper roll and upper roll become lower roll. The initial values are measured value at this moment. To demonstrate that NDOB can recover the time trajectory of the nominal closed-loop system, we simulate with \u03c4 = 0.0001, \u03c4 = 0.004, and \u03c4 = 0.04. In MATLAB simulation we added two disturbances. They are d = 0.4 sin(2t) and Rlow = R + (\u03c9op/2/\u03c0 \u2217 h)t. Rlow is radius of low roll and h is strip depth. As Fig.4, The radius of lower roll increase as low roll move. A solid line is result of nominal response(Fig.5). As \u03c4 is smaller, NDOB is faster to recover the time trajectory of the nominal response. We simulated when \u03c4 = 0.0001, but in this case simulation result covers nominal result. Therefore, we remove the result when \u03c4 = 0.0001. In the hardware simulator, there exist time delay. As we mentioned in previous chapter, it is 28\u223c34ms. Therefore, we compared simulation result and experiment result, when \u03c4 = 0.04. The initial values of looper system are \u2206\u03980 = ( \u22120.0798 0.4438 ) , \u2206\u03c30 = 2.7987 \u00d7 105. These values are measured by hardware experiment" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001240_gt2006-90435-Figure3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001240_gt2006-90435-Figure3-1.png", + "caption": "Fig. 3 A gear pair system", + "texts": [ + "org/about-asme/terms-of-use The kinetic energy of a disk for lateral motion given by Shiau and Hwang (1993) is modified to include the torsional kinetic energy. The kinetic energy of a disk is described as ( ) ( )[ ] ( ) ( )[ ] ( )( ) ( )2 2222d 2 1 2 1 2 1 2 1 ddP ddddddPdddDddd I IIWVmT \u03b1 \u03b1 & &&&&&&& +\u03a9+ \u0392\u0393\u2212\u0393\u0392+\u03a9+\u0393+\u0392++= (1) where V and W are the two lateral displacements along the Y and Z directions and \u0392 , \u0393 are the corresponding bending angles in the Y-Z plane, respectively. dm , dDI and dPI are the mass, transverse mass moment of inertia and polar mass moment of inertia of the gear, respectively. \u03a9 is the spin speed of the shaft. Fig. 3 shows that a gear pair is modeled as the equivalent stiffness hk together with transmission error )(tet along the pressure line between the teeth. In the present model, the mesh stiffness is considered to be constant. Thus, the gear mesh force along the pressure line can be expressed as ( )])(cos)(sin)[( 22111212 terrWWVVkF tddpddpddhh \u2212\u2212\u2212\u2212+\u2212= \u03b1\u03b1\u03c6\u03c6 (2) where p\u03c6 is the pressure angle of the gear pair. 1r and 2r are radii of base circles of the driving gear and the driven gear, respectively. The mesh stiffness approximately by using a constant value and the term ( )tet is the displacement transmission error and is assumed to be time-varying and may include higher harmonics of tooth passing frequency as described by Ozguven and Houser (1988a), i" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001293_rob.20099-Figure12-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001293_rob.20099-Figure12-1.png", + "caption": "Figure 12. Implementation of step 1 of corrective phase: a minimum offset is determined and b vehicle is translated by this offset, respectively.", + "texts": [ + " The main objective in utilizing the pseudo-PD function is to minimize overshoots and unduly oscillatory behavior in the motion undertaken by the vehicle, hence minimizing the number of corrective actions needed to converge within the random noise limits of the overall system. Herein, t is defined as a measure of the number of steps taken rather than discrete time. Several methods can be used for determining the optimal PD gains. One such method is presented in Section 4. Each of the three corrective steps are described below in detail: Step 1: After the implementation of M2, the new offsets along the three PSDs are measured in each detector\u2019s frame, Fdi, and the minimum offset, emin, is determined e.g., e1, e2, or e3 Figure 12 a . The vehicle e.g., aFc is subsequently translated by C3 with respect to Fw in an attempt to have at least one LOS hit the center of its corresponding PSD, where M3 = Tdemin 18 and C3 = KpM3 + Kd dM3 dt . 19 Once a LOS is \u201clocked,\u201d the subsequent steps require this LOS to remain locked while the other two LOS are also locked to the centers of their corresponding PSDs. Step 2: There exist two types of errors in the system; rotational and translational. The guidance algorithm addresses each type of error in separate substeps, respectively", + " They are subsequently utilized, along with h1, the distance from the center of aFc to the center of Fd1, in determining the rotations about the desired y and z axes of dFc as y = cos\u22121 e1z 2 \u2212 f 3 2 \u2212 h1 2 \u2212 2h1f3 23 and z = cos\u22121 e1y 2 \u2212 f 2 2 \u2212 h1 2 \u2212 2h1f2 . 24 The minimum offset among the two remaining PSDs is used next in order to determine the last rotation angle about the desired x axis of dFc: x = cos\u22121 emin 2 \u2212 f 1 2 \u2212 h2 \u2212 2hf1 , 25 where f1 is the distance from the z or y projection e3 or e2 determined in Fd3 or Fd2, respectively, to the center of aFc depending on whether the minimum offset emin is e3 or e2 Figure 12 , e.g., emin = e2z if min offset is on PSD2, e3y if min offset is on PSD3, and h = h2 if min offset is on PSD2, h3 if min offset is on PSD3. Step 2b: If e2 is emin, the distance from the x projection of e2,e2x to the center of aFc , f1, and the distance from the z-projection of e2,e2z to the center of aFc , f3, are calculated. They are, subsequently, utilized along with h2, the distance from the center of aFc to the center of Fd2, in determining the rotations about the desired x and z axes of dFc as x = cos\u22121 e2z 2 \u2212 f 3 2 \u2212 h2 2 \u2212 2h2f3 26 and z = cos\u22121 e2x 2 \u2212 f 1 2 \u2212 h2 2 \u2212 2h2f1 " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001885_1.5060414-Figure3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001885_1.5060414-Figure3-1.png", + "caption": "Figure 3: Growth competition against the direction of", + "texts": [ + " From these experiment, one important point has been evidenced: the columnar growth competition. In an epitactic configuration, during the solidification, the first dendrites tend to orientate along the strongest thermal gradient. However, this is conditioned too by the local direction of the solidification front. Thus, it is possible that, in the solidification area, the bottom of the melt pool can leads to vertically oriented first dendrites, and, on the top, to horizontally oriented first dendrites. The figure 2 and figure 3 explain clearly this situation. Analysis of loss of epitaxy in the critical branching zone during E-LMF of misoriented substrate. This phenomenon appears when E-LMF is operated on substrates with a crystallographic orientation misoriented with respect to the surface treated normal. In this case and for particular growth conditions the single crystallinity is lost in a region defined by the angle of misorientation (angle between main crystallographic direction and the normal to the substrate\u2019s surface)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002456_1.25389-Figure5-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002456_1.25389-Figure5-1.png", + "caption": "Fig. 5 Picture of the wind-tunnel model in basic configuration.", + "texts": [ + " Thewingwas plastic foamwith a fiber-epoxy resin skin. After some efforts, everything was designed small enough to be contained in the useful part of the test section of our wind tunnel. Figure 4 reports the sketch of the model in its basic configuration (dimensions in millimeters). Additional features are as follows. Wing surface: 60; 625 mm2, aspect ratio 6, trapezoidal wing. Horizontal tail plane: surface 18; 700 mm2, aspect ratio 2.75 (rectangular). Vertical tail plane: surface 12; 600 mm2, sweep angle 50 deg. Figure 5 shows the picture of themodel in basic configuration (as in Fig. 4), whereas Fig. 6 reports the picture of additional configurations that were tested, with blunt nose and/or leading-edge extensions (LEX). The blunt cone forming the forebody had the same height as the sharp-cone configuration and the same base radius. The frontal area of the model is about 2:33 10 2 m2, which corresponds to a blockage of 2% (referred to the wind-tunnel section). The model was statically stable and accurately balanced", + ", for freestream speeds greater than about 5 m=s, and spinning was always in the same direction for a fixed geometry. On the contrary, changes in the geometry (nose and/or LEX) caused changes in the spin direction. This confirmed what could have been supposed by common knowledge of the flow over cones, according to which any asymmetry on the cone surface would always produce the same effects, thus leading to a single direction of rotation for a given geometry. A. Basic Configuration All results discussed in this section refer to the model in Fig. 5. Many tests were carried out in the range of Reynolds numbers between 4 104 and 2:65 105, but here we report only the most relevant plots to describe the typical features observed. Figure 7 refers to Re 4:8 104 and shows the power spectrum of \u2019 (channel 1) as a function of the reduced frequency defined as k fc=U1. The main peak located at k 5:3 10 3 (0.215 Hz) corresponds to the number of revolutions per second around the yaw axis. Almost the same plot is obtained for the power spectrum of (channel 2), reported in Fig", + " Further tests were thus conducted to assess the influence of nose, wing geometry, tail planes, and their mutual interactions on the spin motion. In this section the following configurations are considered: 1) blunt nose, 2) leadingedge extensions on the wings, and 3) blunt nose LEX. The analysis of the model motion without tail planes is reported in Sec. IV.C. The model with blunt nose and LEX is as reported in Fig. 6. Results for the blunt-nose case obtained at Re 1:28 105 are reported in Figs. 22 and 23. In this configuration the wing is trapezoidal, as in Fig. 5, whereas the nose is as in Fig. 6. The power spectrum reveals the presence of four peaks. The first one is the frequency of revolution around the yaw axis, whereas the others are the third, fourth, and fifth harmonics. This is new compared to the previous spectra presented. Even though the spectrum is very rich, all frequencies are clearly in natural ratio, and thus themotion is periodic and not particularly complex. This is confirmed by the projection of the attractor, shown in Fig. 23, which features some big orbits confined in a relatively small region and visited together with some other smaller ones" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003681_icmtma.2009.473-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003681_icmtma.2009.473-Figure1-1.png", + "caption": "Figure 1. The schematic drawing of traditional magnetic drive system", + "texts": [ + " ZHAO Han [3][4] established and simulated the physics mathematical model and the dynamic property model of the rare earth permanent magnetism gear drive system. The torque calculation formula of system was derived. CHEN Kuang-fei [5] used the finite element method to obtain the magnetic field of permanent magnetism gear and established the torque-angle characteristic of permanent magnetism gear according to the principle of virtual work. On the characteristic of the nonbearing magnetic resistance motor, MASATSUGU TAKEMOTO [6] established the torque computation equation of the rotor. The above traditional magnetic drive system can be showed in Figure 1. Its initiative cylindrical permanent magnet is driven by motor directly, and driven cylindrical permanent magnet is driven by magnetic field coupling and magnetic pole interaction between initiative and driven cylindrical permanent magnet. When the distance between them is increasing, the driving force can be realized by increasing the magnetic field intensity of initiative cylindrical permanent magnet. On the other hand, when the magnetic field intensity of initiative cylindrical permanent magnet increase, it will couple with the motor inner magnetic field, where the motor tending to be damaged" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000411_6.2001-3480-Figure7-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000411_6.2001-3480-Figure7-1.png", + "caption": "Figure 7. Drive Arrangement for Hot Rig", + "texts": [ + " The inverter is also fitted with a separate braking circuit, this allows the deceleration of the disk to match the acceleration rates. Acceleration and American Institute of Aeronautics and Astronautics deceleration rates are ultimately controlled by the inertia of the disk. Typically we can accelerate from zero to mil speed in 30 second and stop in the same time. This rapid speed change capability allows the rig to be driven like a gas turbine, the pressure response is as quick so it is only the temperature that tends to lag. The spindle is driven via a flat belt drive as indicated in Figure 7. Both spindles utilise angular contact bearings fitted with silicon nitride balls to enable high speeds to be achieved. The bearings are oil cooled and the nose of the disk spindle is also air cooled. Cross turbo charger sealing rings are used as the primary oil seal on both spindles and they have proved very reliable under all operating conditions. Figure 8 shows the housing arrangement used to simulate an aerospace application where rapid response was required. To further simulate the engine the disk was machined with typically 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000098_03616967808955302-Figure6-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000098_03616967808955302-Figure6-1.png", + "caption": "Fig. 6 , Voltage diagram i n the general case v - : voltage before cornrnuta7ion", + "texts": [ + "5, to start a new conduction period with the armature axes in the positions a (j) and a (j) and the machine terminal voltage v and the internal angle 6. For the general case of a machine with unequal subtransient d and q-axes reactances, and considering the armature resistive drop the vector diagram will be different from the last one. In such a case the terminal voltage and its internal angle will vary during the conduction period, according to equations (38) and (39) ; from v. and 6 at the beginning of conduction to v and 0 at the end of ~onductlon as shown in fig.6. These variations are ~ten ver~ small for conventional machines. On the other hand, in the general case when the commutation is not instan taneous, the terminal voltage during the commutation period is actip9 in the ~ axis and rotating with it over a ~elec. rad. angle, while the B -component of the voltage is equal to zero during the commutation period as discussed before. This fact is shown in fig.6, where the voltage at the beginning of commutation is given by the vector v and vary during the commutation to become at the end of commutation v~:. J CALCULATION PROCESS At the beginning of (C'DllIIllutalaon) v . ilo%ll!l:j have been con~idered as data; from them, we have deduced the ~aLues ~f all the machine currents, and the right hand side of (33) and (34). However, the left hand side of (33) and (34) are, by definition of vm and om Va Vmsin (a -Om) Vmcos (a-om) (41 ) (42) Thus, equations (33 - 34) define the values of v and 6 ; strictly speaking, these equations have no solution, since the~r righ~\u00b7hand sides contain constant terms and sinusoIdal functions of 20" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002698_iros.2007.4399277-Figure6-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002698_iros.2007.4399277-Figure6-1.png", + "caption": "Fig. 6. Inverse kinematics of avoidance of a gap on a thick wall.", + "texts": [ + " Avoidance of a gap on a thin wall When the link LM is passing through a gap on a thin wall, one lateral DOF is restricted by the gap. Other DOF of the link are free (see Fig. 5). The link LM is located between the gap by rotating the joint Jk, which is located before the wall. There are two patterns of solutions of k\u03c6 , which are right and left rotations. B.4. Avoidance of a gap on a thick wall When the link LM is passing through a gap on a thick wall, two DOF, which are lateral movement and lateral rotation, are restricted by the gap (see Fig. 6). Therefore, two additional DOF are necessary. The joint JM is located in front of the gap by rotating the joint Jk, which is located before the wall. And, the link LM is located so as to be parallel to the center plane of the gap by rotating the joint JM. There are two patterns of solutions for Jk and JM, respectively, and totally there are four patterns of solutions. However, only the case in which JM is before the wall and JM+1 is after the wall is the real solution. B.5. Avoidance of a cylindrical hole in a thin wall When the link LM is passing through a cylindrical hole in a thin wall, two DOF of lateral movements are restricted by the hole (see Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003040_iccasm.2010.5622573-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003040_iccasm.2010.5622573-Figure2-1.png", + "caption": "Figure 2. Crankshaft FE mesh reasonable", + "texts": [ + " FINITE ELEMENT MODEL AND MULTI-BODY DYNAMICS MODEL We use the finite element (FE) software to build the main bearings and crankshaft three-dimensional solid models, and then divide them into the hexahedral mesh elements. To reduce the computational size, the sub-structure reduction method was used. At the same time, in order to avoid stress concentration, results in singular values, we establish the rigid layer and dynamically reduce model. Based on the same dynamic characteristics of the entire system, the freedom of the original structure is significantly compressed. We can use the smaller matrix to improve the solution efficiency, and the accuracy not be impacted significantly. Figure1, Figure2 are respectively models of the main bearing wall and the crankshaft of FE mesh structure. In those models, the rotation axis is Z-axis and vertical axis is Y-axis. 978-1-4244-7237-6/10/$26.00 \u00a9201O IEEE VlS-179 This paper mainly studies the main bearing lubrication analysis, and the crankshaft and main bearing modules come from the FE model and other modules are simplified models. Among various models, we use non-linear bearing joint and other joints to establish coupled non-linear connections, in which the crankshaft and the main bearing wall use EHD joint to connect models" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003564_1.4000520-Figure5-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003564_1.4000520-Figure5-1.png", + "caption": "Fig. 5 Assembly mode that corresponds with L1", + "texts": [], + "surrounding_texts": [ + "c i t t d t c\nN x a \u00aa\nJ\nf\nn s\nc i\nJ\nDownloaded Fr\nsume all six vertices to be Euclidean points, it is not entirely clear if these are all cases of flexible octahedra with one vertex at infinity. But there are good reasons to conjecture that no other cases exist. \u2022 For TSSM manipulators with three parallel axes the problem reduces to a planar one, as these manipulators possess a cylindrical singularity surface compare Ref. 16 . In this case the degenerated cases correspond to the well known self-motions of 3dof revolute-prismatic-revolute RPR manipulators.\n3.2 Special Cases. In the following we are interested in those ases where more than eight intersection points are located on the maginary spherical circle. A necessary condition for this is that he cyclic points of the circle c3 traced by the end point of the hird leg also belong to at least one of the conic sections ki etermined by Fi=0 of Eq. 7 for i=1,2. Therefore, we compute he intersection points of ki and k0 by eliminating x3 from the orresponding equations.\nThe resulting two expressions Li split up into x1 2+x2 2 Ni where\ni is of degree 2 in the unknowns x1 ,x2. Solving these factors for 1 we get the corresponding x3 values as common solution of k0 nd ki after back-substitution of x1. By setting q1\u00aau\u2212b and q2\nu+b the homogeneous coordinates of the intersection points\n0 , J\u03040 ,Ji , J\u0304i of k0 and ki can be written as\nJ0 = 1 : I : 0 T, Ji = qi 2 \u2212 t2 I : qi 2 + t2 : \u2212 2tqiI T 9\nor i=1,2. Moreover, it should be noted that I denotes the imagi-\nary unit and J\u0304 j the conjugate complex point of Jj for j=0,1 ,2. A ketch of the situation at the plane at infinity is given in Fig. 3.\nAs the carrier plane of c3 has to intersect k0 in conjugate omplex points there are the following possibilities left for choosng e\u00aa :\na ei\u00aa Ji , J\u0304i for i=1,2: Now ei with projective line coordinates\nJi J\u0304i = 2qit:0:qi 2 \u2212 t2 T 10\nintersects in the point Ji and J\u0304i with multiplicity 5, but these points are not fivefold points of , which can be shown as follows: We intersect with the line l spanned\nby the origin and the point Ji and J\u0304i, respectively, by inserting its parametric representation into . Then the resulting equation splits up into 216r1 4r2 4F where F is a\npolynomial of degree 8 in the parameter of l. Therefore,\nournal of Mechanisms and Robotics\nom: http://mechanismsrobotics.asmedigitalcollection.asme.org/ on 03/23/2\nei touches in Ji and J\u0304i, which are still fourfold points of . b e0\u00aa J0 , J\u03040 : As J0 and J\u03040 are located on k1 and k2 the line e0 with projective line coordinates\nJ0 J\u03040 = 0:0:1 T 11\nintersects in these points with a multiplicity of 6.\nMoreover, J0 and J\u03040 are sixfold points of , which can be proven by setting v1=v2=0 and v3=1 in Eq. 8 . If we now plug the parametric representation of the circle c3 into the equation of we end up with a polynomial of degree 12 in the unknown h, which finishes the proof.\nIn the following we have to discuss the special cases e1=e2, e0=e1, e0=e2, and e0=e1=e2, respectively:\n\u2022 If e1=e2 holds this line intersects in J1=J2 and J\u03041= J\u03042 with multiplicity 6. This happens if\ndet 2tq1 qi 2 \u2212 t2\n2tq2 q2 2 \u2212 t2 = 4tb u2 \u2212 b2 + t2 = 0 12\nAs b must be greater than zero we only get as solution t =0 and b2=u2+1, respectively. For the latter it can be\nshown as in a that J1=J2 and J\u03041= J\u03042 are only fourfold points of . t=0 will be discussed as the last point of this case study. \u2022 e0 equals e1 for q1t=0. As t=0 will be treated latter on we\nset u=b and assume t 0. In this case J0=J1 and J\u03040= J\u03041 are sevenfold points of , which can be proven analogously to b . Therefore, such a manipulator cannot have more than ten assembly modes.\nBut in this case the given threshold can be refined because for u=b and t 0 the ruled surface generated by s can only have two real generators. Therefore, the spin surface degenerates into two coplanar circles, which can be intersected by a further circle c3 in a maximum of four real points. Moreover, it should be noted that such a manipulator has only 3dof, namely, the translations in the x and y directions, as well as the rotation about s. \u2022 e0 equals e2 for q2t=0. As we assumed w.l.o.g. u 0 and b 0 this can only happen for t=0. But for t=0 we get e0\n=e1=e2 and the points J0=J1=J2 and J\u03040= J\u03041= J\u03042 are eightfold points of , which can again be proven as in case b .\nIf additionally u=b holds the manipulator can only be assembled for r1=r2. Now the ruled surface traced by s is a cylinder of rotation and the point X can only be located in an annulus. Trivially such a manipulator has again only 3dof and a maximum of four real solutions.\nThe results of this case study are summed up in the following theorem.\nTheorem 3. GTSSM manipulators with two parallel rotary axes (a1 and a2 cannot have more than:\ni 12 assembly modes if the axis a3 is parallel to a1 ,a2 ii eight assembly modes if the axis a3 is parallel to the co-\ninciding axes a1=a2 iii four assembly modes if u=b holds\nexcept in the degenerated cases with infinitely many solutions. In order to show that the upper bounds of assembly modes given in theorems 2 and 3 cannot be improved we must present examples, which possess the maximal number of real solutions for the direct kinematics.\nFor the manipulator of theorem 3 i , which corresponds to the Stewart platform, an example with 12 real solutions was given by Lazard and Merlet 6 . An example with eight real solutions for\nFEBRUARY 2010, Vol. 2 / 011009-3\n018 Terms of Use: http://www.asme.org/about-asme/terms-of-use", + "t a l c s\ns t e\n4\no m G\na e t n e a t a\na f f t a o\no s j G p c\ng n i L d\nc D a\nF u\n0\nDownloaded Fr\nhe manipulator of theorem 3 ii can easily be constructed by pplying the same approach used in Ref. 6 because this manipuator is nothing but a special case of the Stewart platform. The onstruction of a manipulator of theorem 3 iii with four real olutions is trivial anyway.\nTherefore, the only open problem in this regard is to demontrate that there exists a parallel manipulator of GTSSM type with wo parallel rotary axes with 16 assembly modes. How such an xample can be constructed is outlined in Sec. 4.\nMaximum Number of Assembly Modes One possibility to generate an example with a maximal number f assembly modes is the iterative algorithm presented by Dietaier 17 , which was used for the construction of a Stewart\u2013 ough platform with 40 real solutions. We will solve this problem with the help of a graphical tool fter reducing it to a planar one. As we only have to find one xample with 16 real solutions for the direct kinematics we make he following assumptions: As u=0 does not reduce the maximal umber of assembly modes compare theorems 2 and 3 we set u qual to zero. As a consequence the RSSR mechanism with parllel rotary axes degenerates to an ordinary 4-bar mechanism in he xy-plane. Moreover, we can assume that the rotary axis a3 is lso located in this plane.\nNow the circles c3 as well as the circle cF traced by the point B3 bout s\u00aa B1 ,B2 lie in z-parallel planes where F denotes the ootpoint on s with respect to B3 see Figs. 2 and 3 . For the ollowing considerations we also need the spheres 3 and F deermined by their centers A3 and F, respectively, and by c3 3 nd cF F. Trivially the carrier plane of the intersection circle c f these two spheres is also parallel to the z axis. These three planes are mapped by the horizontal projection nto the straight lines l3, lF, and l , respectively. Now the interection point G1,2 of lF and l corresponds to the horizontal proection of the intersection point G1 ,G2 of c3 and cF if and only if\n1,2 is also located on l3. For the case lF= l \u21d2lF= l = l3 all oints G1,2 l3 with G1,2 A3 r3 correspond to points G1 ,G2 of 3=cF.\nIn the following we have to find parameters such that the curve generated by G1,2 during the motion of the the 4-bar mechaism intersects l3 in four real points Li . These points correspond n each case with two real intersection points of c3 and cF if i A3 r3 holds. It should be noted that Li A3 =r3 would yield a ouble solution G1=G2 .\nBased on these considerations it is not difficult to find such a onfiguration set with the help of a graphical tool e.g., EUKLID YNAGEO 1 . For the example illustrated in Fig. 4 the parameters re as follows:\n1\nA1\nA2A3\nB1\nB2\nF\ng\u2032\nG\u2032 1,2\nl\u0393\nlF\nl3\na3\nL1\nL2\nL3\nL4\nig. 4 Screenshot of the software EUKLID DYNAGEO, which was sed to generate an example with 16 assembly modes\nhttp://www.dynageo.com.\n11009-4 / Vol. 2, FEBRUARY 2010\nom: http://mechanismsrobotics.asmedigitalcollection.asme.org/ on 03/23/2\nt = 100, r1 = 58, r2 = 119, r3 = B3F = 45\nA3 = 25,0 , b = 85, B1F = 15, B2F = 70 13\nand a3 equals the x-axis. Due to the symmetry of the chosen example the curve g for the second assembly mode of the 4-bar mechanism can be obtained by reflecting g on a3. This yields the\npoints L\u0304i for i=1, . . . ,4. We only illustrate four solutions see Figs. 5\u20138 of the direct kinematics problem because the remaining 12 can be obtained by reflecting the given four solutions on the xy-plane and by applying a further reflection to the resulting eight configurations on the xz-plane.\n\u03c01\n\u03c02\n\u00af\nFig. 8 Assembly mode that corresponds with L4\nTransactions of the ASME\n018 Terms of Use: http://www.asme.org/about-asme/terms-of-use", + "b c\nl\n5\nR 6 n u t a r\nt p t\nA\nt e\nu t\nT i\nS\nL\nL\nL\u0304\nL\u0304\nJ\nDownloaded Fr\nMoreover, it should be noted that the given example not only elongs to the class of GTSSM manipulators but also to its sublass of TSSM type manipulators.\nThe numerical data of the platform-anchor-points Bi of the ilustrated four solutions are given in Table 1.\nConclusion We demonstrated that the circularity of the spin-surface for an\nSSR mechanism 18\u201322 with parallel rotary axes is 4 instead of as given in the literature 6 . As a consequence a parallel ma-\nipulator of GTSSM type with two parallel rotary joints can have p to 16 solutions instead of 12 compare Ref. 7 . We showed hat this upper bound cannot be improved by constructing an exmple for which the maximal number of assembly modes is eached.\nMoreover, we analyzed all parallel manipulators of GTSSM ype with two parallel rotary axes where more than 4 2=8 oints are located on the imaginary spherical circle. As one of hese special cases the Stewart platform appears.\ncknowledgment This research was carried out as part of the project under Conract No. S9206-N12, which was supported by the Austrian Scince Fund FWF .\nThe author would also like to thank the reviewers for their seful comments and suggestions, which have helped to improve he quality of this article.\nable 1 Numerical data of the platform-anchor-points Bi of the llustrated four solutions\nolution Point x y z\n1 B1 10.19 57.10 0 B2 55.91 110.53 0 B3 25 37.42 24.98 2 B1 55.02 18.36 0 B2 38.16 101.67 0 B3 25 27.59 35.55 3 B1 56.25 14.14 0 B2 14.22 33.39 0 B3 25 33.65 29.88 4 B1 16.30 55.66 0 B2 15.33 29.33 0 B3 25 41.13 18.25\nournal of Mechanisms and Robotics\nom: http://mechanismsrobotics.asmedigitalcollection.asme.org/ on 03/23/2\nReferences 1 Robertson, G. D., and Torfason, L. E., 1975, \u201cThe Qeeroid\u2014A New Kinematic\nSurface,\u201d Proceedings of the Fourth World Congress on the Theory of Machines and Mechanisms, pp. 717\u2013719. 2 Fichter, E. F., and Hunt, K. H., 1977, \u201cMechanical Couplings - A General Geometrical Theory,\u201d ASME J. Eng. Ind., 99, pp. 77\u201381. 3 Hunt, K. H., 1978, Kinematic Geometry of Mechanisms, Clarendon, Oxford. 4 Hunt, K. H., 1983, \u201cStructural Kinematics of In-Parallel-Actuated Robot-\nArms,\u201d ASME J. Mech., Transm., Autom. Des., 105, pp. 705\u2013712. 5 Merlet, J.-P., 1989, \u201cManipulateurs parall\u00e8les, 4eme partie: mode\nd\u2019assemblage et cin\u00e9matique directe sous forme polynomiale,\u201d INRIA Technical Report No. 1135. 6 Lazard, D., and Merlet, J.-P., 1994, \u201cThe True Stewart Platform Has 12 Configurations,\u201d Proceedings of the IEEE International Conference on Robotics and Automation, pp. 2160\u20132165. 7 Merlet, J.-P., 2006, Parallel Robots, 2nd ed., Springer, New York. 8 Merlet, J.-P., 1992, \u201cDirect Kinematics and Assembly Modes of Parallel Ma-\nnipulators,\u201d Int. J. Robot. Res., 11 2 , pp. 150\u2013162. 9 Husty, M., 2000, \u201cE. Borel\u2019s and R. Bricard\u2019s Papers on Displacements With\nSpherical Paths and Their Relevance to Self-Motions of Parallel Manipulators,\u201d Proceedings of the International Symposium on History of Machines and Mechanisms, M. Ceccarelli, ed., Kluwer, Dordrecht, The Netherlands, pp. 163\u2013172. 10 Husty, M., and Karger, A., 2000, \u201cSelf-Motions of Griffis-Duffy Type Platforms,\u201d Proceedings of the IEEE International Conference on Robotics and Automation, pp. 7\u201312. 11 Karger, A., 2008, \u201cNew Self-Motions of Parallel Manipulators,\u201d Advances in Robot Kinematics\u2014Analysis and Design, J. Lenarcic and P. Wenger, eds., Springer, New York, pp. 275\u2013282. 12 Karger, A., and Husty, M., 1998, \u201cClassification of All Self-Motions of the Original Stewart-Gough Platform,\u201d CAD, 30, pp. 205\u2013215. 13 Bricard, R., 1897, \u201cM\u00e9moire sur la th\u00e9orie de l\u2019octa\u00e8dre articul\u00e9,\u201d J. Math. Pures Appl., 3, pp. 113\u2013148. 14 Stachel, H., 1987, \u201cZur Einzigkeit der Bricardschen Oktaeder,\u201d J. Geom., 28, pp. 41\u201356. 15 Stachel, H., 2002, \u201cRemarks on Bricard\u2019s Flexible Octahedra of Type 3,\u201d Proceedings of the Tenth International Conference on Geometry and Graphics, pp. 8\u201312. 16 Nawratil, G., 2009, \u201cAll Planar Parallel Manipulators With Cylindrical Singularity Surface,\u201d Mech. Mach. Theory, 44 12 , pp. 2179\u20132186. 17 Dietmaier, P., 1998, \u201cThe Stewart-Gough Platform of General Geometry Can Have 40 Real Postures,\u201d Advances in Robot Kinematics: Analysis and Control, J. Lenarcic and M. L. Husty, eds., Kluwer, Dordrecht, The Netherlands, pp. 7\u201316. 18 Arakelian, V. H., 2007, \u201cComplete Shaking Force and Shaking Moment Balancing of RSS\u2019R Spatial Linkages,\u201d Proc. Inst. Mech. Eng., Part K: Journal of Multi-body Dynamics, 221 2 , pp. 303\u2013410. 19 Chaudhary, H., and Saha, S. K., 2008, \u201cAn Optimization Technique for the Balancing of Spatial Mechanisms,\u201d Mech. Mach. Theory, 43, pp. 506\u2013522. 20 Chung, W.-Y., 2005, \u201cMobility Analysis of RSSR Mechanisms by Working Volume,\u201d ASME J. Mech. Des., 127, pp. 156\u2013159. 21 Lin, P. D., and Hsieh, J.-F., 2007, \u201cA New Method to Analyze Spatial Binary Mechanisms With Spherical Pairs,\u201d ASME J. Mech. Des., 129, pp. 455\u2013458. 22 Ting, K.-L., and Zhu, J., 2005, \u201cOn Realization of Spherical Joints in RSSR Mechanisms,\u201d ASME J. Mech. Des., 127, pp. 924\u2013930.\nFEBRUARY 2010, Vol. 2 / 011009-5\n018 Terms of Use: http://www.asme.org/about-asme/terms-of-use" + ] + }, + { + "image_filename": "designv11_69_0000749_pime_proc_1973_187_029_02-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000749_pime_proc_1973_187_029_02-Figure1-1.png", + "caption": "Fig. 1. Typical diaphragm spring clutch assembly", + "texts": [], + "surrounding_texts": [ + "Having seen that an analogue computer can be applied to the problem, an attempt was made to develop a simulation which would produce the type of results measured on a vehicle. Previous workers in the field had represented the vehicle by a mathematical model comprising of two inertias connected by a shaft, as shown in Fig. 8. In this I , represents the inertia of the vehicle, referred through the gearing to the gearbox input shaft speed, and K is the stiffness of the transmission, also referred to the gearbox input. The significance of Il depends on whether the clutch is fully engaged and not slipping, in which case it represents the combined inertia of engine, clutch cover Prcc lnstn Mech Engrs 1973 Vol 107 27/73 at WEST VIRGINA UNIV on June 5, 2016pme.sagepub.comDownloaded from CLUTCH JUDDER IN AUTOMOBILE DRIVELINES 371 Proc lnstn Mech Engrs 1973 Vol 187 27/73 at WEST VIRGINA UNIV on June 5, 2016pme.sagepub.comDownloaded from 372 R. F. jARVIS AND R. M. OLDERSHAW 0 5 I.0 1.5 OV 0 T I M E seconds Fig. 7. Practical trace showing natural frequency a s 2.5 Hz Fig. 8. Two-inertia system and driven plate, or whether the clutch is slipping, in which case Il represents the driven plate only. When checked against the practical result, such as in Fig. 7, the correlation on frequency could not be obtained with the calculated value of stiffness, and was only achieved by using an artificial value of 6 8 4 Nm/rad instead of the 120 Nm/rad calculated. In an indirect gear the difference between the output torque and the input torque is supplied by the reaction of the gearbox casing. With an engine and gearbox assembly on rubber mountings this whole inertia must participate in the vibration. This system is represented in Fig. 9 as a branched system with three inertias. This simulation gave good agreement on the natural frequency and proved to be adequate for examining many of the possible or postulated causes of judder, but the representation of damping in the system proved to be unsatisfactory. Linear behaviour had been obtained by assuming damping proportional to velocity-and a sufficient amount had to be associated with the vehicle inertia I,. This damping appeared then as a drag to be overcome by the engine, which would have given the car a maximum speed of only 32 km/h. In practice it is known Crankshaft Driven and flywheel * plate Vehicle i ne r t i a C L U T C H JU D D E R IN A U T O M O B IL E D R IV E L IN E S 373 n 9 I \u2018 I Proc lnstn M ech Engrs 1973 V ol 187 27/73 at W E S T V IR G IN A U N IV on June 5, 2016 pm e.sagepub.com D ow nloaded from at WEST VIRGINA UNIV on June 5, 2016pme.sagepub.comDownloaded from CLUTCH JUDDER IN AUTOMOEILE DRIVELINES 375 2oo, T I M E s a Driven plate speed. b Speed difference across clutch. c Equivalent wheel speed. assumed that the effect of misalignment could be represented by a sinusoidal torque variation which synchronized with the slip of the driven plate relative to the pre, .sure plate and flywheel. The justification of this assumption is considered in detail later in the paper. Fig. 15 shows clutch take-up performances when the disturbing torque was set to f 6 7 5 N m and f2 .0Nm respectively. Judder is clearly evident coming in and then fading in the manner to be expected from experience. Although the effect of clutch misalignment may be represented in this manner, any form of exciting force which produces a sinusoidal torque variation will have the same effect." + ] + }, + { + "image_filename": "designv11_69_0003793_imece2009-10092-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003793_imece2009-10092-Figure1-1.png", + "caption": "Figure 1. NEURAL NETWORK STRUCTURE OF RECEPTIVE FIELDWEIGHTED REGRESSION", + "texts": [ + " Several authors have used local polynomial of low order to model the relationship between input and output data within each receptive field, particularly linear models because the achieve a favorable compromise between computational complexity and quality of result (25): y\u0302k = (x\u2212 ck) T \u2219bk+b0,k = xT 1 \u2219\u03b2k (4) where \u03b2k = (bT k ,b0,k) T denotes the parameters of the locally linear model, formed by the coefficent vector bk and the bias b0,k of the linear model and x1 = ( (x\u2212 ck) T ,1 )T is a compact form of the center-subtracted, augmented input vector. Figure 1 shows the architecture of the neural network for the implementation of local, receptive field-based learning. Bontempi et al. (22) propose a model identification methodology based on the use of an iterative optimization procedure to select the best local model among a set of different candidates. This technique is based on recursive least squares methods to compute PRESS in an incremental way. Schaal et al. (25) introduce a modified method in which the parameters of the locally linear model as well as the size and shape of the receptive field itself are learned independently without the need for competition or any kind of communication" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003889_pes.2010.5589727-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003889_pes.2010.5589727-Figure2-1.png", + "caption": "Fig. 2. Experimental setup of sensorless DTFC drive system using matrix converter fed IPM synchronous motor.", + "texts": [ + " 1 \u02c6 \u02c6\u02c6 ( ) /re d q d q\u03c9 \u03bb \u03bb \u03bb \u03bb \u03b3= \u2212& % % (4) 2 \u02c6 /f S dR i\u03bb \u03b3=& % (5) To improve the dynamic behavior, a PI estimator is employed. \u02c6 \u02c6\u02c6 ( / )*( )re P I d q d qK K S\u03c9 \u03bb \u03bb \u03bb \u03bb= + \u2212% % (6) From experiments, it is found that the position error is increased with the increase of the load. Using a simple and fast compensation strategy, which is immune to the noise, the error signal, rather than the estimated position, is corrected at the input of the PI-PLL observer by an offset angle. A properly selected hysteresis-band enables the offset independent of torque noise. Fig. 2 shows the overall experimental setup. The matrix converter drives a 230V, four-pole, 0.97kW IPM synchronous motor mechanically coupled with a PMDC machine. The DC machine is fed by a diode rectifier and H-bridge converter, which provides load torque. The effect of PM flux adaptation on the performance of the sensorless drive was investigated experimentally. Fig. 3 shows that the errors in the rotor position and current are reduced after the PM flux estimation is enabled at 3.4s. The position error is decreased by 2\u00b0 and d-axis current error converges to zero from \u22120" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000162_robot.1998.677383-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000162_robot.1998.677383-Figure1-1.png", + "caption": "Fig. 1 A model of the human finger", + "texts": [ + " The motion of the finger is restricted within the sagittal plane. Assuming that the four fingers of the hand excepting the thumb perform the similar movement, they are thus all modeled in the same way. The model includes five kinds of muscles: (1) the extensor digitorum (ED); (2) the flexor digitorum profundus (FDP); (3) the flexor digitorum superficialis (FDS); (4) the interosseous muscle (Int); (5 ) the lumbricalis muscle (Lum), and three kinds of tendons: (6) the medial band (MB); (7) the lateral band (L); and (8) the terminal band (T) (Fig.1). The excursions in the muscles or tendons have positive values when they contract ,and have negative values when they relax. The extensor excursion-joint model[ 141 is represented as follows: (Landsmeer\u2019s model 1) - R e , (2-1) and the flexor excursion-joint model[ 141 is represented as follows: (Landsmeer\u2019s model 3) e=(R+R\u20198)8 , (2-2) where e\u2019 is the tendon excursion, R is the radius of a pulley on the joint, R and R\u2019 are constant and e is the joint angle. Consequently, the excursions of the muscles are described as follows: eELt=-R1 l e M P - R l 2 e P I P (2-3) (2-4) eFDS=(R3 1+Rj 10 M P ) ~ I U P + ~ ~ ~ ~ P I P (2-5) =Iit41e M P - R 4 2 @ PIP (2-6) + i R ; 2 e P I P ) e p I P - ( R 1 3 + R 2 3 + R i 3 ~ D I P ) e D I P 9 (2-7) where e# represents the excursion of #, &j and R\u2019,j are defined in Table 1, and e, represents the joint angle of i" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001415_ijmr.2006.010703-Figure4-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001415_ijmr.2006.010703-Figure4-1.png", + "caption": "Figure 4 Schematic of cylindrical grinding", + "texts": [ + " Among them, a threshold force can represent the sum of ploughing and sliding force components. Only when the amount is above the threshold will it contribute to material removal. For easy to grind material, the energy consumption associated with sliding and ploughing will be insignificant compared to chipping energy, and hence almost all energy is used for material removal and the threshold force can be ignored. Therefore, the grinding force can be modelled to be proportional to the material removal rate. For the cylindrical plunge grinding as illustrated in Figure 4, the material removal can be computed as s wb d v (8) where bs is the grinding wheel width, dw is the diameter of work piece and v is the actual infeed velocity. However, because of final stiffness associated with work, wheel and contact, the actual radial infeed velocity will be different from the commanded radial infeed velocity (Malkin, 1989). Neglecting wheel wear for the moment, continuity requires that the difference between the controlled )(tu and the actual )(tv infeed velocities be equal to the time rate change of the radial elastic deflection of the grinding system: ( ) ( ) d u t v t dt (9) n e F k (10) where Fn is the normal force component and ke is the effective stiffness" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002928_978-3-540-70534-5_16-Figure16.22-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002928_978-3-540-70534-5_16-Figure16.22-1.png", + "caption": "Figure 16.22: Bug1 and Bug2 examples [Ng, Br\u00e4unl 2007]", + "texts": [ + "23 shows two more examples that further demonstrate the DistBug algorithm. In Figure 16.23, left, the goal is inside the E-shaped obstacle and cannot be reached. The robot first drives straight toward the goal, hits the obstacle, and records the hit point, then starts boundary following. After completion of a full circle around the obstacle, the robot returns to the hit point, which is its termination condition for an unreachable goal. To point out the differences between the two algorithms, we show the execution of the algorithms Bug1 (Figure 16.22, left) and Bug2 (Figure 16.22, right) in the same environment as Figure 16.21, right. Figure 16.23, right, shows a more complex example. After the hit point has been reached, the robot surrounds almost the whole obstacle until it finds the entry to the maze-like structure. It continues boundary following until the goal is directly reachable from the leave point. 16.10 Dijkstra\u2019s Algorithm Reference [Dijkstra 1959] Description Algorithm for computing all shortest paths from a given starting node in a fully connected graph. Time complexity for naive implementation is O(e + v2), and can be reduced to O(e + v\u00b7log v), for e edges and v nodes" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003666_icinfa.2010.5512357-Figure4-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003666_icinfa.2010.5512357-Figure4-1.png", + "caption": "Fig. 4 Top view of the solar vehicle", + "texts": [ + " For example, when the dynamics is unclear, it leads to a task that derives the dynamics of the module. Based on the kinematic relationship, the following equations are derived where the forward velocity and the steering angle are considered as inputs which are also the boundary conditions of the other two modules, is the Cartesian coordinate of the centre of mass of the vehicle body, is the heading angle of the vehicle and is the radius of the cross-section of the cylindrical vehicle body (see Fig. 4). Planning various tasks corresponds to Step four in the method. These tasks are then given to students. Once they complete these tasks, the system candidate is completely designed. Then, the performance of the entire system can be tested and evaluated. The four-wheel chassis module based system candidate can be verified and tested through simulation using computer tools like MATLAB/SIMULINK. Simulating the dynamics of the solar vehicle yields certain responses with respect to cloud shading effect and steering controller which students know little about" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002821_6.2007-4340-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002821_6.2007-4340-Figure1-1.png", + "caption": "Figure 1. Flapping-wing MAV model14", + "texts": [ + "12 experimentally studied the performance of an aquatic propulsion system inspired from the uniform swimming mode to investigate the effects of flapping parameters on the thrust force and the hydromechanical efficiency. In their recent study,13 Muniappan et al. concentrated on the effect of flap angle and flapping frequency on the lift characteristics of the flapping wing micro air vehicle. Jones and Platzer14 demonstrated a radio-controlled micro air vehicle propelled by flapping wings in a biplane configuration (Figure 1). The experimental and numerical studies by Jones et al.14\u201317 and Lai and Platzer18 on flapping-wing propellers points at the gap between numerical flow solutions and the actual flight conditions over flapping wings. Tuncer and Kaya19\u201321 investigated the optimization of flapping motion parameters for maximizing thrust and propulsive efficiency of single flapping airfoils and airfoils flapping in a biplane configuration. More recently, Kaya et al.22 optimized flapping parameters of sinusoidal plunge and pitch motions of NACA0012 airfoils in a biplane configuration using moving and deforming overset grids (Figure 3)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003428_icicisys.2010.5658813-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003428_icicisys.2010.5658813-Figure1-1.png", + "caption": "Figure 1. Parallel-type double inverted pendulum system", + "texts": [ + " More specifically, the mathematical model in state space is derived as follows: 0 0 1 0 0 0 0 0 XI 0 0 0 1 0 0 0 0 XI 8 1 0 0 0 0 0 0 0 0 81 Xl AfJ A(mlg/I - fJh) -AfJ AfJh XI 0 0 0 0 81 JI JI JI JI 8 1 x, 0 0 0 0 0 0 1 0 x, 8, 0 0 0 0 0 0 0 1 8, x2 0 0 0 0 0 0 0 0 x, 8, -AfJ AfJh 0 0 AfJ A(m,g/, - fJh) 0 0 8, J, J, J, J, 0 0 0 0 0 A/lml 0 + JI [:: ] 0 0 0 0 0 0 Am212 J, Substitute the value of the parameters mentioned in Tab.1 into matrices A and B, the matrices as follows can be obtained: 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 19.782 8.8879 0 0 -19.782 7.9128 0 0 A= 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 -19.783 7.9128 0 0 19.782 8.8879 0 0 0 0 0 0 0 l.7143 0 B= 0 0 0 0 0 0 l.7143 c. Comparison with Newtonian Mechanics Modeling Choose the two pendulums in Fig.1 as research objects, the force analysis of pendulum 1 is shown in Fig.3. The resultant force acting upon the two pendulums in the vertical direction is: (11) (12) The resultant force acting upon the two pendulums in the horizontal direction can be obtained as: d2 \u2022 N1 + Sh = m1-2 (X1 -/1 smB1 ) (13) dt d2 \u2022 N2 - Sh =m2-2 (x2 -/2 smB2 ) (14) dt The tension generated by the spring along the slide rail is: \ufffdw Lo L S =k ( L-L)-=k\ufffdw ( I--)=k\ufffdw ( l- ) h s O L s L s .J 2 2 d +\ufffd(l) matter which is based on Lagrange Equation or Newtonian Mechanics" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002877_epe.2007.4417421-Figure3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002877_epe.2007.4417421-Figure3-1.png", + "caption": "Fig. 3: Modified bearing housing with eccentric sleeves and digital camera.", + "texts": [ + " As already mentioned the identification of the different asymmetry components is done using a digital camera placed at different angular positions as well as on both sides of the shaft. For the calculation of magnitude and direction of the different phasors in Fig. 1 using data of the digital camera it is important to consider the fact that the measurement of the air gap length with the camera is done at a different axial position than the positioning of the bearings (see also Fig. 2). This displacement is important especially when the axis of the rotor is not exactly aligned with that of the stator lamination \u2013 which generally is the case. In Fig. 3 the modified bearing housing with the two eccentric sleeves is shown together with the digital camera mounted. As already mentioned there are four possible positions to measure the airgap optically through holes in the housing. For the detection of the airgap length it is essential to realise a proper illumination inside the machine. Though the need for an illumination is obvious, its proper placing and tuning is tricky and has a strong influence on the magnitude of the airgap detected. It was found that a balanced illumination of both the foreground and the background (trough the air gap) leads to the best results" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001203_esda2006-95425-Figure6-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001203_esda2006-95425-Figure6-1.png", + "caption": "Fig. 6 \u2013 Reaction damper force on body and wheel", + "texts": [ + " Here we want to influence the wheel load by controlling the active dampers to finally influence the braking force and the braking slip. But how can the wheel load be influenced by an active damper? How does the causal direction look like? To find out more about this we want to look at a simple switch from hard to soft damping in detail. For this we look at the characteristic diagram shown in Fig. 5 first. There we can see that for any given velocity of the damper the soft damping produces a smaller damper force than the hard setting. 4 Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 07/06/2018 T In Fig. 6 we see a part of a quarter-car model where only the damper force is shown. The force of the spring is not of any interest for the following thought. We consider the damper to be in rebound at a constant velocity. In rebound the damper force points upward with respect to the wheel and downward with respect to the body. That means that the damper force reduces the wheel load in rebound, because the damper force was not there, the wheel would not be pulled upward and therefore the wheel load would be higher than with the damper force being present" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003334_icelmach.2010.5608064-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003334_icelmach.2010.5608064-Figure1-1.png", + "caption": "Fig. 1. Cross-sectional geometries of the three simulated motors.", + "texts": [ + " TABLE I MAIN PARAMETERS OF THE SIMULATED MOTORS Parameter Motor 1 Motor 2 Motor 3 Number of pole pairs 2 2 10 Number of phases 3 3 3 Number of stator slots 48 48 180 Number of rotor slots 40 0 100 Stator inner diameter (mm) 200 200 2160 Rotor outer diameter (mm) 198.4 196.4 2142 Radial air-gap length (mm) 0.8 1.8 9 Core length (mm) 249 249 1744 Connection star star star Rated voltage (V) 380 380 2900 Rated frequency (Hz) 50 50 9.66 Rated slip 0.016 0 0 Rated current (A) 72 59 703 Rated power (kW) 37 37 3400 Rated torque (kNm) 0.24 0.236 560 Simulation point Voltage (V) 20 20 2900 Current (A) 65 126 703 Frequency (Hz) 1.6 0.8 9.66 Slip 0.5 0 0 Torque (kNm) 0.223 0.23 560 Fig. 1 presents the cross-sectional geometries of the motors. Harmonic Torque Suppression by Manual Voltage Injection I. K\u00e4rkk\u00e4inen, A. Arkkio P XIX International Conference on Electrical Machines - ICEM 2010, Rome 978-1-4244-4175-4/10/$25.00 \u00a92010 IEEE ( ) ( )( ) ( ) ( )( ) 1 2 1 2 mkr mkr mkt mkt mkr mkt j m k t j m k t mkr mkr j m k t j m k t mkt mkt B B A e A e A e A e \u03d5 \u03c9 \u03b3 \u03d5 \u03c9 \u03b3 \u03d5 \u03c9 \u03b3 \u03d5 \u03c9 \u03b3 \u2212 + \u2212 + + + \u2212 +\u2212 + + + \u2212 +\u2212 + + = + + + Motor 1 has a double-cage squirrel rotor with closed slots and the stator has semi-closed slots" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001979_50037-5-Figure7.80-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001979_50037-5-Figure7.80-1.png", + "caption": "Figure 7.80 Symmetry constraints applied on three horizontal areas.", + "texts": [ + " Also, Contact Manager summary information frame should be closed. Before the solution process can be attempted, solution criteria have to be specified. As a first step in that process, symmetry constraints are applied on the half-symmetry model. From ANSYS Main Menu select Solution \u2192 Define Loads \u2192 Apply \u2192 Structural \u2192 Displacement \u2192 Symmetry BC \u2192 On Areas. The frame shown in Figure 7.79 appears. Three horizontal surfaces should be selected by picking them and then clicking [A] OK. As a result, image shown in Figure 7.80 appears. The next step is to apply constraints on the bottom surface of the block. Form ANSYS Main Menu select Solution \u2192 Define Loads \u2192 Apply \u2192 Structural \u2192 Displacement \u2192 On Areas. The frame shown in Figure 7.81 appears. After selecting required surface (bottom surface of the block) and pressing [A] OK button, another frame appears in which the following should be selected: DOFs to be constrained = All DOF and Displacement value = 0. Selections are implemented by pressing OK button in the frame" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001089_s11741-005-0012-3-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001089_s11741-005-0012-3-Figure1-1.png", + "caption": "Fig. 1 Structure of crown gear coupling", + "texts": [ + " It is concluded that the meshin~ of conjugate CGC is line-contact, there are several paim of teeth engage simultaneottsly, and non-coqjugate CGC has point-contact condition of meshing and only 2 pairs of teeth engage in theory. Key words kinematic analysis, crown gear coupling, meshing. 1 Introduction Crown gear coupling (CGC) is a type of important mechanical parts widely used in metallurgy, mine, transport and ship machinery. It is developed from gear coupling to meet the needs of large and changing angle between coupled shafts and greater torsion with small size and high reliability E~ . CGC consists of two pairs of crown gears and internal gears. Fig. 1 shows the slructure of a CGC. Benlder studied the geometry and mechanics of CGC and gave formulas used as an imperfect design theory and method. Natashima and Gerhardt have studied the contact behavior and operating chamc~rist ics t2\"s~ . The kinematic analysis was based on the simplified plane development and the mechanical analysis on the dead load, while the crown curve was not studied [4'5~ . The real 3-D mo~ion state of CGC is very complicated and there exist problems of dynamic load caused by uniform rotation and vibration t6~ " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001118_0471758159.ch5-Figure5.38-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001118_0471758159.ch5-Figure5.38-1.png", + "caption": "FIGURE 5.38 A dc motor illustrating (a) physical construction, (b) brushes and commutator, and (c) arc suppression elements.", + "texts": [ + " The purpose of this section is to highlight these problem areas and increase the awareness of the reader for their potential to create EMC problems. DC motors are used to produce rotational motion, which can be used to produce translational motion using gears or belts. They rely on the property of magnetic north and south poles to attract and like poles to repel. A dc motor consists of stationary windings or coils on the stator, along with coils attached to the rotating member or rotor, as illustrated in Fig. 5.38a. The coils are wound on metallic protrusions, and a dc current is passed through the windings, creating magnetic poles. A commutator consists of metallic segments that are segmented such that the dc current to the rotor windings can be applied to the appropriate coils to cause the rotor to align with or repel the stator poles as the rotor rotates. Carbon brushes make contact with the rotor segments and provide a means of alternating the current and magnetic fields of the rotor poles using a dc current from a source, as illustrated in Fig. 5.38b. As the current to the rotor coils is connected and disconnected to the dc source through the commutator segments, arcing at the brushes is created as a result of the periodic interruption of the current in the rotor coils (inductors). This arcing has a very high-frequency spectral content, as we saw in Chapter 3. This spectral content tends to create radiated emission problems in the radiated emission regulatory limit frequency range between 200 MHz and 1 GHz, depending on the motor type. In order to suppress this arcing, resistors or capacitors may be placed across the commutator segments as illustrated in Fig. 5.38c. These can be implemented in the form of capacitor or resistor disks that are segmented disks of capacitors or resistors attached directly to the commutator or in resistive ring placed around the commutator. In some cases it may be necessary to insert small inductors in the dc leads to block those noise currents that are not completely suppressed by the capacitor or resistor disks. An additional source of high-frequency noise and associated radiated and conducted emission comes not from the motors themselves but from the driver circuits that are used to change the direction of rotation to provide precise position control of the motor" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002456_1.25389-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002456_1.25389-Figure1-1.png", + "caption": "Fig. 1 Wind-tunnel test section setup, flow from right to left.", + "texts": [ + " Because of the formidable complications imposed by the first option, the simplest solution was to support the model by its center of gravity, allowing only rotations and employing a sting whose axis would be aligned with the freestream flow. To assess how well this 3-degree-of- freedom gimbaled rig represents the full 6-degree-of-freedom spin motions, free-flight tests (e.g., in vertical wind tunnels) would be needed. However, for the purpose of investigating the possible chaotic nature of spin, which requires two independent spinmotions, the present configuration with only 3 degrees of freedom is expected to be satisfactory. Figure 1 reports a picture of the wind-tunnel setup. A tripod is set in the divergent so as to anchor the sting to the wind tunnel. Themodel had to be connected to the sting by a universal joint and a hinge around the yaw axis to allow the three rotations \u2019, , . The joint is shown in Fig. 2 along with the degrees of freedom ( denotes the yaw rotation). The model is anchored to the universal joint through four screws (see the holes on the picture) so that its lower side would be visible on Fig. 2 if bolted" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001644_icarcv.2006.345182-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001644_icarcv.2006.345182-Figure2-1.png", + "caption": "Figure 2. Coordinates of mobile manipulator and the conference coordinates of workspace", + "texts": [ + " They add the input of system stability in vehicle fuzzy navigation so that the mobile manipulator can avoid stably unknown or dynamic obstacles. Integration of robust controller and modified Elman neural network (MENN) is to deal with uncertainties. III. COORDINATING CONTROL FOR HEBUT-\u2161 MOBILE MANIPULATOR Here we will introduce our mobile manipulator system, which consists of a five degree-of-freedom manipulator mounted on a differentially driven mobile platform, see in Fig.1. The Cartesian coordinates of the mobile manipulator is shown as Fig.2 Where, O\u2014XYZ is the conference coordinates of the workspace; G\u2014XGYGZG is the coordinates for the center of mobile platform; G\u2014XMYMZM is the coordinates for the base of manipulator. A wheeled mobile manipulator is a nonholonomic system, the base of which is subject to nonholonomic kinematics constraints. The kinematics of the wheeled mobile platform is subject to velocity constraint model. Suppose that the goal posture and rotation angle for each joints are known, to get the posture of the end-effectors, we need two steps: 1\uff09 Trough the mobile platform center posture [ ]T G G Gx y \u03b8 , we can calculate the transform matrix from mobile platform to the conference coordinates: 0 ( , ) ( , ) ( , 90 )M G G GA Trans x x Trans y y Rot z \u03b8= \u2212 + (1) 2\uff09 The posture of the mobile manipulator 0 5T 0 0 5 5 M MT A T= (2) 5 MT is the transformation from end coordinates of the manipulator to the basement coordinates G\u2014XMYMZM" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002817_660033-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002817_660033-Figure1-1.png", + "caption": "Fig. 1 shows Sommerfeld number So as functions of e c centricity, ϵ, or relative minimum film thickness, h o , re spectively, and relative bearing width, \u03b2. Such a diagram can be used for all values of load and speed, for different", + "texts": [ + " In engine bearings, there are almost no periods in which constant load rotates with constant half shaft speed for a long t ime. But this case gives a very good impression of what happens in a bearing i f the load rotates with almost half shaft speed. Load cannot be supported by a film created by shaft rotation as usual, but must be supported by a squeeze effect. That means that minimum film thickness decreases even i f the load is very low. Calculations of real engine bearings and experiments will show this effect. For the third component of film pressure caused by squeeze effect, diagrams similar to Fig. 1 can be calculated. Fig. 5 shows the load capacity SoV in dimensionless form as func tions of relative eccentricity or relative minimum film thick ness and bearing width. Negative values for the relative eccentricity are neces sary for radial movements of the shaft towards the center of the bearing. The combination of these two effects leads to a theory suitable for calculations of engine bearings. Hahn (11, 12) and Holland (13) have used two different ways to arrive at char acteristic diagrams for this calculation (see Appendix B ) ", + " For loads created by gas pressure, in general there is more favorable behavior. At the beginning of the peak pressure period, the center of the shaft very often moves in the direc tion of the rotation. As a result of the increasing load, the center then moves backward with high velocity and so c re ates a high tangential velocity which is added to rotation speed (opposite direction), thus building up a high film pres sure. Therefore a low radial movement of the shaft is nec essary to support the load. As can be seen with bearings 0 and 3 in Fig. 1 1 , the tangential speed is so high that the loadcarrying capacity is higher than the external load. Despite the increasing load, the oil film also increases. From these diagrams, it is possible to ascertain that the peak load can be supported easier the higher the center of the shaft has climbed up in the direction of rotation. On the other hand, it becomes rather dangerous if the center has almost reached a position on the top; because then there is no effect by re turning back and the center is moving in the direction of rotation (and so diminishing film pressure formation)", + " Holland (13) solves the problem in two different steps, according to the two different right sides of differential Eq. 6. For rotation of the shaft and tangential movement of the center: For radial movement of the center (squeeze effect) : For general movements, it can be seen that the boundary conditions at the same point of the bearing and at the same point cannot be reached. Therefore, two pressures with different boundary conditions are added which is impossible by physical reasons. The solutions for Eqs. 33 and 35 result in the diagrams for constant load (Fig. 1) and for radial movement (Fig. 5 ) . APPENDIX C The solution of the general Reynolds equation leads to the pressure distribution in the film, and the integration of all positive film pressures results in the external load of the bearing which can be defined by the magnitude and direction (angle \u03c6p relative to minimum film thickness) The results are given in Figs. 6 and 7. At the same time, both Eqs. 36 and 37 constitute a system of two differential equations for determining the relative eccentricity ϵ and the angle \u03b3 i f the load in magnitude So and angle \u03c6p is given" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000314_lfa.1988.24952-Figure6.1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000314_lfa.1988.24952-Figure6.1-1.png", + "caption": "Figure 6.1. Example for Explanation or Waiting more eficirnt than the paths generated by the algorithrn in Section 4. 0", + "texts": [ + " Let t l( i) be the value of t, when the above condition is L t i l satisfied. The minimum time of the motion assignment is min(max(tl(i),t~'(i)),max(t,(2),t;'(2))~ Therefore, the paths of B, and B, tha t uses the minimum time is selected as the solution. 6. Effect of W a i t i n g In our previous discussion, we assumed t h a t t he robots do not wait for each other. In some situations, however, it may be more efficient i f we allow one robot wait at the s ta r t ing point and the other robot move first. The following is a n example. Example 6.1 : In Figure 6.1, there a re two robots B, and B 2 where r2 = l . l r l and v2 = 2v1. The task of B, is to move from Pi, t o Pf, and t h a t of B, is t o move from Pi, t o Pf,. If B, moves along the pa th Pi,Pf,, the moving t ime t , of B, is 11.5s. If B, moves along pa th Pi,Pf,, the moving t ime t, of B, is 8s. Because t, > t,, B, should change its path. Therefore we obtain Path;(Pi,Pf,) = Pi ,K,Pf2 with t i = 12s. Since t, < t 2 , we should change Path, according t o the algori thm in Section 4. In t h a t case, t, would be greater than 11" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003881_j.precisioneng.2010.10.003-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003881_j.precisioneng.2010.10.003-Figure2-1.png", + "caption": "Fig. 2. Design variables in wedge-roller-traction drive.", + "texts": [], + "surrounding_texts": [ + "where is the traction coefficient and N is the contact force applied between the rollers. A large contact force is necessary to generate a large traction force, because is almost constant at 0.1. However, a large contact force causes heat to be generated by the shearing of the oil film. In large-power TD transmission, the traction oil is generally circulated for cooling. The mechanism of planetary-roller TDs is shown in Fig. 1 [7]. In a conventional planetary-roller TD, the ring and sun rollers are concentric, the diameters of the three idler rollers are the same, and all rollers can rotate in the CW and CCW directions. The contact force is generated by the shrink fit of the ring roller. When a TD is designed, the maximum transmission torque is decided according to the specifications. The torque is expressed F w a f b b t t a a h N w t a t a t t i r t d m p c b p c c s F d f r t T W t T [ s t e i w t a r w End milling was carried out with various axial depths of cut Ad, radial depths of cut Rd and feed speeds C in the ranges shown in Table 1. Other conditions are also shown in the table. The measured surface roughnesses are shown in Fig. 4. The results of the WTD spindle are better under most cutting conditions. The main reason ig. 1. Mechanism of planetary-roller-traction drives: (a) conventional type and (b) edge-roller type. s Ft\u00b7r, where r is the radius of a roller. Then the maximum contact orce is obtained by Eq. (1). The interference of the shrink fit is set to e sufficiently large to generate the maximum contact force, which ecomes excessive when the transmission torque is small, causing he generation of heat. On the other hand, the ring and sun rollers are nonconcentric in he WTD. The small and large rollers among the three idler rollers re restrained by shafts and can rotate along them. The wedge, ring nd sun rollers can move freely. The WTD is used as an increaser ere for the ring roller input and sun roller output, respectively. o contact force is generated between rollers without loads. The edge roller digs into point A as a result of the wedge action when he input ring roller rotates in the CCW, and a contact force is generted between the ring and sun rollers according to the load. Because he small and large rollers are fixed by shafts, the contact force is pplied between all the rollers, and the input torque is transmitted o the sun roller through the three idler rollers. Little heat generaion compared with the planetary-roller TD in a fixed contact force s a feature of the WTD. No contact force is applied to the wedge oller when the input ring roller rotates in the CW, and no torque is ransmitted. The WTD acts as a one-way clutch. A difficulty in the esign is supporting the ring and sun rollers while allowing their ovement. One of the applications of the WTD is as a reducer for ower-assisted bicycles [8]. The commercially available TD speed-increasing spindle uses a onventional planetary-roller TD. It is necessary to cool the jacket y circulating a coolant to limit the temperature rise. The temerature rise at the chuck is 25 K at 12,000 min\u22121 [2]. The authors onsidered that reason for the temperature rise was the constant ontact force. Thus, we propose the application of the WTD to a peed-increasing spindle. The design variables in the WTD are \u02db, \u02c7, and \u03b5, as shown in ig. 2. The ratio between the diameters of the sun and ring rollers is ecided by the speed ratio given by the specifications. \u02db is obtained rom the traction coefficient. The Hertzian pressure between the ollers changes with \u02c7 and . They should be the same to improve he durability of the rollers and the power transmission efficiency. he optimum values of \u02c7 and can be obtained by calculation. hen \u02db, \u02c7 and are decided, \u03b5 is obtained geometrically, and then he diameter of the idler rollers and their arrangement are decided. he derivation of the design variables is described in detail in Ref. 9]. The composition of the developed spindle is shown in Fig. 3. The peed-increasing ratio is 5 and the size of the shank is BT40. When he input shaft attached to the tool spindle of the MC rotates, the nd mill connected to the output shaft rotates through the ring, dler and sun rollers. The ring and sun rollers are slightly displaced hen the wedge roller digs into them by the wedge action during orque transmission. In particular, if the sun roller and output shaft re in a body, the end mill will be eccentric and its cutting accuacy may become worse. Therefore, the output shaft and sun roller ere connected by an Oldham\u2019s coupling, as shown in the figure, and the displacement was absorbed. The ring roller and input shaft are coupled by a spline, and the ring roller can move in the radial direction. Traction oil, which is a special oil for the TD, is used for ensuring traction contact and for lubricating the bearings. First, we attempted to use an oil bath to supply the oil at the traction contact and bearings, because we wanted to disuse a tube for the oil supply, which was used in a conventional TD spindle. If there is no tube, an automatic tool changer (ATC) can be used in an MC. However, an oil seal was required at the end of the output shaft, which caused a large temperature rise. Therefore, an oil mist is used instead of an oil bath. It was confirmed experimentally using a two-disk testing machine that the TD was achieved with a very small amount of traction oil [10]. The oil mist is also effective for cooling. Although an oil seal was not used, resulting in a lower temperature rise, it was necessary to use the tube. 3. Evaluation of performance The temperature rise at the chuck in the developed WTD spindle is 23 K at 15,000 min\u22121. The cooling ability of the oil mist is lower than that of a circulating coolant. However, the temperature rise is almost the same in the two cases. The feature of low heat generation in the WTD was confirmed. To evaluate the machining performance of the developed WTD spindle, cutting tests were carried out using a vertical MC and end mills. The results are compared with those obtained using a commercially available speed-increasing spindle using planetary gears. Input shaft Output shaft Spline Ring roller Large roller Sun roller Small roller Wedge roller Housing" + ] + }, + { + "image_filename": "designv11_69_0000163_iros.1998.724799-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000163_iros.1998.724799-Figure1-1.png", + "caption": "Fig. 1: A model of space robot system", + "texts": [ + " We have been studying on control problems for realising such cooperative manipulation and reported that a system consisting of two space robots having manip ulators and a floating object can be treated as a kind of distributed system [SI. In this paper, an adaptive RMRC method for handling an unknown floating object cooperatively by two space manipulators is proposed. And its validity is successfully confirmed by computer simulation. 2 System Formulation Formulation of space robot system, which consists of two free-based robots with 3-DOF manipulators and a floating object as shown in Fig. 1, and a functional partition of the robot system are shown here. 2.1 Model of Space Manipulators A space robot system shown in Fig. 1 is analyzed here. Assumptions and symbols used in this paper are as follows; [Assumptions] A1 : All elements of the space robot are rigid. A2 : The robot system is standing still at an initial state, i.e., both an initial linear momentum and an angular momentum of the space robot are zero. A3 : No external force acts on the robot system. A4 : Positions and attitude angles of the robots and an object in an inertial coordinate frame can be measured. [Symbols] E : an inertial coordinate frame Pint : vector fiom an origin to an interest position on Pf : vector from an origin to joint i of robot k in E P", + " : vector of mass center of link i of robot k in E rg : vector of mass center of total robot system in T O : vector of mass center of an object in E vo : linear velocity of an object in E 40 : angle of an object 4: : angle of link i of robot k WO : angular velocity of an object W: : angular velocity of link i of robot k mo : mass of an object m! : mass of link i of robot k w : total mass of robot system wk : total mass of robot k 10 : moment of inertia of an object about its mass an object in E E center mass center link i of robot k It : moment of inertia of link i of robot k about its kf : unit vector of indicating joint axis direction of E : unit matrix ii : a skew symmetric matrix constructed by vec- agi : a vector constructed by vector ai and ro (= Although the robot system shown in Fig. 1 consists of two robots and an object, this system can be regarded to be equivalent to a robot system having two manipulators, whose kinematic formulation has been already derived by Yoshida et al. [7]. According to their analysis result, the relation between velocity vector k of the object and angular velocity vector (ii of manipulator's joints, which is obtained fiom robot's geometrical relationships and the conservation laws of linear and angular momentum under the above assumptions, can be written as follows; tor a ai - PO) P = - J , w ; ~ H , + = J'+ (1) where T k = [kLt w;f] : velocity vector of object, angular velocity vector of manipulator's joints, r _ -I 2 3 2 3 k = l i=l 2 3 k=l i=l Jki = [o, . . . , 0 , k: x (r: - p i ) , . , kf x (r: - p i ) , O , + - . , O ] , &a = b,. . . , o , le:, . . . ,le:, 0 , . . . , o ] . A matrix J' is a Generalized Jacobi Matrix (GJM) of the system shown in Fig. 1. 2.2 System Partition By examining parameters and variables included in a matrix H a and a vector Hmq5 in Eq.(l), they can be rewritten as follows; where upper manuscripts \"O\", \"1\" and \"2\" of matrices H , , Hm denote a floating object, Robot 1 and Robot 2, respectively. Vectors q5 and q5 are joint angle vectors of Robot 1 and Robot 2. It is clear from these equations that matrix H i is constructed by the only parameter of Robot 1. Matrices H : and H : are also determined by only parameter of Robot 2 and the object, respectively, as well as Hf", + " 3 shows difference of tracking error between with and without adaptive estimation. Effectiveness of adaptive estimation is clear by this figure. These figures show that adaptive RMRC method proposed here is useful. 5 Conclusion We proposed an adaptive RMRC method for c e operatively manipulating unknown floating object by two space manipulators. The method proposed here is available, even if a number of robots having manipulator increases, under the condition that the robots handle the object in parallel form as shown in Fig. 1. Validity of the proposed method was confirmed by computer simulation. Now, experiment using a Space Manipulator Robot Testbed I1 (SMART-11) developed by us are in progress. References [l] Z. Vafa and S. Dubowsky, \u201cOn the Dynamics of Manipulators in Space Using the Virtual Manipulator Approach,\u201d Proceedings of the IEEE International Conference on Robotics and Automation, pp. 579-585, 1987. [2] Y. Umetani and K. Yoshida, \u201cResolved Motion Rate Control of Space Manipulators with Generalized Jacobian Matrix,\u201d IEEE Transactions on Robotics and Automation, Vo1" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002716_robio.2007.4522162-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002716_robio.2007.4522162-Figure2-1.png", + "caption": "Fig. 2. Specification of the T-shaped clippers of the Single hanging-point", + "texts": [ + " The mechanical system of the robot on lines could be divided into the mechanism of the robot entity, the manipulators, the auxiliary mechanism, and the housing mechanism. A mobile control station is the ground working station, which can monitor the robot by the wireless transmission system and control the robot by image data analysis. The left is the ground working station and the right is the robot entity on the lines in Fig.1. In the EHV power transmission lines environments, obstaclenavigation has tough requirements for the inspection robot from environments interference as illustrated in Fig. 2. 978-1-4244-1758-2/08/$25.00 \u00a9 2008 IEEE. 213 1) Definitions: First some symbols have to be presented. L\u2014the length of damper; M\u2014the length of the damper on the lines for mounting; D\u2014the diameter of the damper; N\u2014the length of the clipper; P\u2014the height from the top of the clipper to the tension string; e\u2014the empirical value (It is assigned a value in 5~10mm). The width of the clipper can be expressed by a maximum estimation of 100mm in width. 2) Expressions for Obstacle Space: The space of the damper and the clipper should be as followed" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002236_j.topol.2007.12.004-Figure11-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002236_j.topol.2007.12.004-Figure11-1.png", + "caption": "Fig. 11.", + "texts": [ + " , n + 2 of Jn and, hence, Jn does not contain any pair of parallel stripes. Thus Jn is incompressible for all n. Example 30. Our next example concerns the surfaces constructed by R. Qiu in [7]. These are separating incompressible surfaces Qn, n 1 of arbitrarily high genus, one boundary component and properly embedded in H2. We demonstrate how these examples can be obtained by the Main Construction and how Theorem 11 can be used to show that these surfaces are incompressible. Start with the surface Q as shown in Fig. 11(a). This surface can be obtained by the main construction by using two stripes of type \u03b11, four stripes of type \u03b21, two stripes of type \u03b20 and seven components (disks) Ki , i = 1, . . . ,7 inside the 3-ball C\u03b1,\u03b2 . We may construct a surface Q1 out of Q as shown in Fig. 11(b). In fact, this can be done by using an (not properly) embedded disk, denoted by DQ1 in order to connect two boundary components of K5. The surface Q1 has 4 stripes of type \u03b11, 8 stripes of type \u03b21 and 4 stripes of type \u03b20. Observe that the number of stripes of type \u03b11 is multiplied by 2 (and similarly for \u03b21). This is because the embedded disk DQ1 used to construct Q1 out of Q \u201ctraces\u201d all stripes of type \u03b11 and \u03b21 used to construct Q. Moreover, for the same reason, the number of components (disks) inside the 3-ball C\u03b1,\u03b2 is increased by 5. Q does not contain any pair of parallel stripes and, by construction, the same is true for Q1. As in the previous example, we may perform the same procedure with a disk DQ2 (see Fig. 11(d)) to obtain a surface Q2. With a repeated application of this attaching procedure we obtain the family Qn,n 1. As above, Theorem 11 guarantees that Qn is incompressible for all n. All surfaces Qn, n 1 above have one boundary component. The construction in the above example can be modified in order to produce separating incompressible surfaces Sn, n 1 of arbitrarily high genus properly embedded in H2 each of which has two boundary components. This can be done by starting with the surface Q0 and adding one stripe of type \u03b11 (this stripe is indicated by an arrow in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002591_icma.2007.4303753-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002591_icma.2007.4303753-Figure1-1.png", + "caption": "Fig. 1. cooperative communication with 3 nodes", + "texts": [ + " The outline of this paper is as follows. Section II provides an overview of the system model of cooperative communication. Section III investigates the communication range of virtual node when two nodes are cooperating. Section IV analyzes the transformation condition of the virtual node\u2019s 1-4244-0828-8/07/$20.00 \u00a9 2007 IEEE. 1395 communication range . Section V studies the communication range with different power allocation strategies. The paper ends with conclusions Section VI. The system model is shown in Fig.1,which denotes the communication model of two nodes are transmitting cooperatively. We make the following assumptions: 1) Using Decode-and-Forward(DF) cooperative transmit strategy 2) All path undergo Rayleigh fading 3) The system can achieve the performance of diversity order 2 4) Total transmission power is constrained 5) The coordinates of two cooperative nodes are (R, 0) and (\u2212R, 0) respectively, destination node\u2019s coordinates are (x, y) III. COMMUNICATION RANGE OF VIRTUAL NODE In Rayleigh fading environment, the relationship between average BER, transmission distance and receiver SNR can be rewritten as [5][6]: Pb(e) = 1 2 ( 1\u2212 \u221a \u03b3\u0304 1 + \u03b3\u0304 ) , (1) \u03b3\u0304 = \u03b3 \u00b7 r\u2212\u03b1 where \u03b3 is the input SNR of system, r is the distance between transmitter and receiver, \u03b1 is the channel attenuation factor which usually chooses 2 \u223c 6", + "952946 Pb(e) = 10\u22123, r = 0.532229 Pb(e) = 10\u22124, r = 0.299092 Pb(e) = 10\u22125, r = 0.168181 Pb(e) = 10\u22126, r = 0.094574 Now we consider the range of two nodes participated in cooperative transmission. Therefore, we need to establish the relationship between distance and BER firstly. In traditional analysis of BER, little attention is paid on the difference of distance between different links. With cooperative communication is introduced into WSN, we must take the difference mentioned above into account(look at r1 and r2 in Fig.1). It is necessary for us to deduce the BER performance of destination node under cooperative mode. Make use of the method based on MGF(moment-generating function)[5],the MGF of receiver SNR over Rayleigh fading channels is: Mrayleigh ( \u2212 g sin2 \u03c6 ; \u03b3 ) = ( 1 + \u03b3 sin2 \u03c6 )\u22121 ; (3) The BER is: Pb(e) = 1 \u03c0 \u222b \u03c0 2 0 ( 1 + \u03b3r\u2212\u03b1 1 sin2 \u03c6 )\u22121( 1 + \u03b3r\u2212\u03b1 2 sin2 \u03c6 )\u22121 d\u03c6 (4) where \u03b3 = Eb N0 ,r1 and r2 is distance of node 1 to destination and node 2 to destination respectively. We will require an integral of (4), written as: Pb(e) = \u23a7\u23aa\u23a8 \u23aa\u23a9 1 2 ( 1\u2212 \u221a \u03b3 ( r\u03b1 1 \u221a r\u03b1 1 +\u03b3\u2212r\u03b1 2 \u221a r\u03b1 2 +\u03b3 ) (r\u03b1 1 \u2212r\u03b1 2 ) \u221a r\u03b1 1 +\u03b3 \u221a r\u03b1 2 +\u03b3 ) (r1 = r2) 1 2 \u2212 \u221a \u03b3 (3 r\u03b1 1 +2 \u03b3) 4 (r\u03b1 1 +\u03b3) 3 2 (r1 = r2) (5) 1) r1 = r2: It is a complicated work to get the exact expression of r1 and r2 from (5)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001979_50037-5-Figure7.93-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001979_50037-5-Figure7.93-1.png", + "caption": "Figure 7.93 A quarter-symmetry model.", + "texts": [ + " The objective of the analysis is to observe the stresses in the cylinder and the rail when an external load is imposed on them. The dimensions of the model are as follows: diameter of the cylinder = 1 cm and cylinder length = 2 cm. Rail dimensions: base width = 4 cm; head width = 2 cm; head thickness = 0.5 cm; and rail height = 2 cm. Both elements are made of steel with Young\u2019s modulus = 2.1 \u00d7 109 N/m2, Poisson\u2019s ratio = 0.3 and are assumed elastic. Friction coefficient at the interface between cylinder and the rail is 0.1. For the intended analysis a quarter-symmetry model is appropriate. It is shown in Figure 7.93. The model is constructed using GUI facilities only. First, a two-dimensional (2D) model is created (using rectangles and circle as primitives). This is shown in Figure 7.94. Next, using \u201cextrude\u201d facility, areas are converted into volumes and a 3D model constructed. When carrying out Boolean operations on areas or volumes, it is convenient to have them numbered. This is done by selecting from the Utility Menu \u2192 PlotCtrls \u2192 Numbering and checking appropriate box to activate AREA (area numbers) or VOLU (volume numbers) option", + " Figure 7.104 shows the frame and selections made. It should be noted that in order to have the quarter cylinder oriented as required, [A] length of extrusion = \u22121 cm should be selected. The minus sign denotes direction of extrusion. From Utility Menu select PlotCtrls \u2192 Numbering and check in VOLU and check out AREA. This will change the system of numbering from areas to volumes. Rail is allocated number V1 and cylinder number V2. This completes the construction of a quarter-symmetry model, as shown in Figure 7.93. Before any analysis is attempted, it is necessary to define properties of the material to be used. From ANSYS Main Menu select Preferences. The frame in Figure 7.105 is produced. As shown in Figure 7.105, [A] Structural was the option selected. From ANSYS Main Menu select Preprocessor \u2192 Material Props \u2192 Material Models. Double click Structural \u2192 Linear \u2192 Elastic \u2192 Isotropic. The frame shown in Figure 7.106 appears. Enter [A] EX = 2.1 \u00d7 109 for Young\u2019s modulus and [B] PRXY = 0.3 for Poisson\u2019s ratio" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000215_icsmc.1995.538489-Figure7-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000215_icsmc.1995.538489-Figure7-1.png", + "caption": "Figure 7: Simulation results of the PI method for Example 1.", + "texts": [ + " The translational motion planning is defined in Fig. 6 with motion in x and y directions of Base coordinate, XB and Y E , only. The orientation of the End-effector is planned to be fixed through the whole trajectory. It is noted that it takes 10 seconds to finish the whole motion and the desired trajectory passes through the interior singular point at t=5 seconds. In this example, ko=100 and ~ i = 3 . The simulation results are shown in Figs. 7 to 10. The restricted region lies within the vertical lines at t=4.68 and 5.40 seconds. Referring to Fig. 7, it is obvious that the PI method will produce discontinuous solutions of 81 when the End-effector is passing through the singular point. The discontinuity of 81 will cause abrupt change of the End-effector. As for (22). In accordance with W.1 and W.2 selecte h in the SRI the results of the SRI method, the solution curve of 04 revealrr a lump within the restricted region. (See Fig. 8 small lump can also be found in the solution curve o 6'2. Therefore, the End-effector will deviate when it is passing through the singular point" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000197_6.2002-4081-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000197_6.2002-4081-Figure1-1.png", + "caption": "Fig. 1 Schematics of example rotor. Nondimensional parameters: k1 = 1 (tuned disk), k2 = 1.1, k3 = 493, m1 = 1, m2 = 426.", + "texts": [ + " LFTLB stops when the cost function makes no further improvement for a pre-specified number of consecutive iterations. The calculation of the optimal step-size can be done in several ways. Note from (22) that \u03b1o is a maximizer of the largest singular value of a rational matrix function of \u03b1. The method to calculate \u03b1o is reminiscent of the methods used to calculate infinity norms of rational transfer functions.33 In this method \u03b1o is determined by solving a sequence of eigenvalue problems. Details of the method are in the full version28 of this conference paper. The rotor example (see Fig. 1) is a mass-spring model with n = 56 sectors. There are two degrees of freedom (DOF) per sector, one DOF represents the blade motion and the other DOF the disk motion. All model parameters are known except for the blade-alone stiffness which, for the i-th blade, is modeled as k1i = k1(1 + \u03b4i) (23) 5 of 11 American Institute of Aeronautics and Astronautics where k1 is the nominal (tuned) blade-alone stiffness and \u03b4i is the unknown but bounded stiffness perturbation or mistuning. The equation for this model is of the form shown in (1)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002719_pesc.2008.4592623-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002719_pesc.2008.4592623-Figure2-1.png", + "caption": "Figure 2. Phasor diagram of PM synchronous motor", + "texts": [ + " The voltage equation in the synchronously rotating reference frame can be obtained from the voltage equation in (1) using the relationship as follows [2]: 2 2 cos cos cos 3 32 3 2 2 sin sin sin 3 3 ase e e qs bs ds e e e cs v v v v v (3) and the result becomes 0 S S S e qs qs m e ds ds S e S S d R L Lv idt v id L R L dt (4) where vqs and vds are the q- and d-axis voltages, iqs and ids are the q- and d-axis currents, and e is the electrical angular velocity of the rotor. III. FLUX OBSERVER AND POSITION ESTIMATION The phasor diagram of a PM synchronous motor is shown in Fig. 2, the current vector is and the voltage vector vs being derived from the measured phase current and voltage, and the transformation relationship from the abc reference frame to the stationary reference frame. The vector s represents the resultant stator flux linkage, while the vector m represents the excitation flux linkage due to the permanent magnets, which is in phase with the rotor d-axis. The angle e represents the rotor position angle which is the angle of q-axis. From the (1), s is observed as follows: 0 0 t s s S s sv R i d (5) with zero current in the stator windings, the stator flux linkage vector is simply the excitation flux linkage vector, s(0) = m, which is obtained by initially aligning the rotor before the flux observer is applied. From Fig. 2, it is shown that m can be calculated from the stator flux linkage vector s, the current vector is, and winding inductance. Further, if surface-mounted type PM synchronous motor is used, saliency can be neglected, Ld Lq LS, and m is simply calculated as follows: m s S sL i . (6) In the reference frame, m is expressed by the projections of m and m on the - and - axes as shown in Fig. 2. Therefore, the rotor position is obtained as follows: 1tan m e m . (7) The estimation of the stator flux linkage vector s according to (5) requires performing the integration. The offsets contained in the integrator input make its output drift away beyond limits as shown in Fig. 3(a), because the integrator has an infinite gain at zero frequency. Therefore, the integrator amplifies any dc offset or error in the voltages, currents, and resistances until saturation is reached [3]. Clearly, it is not possible to deduce the rotor position from such a flux vector" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002979_978-3-540-70534-5_7-Figure17.2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002979_978-3-540-70534-5_7-Figure17.2-1.png", + "caption": "Figure 17.2: Micromouse generations, Univ. Kaiserslautern [Hinkel 1987]", + "texts": [ + " All of a sudden, people had a goal and could share ideas with a large number of colleagues who were working on exactly the same problem. Micromouse technology evolved quite a bit over time, as did the running time. A typical sensor arrangement was to use three sensors to detect any walls in front, to the left, and to the right of the mouse. Early mice used simple Maze Exploration Algorithms 273 micro-switches as touch sensors, while later on sonar, infrared, or even optical sensors [Hinkel 1987] became popular (Figure 17.2). While the mouse\u2019s size is restricted by the maze\u2019s wall distance, smaller and especially lighter mice have the advantage of higher acceleration/deceleration and therefore higher speed. Even smaller mice became able to drive in a straight diagonal line instead of going through a sequence of left/right turns, which exist in most mazes. One of today\u2019s fastest mice comes from the University of Queensland, Australia (see Figure 17.3 \u2013 the Micro Mouse Contest has survived until today!), using three extended arms with several infrared sensors each for reliable wall distance measurement" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000914_robot.2004.1302473-Figure3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000914_robot.2004.1302473-Figure3-1.png", + "caption": "Fig. 3. The ideal beam sensor model without sensory noise.", + "texts": [ + " The expected information gain (IG) is formulatcd as I G c ( s ) = - E { A H ( C ) ) . 9 where E denotes the expectation operation. In [8], explicit closed form expressions were derived for IGc ( s ) for an ideal beam sensor, assuming a Poisson point process model [I 11 for obstacles in the physical space3 and ignoring mutual information terms (for simplicity of computations). A beam sensor is characterized by a sensing ray of length L starting from the sensor origin. it gives the distance of the closest obstacle point (called hit point) along the beam. Fig. 3 shows the beam sensor model. Note that it has a zero volume field of view (FoV), and hence we compute the information gain density (IGD) [4], [6], [8]. The final expression of IGD, the approximation of IGD that omits mutual information, for the ideal beam sensor is given by I F D ~ = igdq(s) = 2. len(A(9)nv,(s)).log(l-p(q)) (2) where igd,(s) describes the information gain density due to a configuration q; V(s ) denotes the sensing beam; and VZL(s) is the unknown part of V ( s ) that is in front of the first known obstacle along the sensing direction" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001563_095440605x8397-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001563_095440605x8397-Figure1-1.png", + "caption": "Fig. 1 Surface geometry of the special rack cutter", + "texts": [ + " In the present study, an innovative modified spur gear with crowned teeth and its generating mechanism, which can manufacture the surface of one gear tooth in only one working procedure, are proposed. Mathematical models of surface design and tooth contact analysis of the gears have been developed on the basis of gearing theory and computer programs established according to derivedmodels. Consequently, numerical examples are intended to illustrate generated tooth profiles, the results of tooth contact analysis, and potential transmission errors caused by assembly error. As shown in Fig. 1a, if the gear blank is rotated with an angular velocity v about its axis A\u2013A, and the special rack cutter is translated with a linear velocity of rgv, where rg is the pitch radius of the gear, then the modified crowned spur gear can be generated by a special rack cutter. As demonstrated in Fig. 1b, a special rack cutter tooth is generated by rotating the section profile of an ordinary involute rack (as illustrated in Fig. 1c) about axis B\u2013B, where rc is the distance between axes B\u2013B and Xr, and W is the tooth face width. Each radial section of the special rack cutter tooth, such as the section of plane SR passing through axis B\u2013B, as shown in Fig. 1b, has a straight-edged profile as in Fig. 1c. As to those vertical sections of the special rack cutter (such as the section of plane SP in Fig. 1b), both sides of their profiles are concave curves (Fig. 1d). It can also be noted that vertical sections have different dimensions; those closer to the lateral edges of the special rack cutter are taller. Both lateral sides of the special rack cutter, therefore, exhibit maximum height, while the middle one of plane SQ, in Fig. 1b, also a radial section, exhibits the minimum height. In practice, the special rack cutter can be replaced by a ring gear cutter. The outline of the manufacturingmechanism for generating themodified spur gear is illustrated in Fig. 2, where A\u2013A is the rotational axis of the gear blank and B\u2013B is the rotational axis of the ring gear cutter. Here, rc and vc serve as the pitch Proc. IMechE Vol. 219 Part C: J. Mechanical Engineering Science C11604 # IMechE 2005 at RMIT UNIVERSITY on July 12, 2015pic.sagepub.comDownloaded from radius of the cutter and its angular speed respectively. The profile of the radial section (K\u2013K section) of the cutting tool is identical to the straight-edged rack cutter in Fig. 1c which generates the tooth profile of involute gears. As presented in Fig. 1a, the special rack cutter can be used as the generating cutter. Based on its geometry, the mathematical model for surface generation is developed as articulated below. As shown in Fig. 1c, the profile of the radial section of the special rack cutter includes two parts: the straight-line segment generates the working portion of the gear tooth surface, while the other, the circular arc, produces the fillet of the gear surface. Here, d denotes the half tooth thickness on the pitch line, u is a surface parameter, a denotes the pressure angle, a is the addendum, m represents the module, and Crd, rrd, and ard become the centre, radius, and central angle of the circular arc. The position vector of a point on the straight-line segment with respect to coordinate system Sr(XrYrZr) can be expressed as R(0) r \u00bc + d \u00fe u tana\u00f0 \u00de u 0 2 4 3 5 (1) where superscript (0) represents the generating surface S0, and the upper (plus) sign and lower (minus) sign represent the right-side and left-side profiles of the radial-section of the special rack cutter (Fig. 1c) respectively. C11604 # IMechE 2005 Proc. IMechE Vol. 219 Part C: J. Mechanical Engineering Science at RMIT UNIVERSITY on July 12, 2015pic.sagepub.comDownloaded from As reflected in Fig. 1b, the surfaces of the special rack cutter can be acquired by revolving the profile in Fig. 1c about axis B\u2013B with a radius of rc. The surfaces of the special rack cutter with respect to coordinate system Sc(XcYcZc) can then be represented by R(0) c \u00bc + d \u00fe u tana\u00f0 \u00de u\u00fe rc\u00f0 \u00de cosb rc u\u00fe rc\u00f0 \u00de sinb 2 4 3 5 (2) where b represents the revolving angle, shown in Fig. 3a, and is also a surface parameter of the special rack cutter. The normal vector pointing to the inward side of generating surface S0 can be determined with N(0) c \u00bc @R(0) c @u @R(0) c @b \u00bc + rc \u00fe u\u00f0 \u00de (rc \u00fe u) cosb tana (rc \u00fe u) sinb tana 2 4 3 5 (3) Thus, the unit normal vector is n(0)c \u00bc + cosa cosb sina sinb sina 2 4 3 5 (4) To derive mathematical models for the surface generation and curvature analysis, coordinate systems are defined in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001967_978-1-4302-0088-8-Figure3-5-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001967_978-1-4302-0088-8-Figure3-5-1.png", + "caption": "Figure 3-5. The differential drive of a CubeBot", + "texts": [ + "java: Super-class used by the MiniSSC and other classes for servo communication \u2022 MiniSSC.java: Implementation class for the Scott Edwards MiniSSC-II Now that you know how to control servos with your PC, you\u2019re ready to get a robot to move. In the next section, I\u2019ll talk about differential drive robots (with two wheels) and there I\u2019ll use the MiniSSC and your PC\u2019s serial port to make it move. Using a servo controller connected to an electronic speed controller or a pair of \u201chacked\u201d or continuous rotation servos is an excellent way to facilitate wheeled motion. Figure 3-5 shows a picture of the differential drive of a CubeBot connected to a MiniSSC-II. Notice that the servo wires are to the rear of the platform. This means the motors are inverted, so I\u2019ll have to account for this in the classes in this section. 60 C H A P T E R 3 \u25a0 M O T I O N Three classes and one interface will be discussed in this section. Figure 3-6 shows a class diagram that summarizes the classes. C H A P T E R 3 \u25a0 M O T I O N 61 The objective in this section is to provide basic movements: forward, reverse, pivotRight, pivotLeft, and stop" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001336_iros.2006.281764-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001336_iros.2006.281764-Figure1-1.png", + "caption": "Fig. 1. (a) Tire-Roadl syst4tii: (b): Tire-Road Systeml Word Bowl(I Graph", + "texts": [], + "surrounding_texts": [ + "A. System Descript;ion In this part, we limited ourselves to a tire-road system, made up of six parts: the gum. the tube tire, the wave front, the wheel, the environment and the roadway as it is shown in Figure (1-a). In the following modelling development, only lonigitudinal and uormal dyuamics are taken in corisid eration for mechanical aspect due to the tire-road contact. Thermal power exchange is also modelled betwveen the gum aud the tire tube and betwveen the enviroument and the tire tube. Finally, pneumatic and hyldrodynaamnic phenomenon are described successively by the pressure variation inside the tire tuLibe auid hydrodynamics force generated by the waw front. [11: the longitudinal qarter of vehicle velocity 'Land the linear tire velocity r.w, with r the tire radiuis and considered constant due to the tire rigidity and w the angular velocity of tire. X rwx t, (1) Figure (2-a) shows the longitudinal tire gum behavior when the wheel makes an angular and translation motions after contact with the roadway. At the contact level, the generated longitudinal effort Fa is decomposed into three forces: inertial torce due to mass M, elastic force from the spring k and viscous friction force through the resistance RS. The resistance element RS is used in bond graph theory to model the active resistance which generate the entropy flow from mechanical friction ect. This irreversible transformation fomn echanical to theriral power provide the thermal flow Q to the tire tube. B. Wod Bond Grah The word bond graph represents the technological level of the model where global systern is decomiposed into six subsystems (see Figure (1 -b)). Comparing to classical block diagiram, the imrput and outpuLit of each subsystems define a power variables represented by a conjlugated pair of effort-flow labelled by a half arrow. Power variables used for the studied system are: (Force, linear velocity) (F x), (Tempernture, ThermaiFlow) (T, Q), (Pressure, Volume Flow) = (P. V) (Torque. Angnlar velocity) = (r, ). These true and psetudo bond graph variables are associated respectiviely with mechanical translation and hrydrodynamic, thermal, pneumatic, amrd mechanical rotation. C Bond Craph Modes In this subsection, dynamic bond graph models of each subsystern descriibed in Figure (1-b) are presented. C. l Tire Gum C.iLa Lonigitudinal tire gum behavior bond graph. Tire gum is considered as a viscoelastic muaterial, xhich is deforming with a behavior located between a viscous liquid and an elastic solid. Vhen the tire is in coitact with tire road, kinemnatically, one definles a slip velocity Qi, This latter is the difference between two eollinear velocities at the center of tire contact Bond graph model of this subsystem is given in Figure (2-b). Thus, the corresponding bonid graph which reflects the viscoelastic phenomenon of the tire-road contact is developed in Figure (2-b). In our case, the known and measurable inputs are respectively the vehicle velocity xLand the angular wheel velocity w. They are represented by a flow sources (SF: XL) and (SF c). Then, according to the viscoelastic characteristic of the gum, tranmsmitted mechanical power (in the case where the tire is in contact with the roadway) is decomposed in to two parts: The first one is transformed into kinetic through inertia of mass M modelled by I element and which describe the dynamic of all tire points outside of physical contact. The second part and due to the contact conifiguration, the slip velocity xi5is transformed into friction (generating a thermal powem RS element) and into elasticity modelled by a storage element C of value kL The longitudinal effort FI is estimated using bond graph I element in derivative causality. This causality conflict introduces implicit equation in numerical simulation. This is due of presence of imposed flow source (SF : XL). The slip velocity i is calculated by 0 junction and modulated transformer used to transfer anguIla-r velocity c to linear one. The deduced equation is givle in (1) The mechanical equation of the tire longitudinal motion, referring to the c;ontact suiface cotuld be synthesis from the bond graph scheme and given as follows: from 1 junction associated with C arid RS elements, we have: Es-FF S-.kx, + RSt5s friction is represented by a dissipative bond graph element (R: RN) while elasticity is represented by an element C of elasticity kT which store a pote-ntial energy. The qnarter vehicle inertia due to its mass M is represented by conservative element (I: M). (2) from 1 junction and I element in derivative causality, equation (3) is deduced: F-1 ALTL RS.x5 ka, (3) xvlere x, describes thre slip displacemernt, S6L the lorgitudinal acceleration of the vehicle and M the quarter vehicle mass. To avoid a derivative causality which introduce implicit equation in numerical simulation is introduced a pad in between (SF: XL) and (I M) element [3]. This pad is given as a rigid stiff spring-damper combination with a big value of stiffniess and friction coefficient (ft6 1 00, ke = oc). Let's consider that during a dynamnic slip motion of the wxheel, the points of the tire which do not beloing to the contact are nlot very deformed than those of contact, so it can be the physical meaning of the added pad. The bonid graph model of the longitudinal tire gum behavior incluLding the new pad is given iiin Figure (3). This last bond graph is in integral causality. Let's prove that flow f4 of element (I : M) is equal to c\u00a3Lafter adding the new pad. According to afected causality, the thermal power flow Q generated by the actix e RS element is given by equation (4) and detailed in [5]. RS is the resistance value antd a a parameter depending on the slide angle. 2Q RS.W a (4) C.1.b Normal tire guru behavior bond graph. The guLru is a viscoelastic material which can be described by (damper spring and niass) systeo of Figure (4-a). The input variable corresponds to the normal effort Fy deduteed from the pressure variationi in the tire tulbe. The viscous normal In this case, the dynamic element (I: M) is in itAegral causality, because the effort FN is known. That is why the normal mass elocity ktcan be deduced by integration. (R. Rv) is a dissipatiw f element with aniy thermal power generation. The corresponding bond graph model is given in Figure (4-b) The following mechanical equation (5) is deduced from bond graph of Figure (4-b) M.i1 + RN.( 1-2) + K.(xj X2) FN (5) x1 and zi are the vertical guim displacement alnd vertical gum velocity due to the load anid x2 and x2 are the vertical gum displacement and vertical gum velocity due to the roadway profile. C.2 Wave Front and hydrodynamic effect Two kinds of forces interaction can generate a wave front phenomena: internal forces interaction between fluid mnolecules and externld forces interaction between fluid mole cules and those of the tire circumference (see Figure (5-a)). Each fluid molecule does not ruLn out at the same vlocity [71 ard follows a velocity profile of Figure (5-a). Bond graph model of hydrodynamic phenomenon is given in Figure (5-b). When each particle located in a cross-section perpendicular to the overall flow is represented by a velocity vPector, the obtained curve from the vectors extremities represents the velocity profile of Figure (5-a). The movement of the fluid can be regarded as resulting from the slip of the fluid layers the ones on the others. The velocity of each layer is a function of distance, h of this curve in the fixed plan: Xf XLf (h). Let us consider two distant contiguous layers of fluid of dh. The friction force F. which is exerted on separation surface of these two lyers, opposes to the slip of a layer on the other. Thus, the force F. is proportional to the difference in layers velocity dif, on their surfice so and invwrsely proportionlAl to dh: Let's consider that distribution of pressure during the contact tire-road is represented by an ellipsoidal function of longitudinal and lateral positions in 3D [4], as it is shown in Figure (6). From (9), FN could be expressed as follows: 1 fj(A y) PoJf 1 a2 b2 (10) After development of equation (10), the following expression of fry is deduced:h+ dh (a) (b) FNt -. (i (1 12)3) with I describes the contact surfaice lenigth. iwhere describes the coefficient of dynamic viscosity. our case, we take the difference in layers velocity dx equal to the velocity slip x5, (Z.e velocity of the first layer is Xf 0 and for the second irf in the slip su-rfice for water height level of h). So, F1; cau be written as: Figz. 6. Pineuimiatic and tlietinical effedts in tie tube = .80. Xs h This transformation of the slip velocity to the lhdrodyniamlicfeorce FR can be represen;ted in a bond graph model of Figure (5-b), by a modulated gyrator elemient GY with the constant as a modulus. C.3 Tire tube Tire ttibe is a synthetic ruibber sheet located inside the tire and describes an elastic enclosure containing a perfect gas under pressure. The tire tube and the correspondinlg bond graph model are given in Figures (6) and (7). Tire tube exchanges the thermal power with the external envi- ronment dLue to the heat conduLctivity of its wall, anid stores two types of eneirgies, pneumatic and thernal. The thermnal exchange is modelled by the dissipation element R and the storage phenomenon is represented by the two port C element. Durinig the slip stage, the temLperature \\ariation acts on the iinternal pressure inside the tire, tube according to the perfect gas equation (8) FU.V n.,ro.T (8) with: P the gas pressure inside the tire trrbe, V the gas vTolume of the tilre tuibe, n is moles number anid T the temperature iniside the tire tube. Knowing that normal effort Fr is proportional to contact pressure P and con:tact surface A as given by equation (9). = P.A (9) with: (x, y) are the two dimensiunus contact sllrface a11d a, b are constants. Thus the bond graph model of the tire tube is given by Figure (7). Resistance (R : R1) represents the heat con- ductivity of the tire tube walls. It ensures the heat transfer with the external environment. Input model corresponds to the heat flow originally from thi frictioni of gutm Q, and the output one represents the normal effort FN5, obtained by ellipsoidal distribution of pressure due to transformer TF. The two ports C element describes entropy S and xoluume V variations inside the tire tuhoe and is detailed in The thermal flow exchanged between environment and tire tube is given by following equation (see Figuire (7)): 1 Q0- 07R 7rI N Ri (12) where: Qo exchanged heat flow between environment and tire tube, T and are respectively tire tube and en- vironnment temperatures. and the global inlet therrmal flow Q1 to the tire is calcu- lated from 0 junction: (11) Q1 = Q0 + Q (13) thus, for the inilet flow entropy S to the tire, it is calen lated by tIhe transformer TF according to Car not equation as follows: s = T (14) this transformer is added just to transform the thermal flow Q, to entropy flow S. We note that the constituitive equation of the last transformer concerns onlly the relationship between flows (x.e. temperature is not transirmed). D. Global system modelling Global Tire-road bond graph mnodel is given in Figgure (8), where interconnection appears clearly between all subsystems represented in word bond graph. For this global model, one needs two principal measurements: linear velocity of thexwhicle L, arid angular xheel velocity w. The external temperature can be added as pas rameter online or offhine and the roadway profile can be considered as external input. So, tire guLr xertical deformation xr x can be estimated through the normal eflort TFx variation dtre to variation of presstrre P inside the tire ttmbe.IThis pressure variation is caused by viation of teniperature T of the tire tube which is generated front transformation to heat flow Q of the tire-raod longitudinal friction force RS.X. In order to represent more the water presence oni roadway, waxe front subsystem car be optionally added to the global nmodel as shown in Figure (8).- III. SIMUIoATIN TESTS A. SoftwomjeIplementation Simulation step is done on a specific bond graph software SMIBOLS 2000, which is an object oriented hierarchical modelling. It alloxs users to create models using bond graph, block-diagram and equation inodels. Differential causaities and algebraic loops are solved ouLt using its powerful symbolic soltution engine. Nonlinearities and user code can be integrated in single editing IDE (integrated development environment). The iconic modelling facility allows system-morphic model layout. It also has manW post-processing facilities over the simulated result. Thanks to a developed genieric item database which consists of a set of predefined models, and has been incorporated as capsules in the software SYMBOLS 2000, the designer can easily build the dynamic models of sewvral transportation systems from the Process and Instrumentation Diagram (P&ID) just connecting different sub models. The global dynamic in syibolic format is obtained connecting diffrent icons. Behind each submodel the bond graph model is hinted. If parameter values are available, the model can be simulated using own Symbols function or Matlab-Simulink (the bond graph model can generate S -Matlab function). B. Simulation results Simulatiou parameters are: Po 2.iO (Pnscnl), V ,0,3 (iM3) To 45 (oC), 1 0, 1 (m), r = 0,25 (at), KN k -6 (N mt), M 500 (k), RS - 12 (N.s/at), RN = 9 (N.s/rat) so 0, 1 (aMt2), R1 = 0, 23 (Wqattsam.0K), A 0, 3 (mt2) Wave frotA phenomenon is present at the tire-road conAact level (see Figure (9-(a))), where the grip mechanism depends inversely on the water height. One notices a fill of effort value from 0, 63kN to 0, 58kN for a water height of 0,3mm. By increasing the water height on the roadway, the grip is degraded and the effort reached a value of 0, 52kN for 2ma/s of slip velocity and 1, 5mm of water height. In this phase, the micro-indenters are flooded, and only the macro ones continue to operate. On a wet roadway, water viscosity increases when temperature decreases. Thus, the ground becomes more slipping, and the grip potential is degraded. The maximum effort value goes from 0, i68kV to 0, 58kW when water viscosity varies from 0, 5. 10 3Pas to 3.10-3Pals with the same slip speed profile (see Figure (9 (b))). The curves of Figure (9- (c)) are obtained for an initial value of temperature inside the tire tube of 450C, and two values of ambient temperature, 150C arid 30 C. The shift between the two curves shows the influence of the ambient temperature on temperature variation inside the tire tube. For the case of 150C and for slip velocity profile of Figure (9-(c)), a temperature decreasing from 45 C to 44.92 C is noticed after a braking period where the tire ttube releases heat to the external environment because its temperature is higher than outside temperature. The temperature increases gradually in the tire tulbe during the phase of increasing slip velocity. Pressure variation is proportional to the variation in the temperature. by comparing the curve of Figure (9-(c)) with that of Figure (9-(d)). OnLe notices that pressure decreases with temperature, which occurs indeed for values of slip going from lam/s to 0.2an/s, when pressure varies from 2Bar to 1.68Bar and temperaturte goes from 450C to 44.920C). Then a progressive increase of pressure is observd when slip and temperature increase." + ] + }, + { + "image_filename": "designv11_69_0001148_50002-1-Figure1.3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001148_50002-1-Figure1.3-1.png", + "caption": "Figure 1.3 (a) npn bipolar transistor; (b) pnp bipolar transistor.", + "texts": [ + " The thin and lightly doped central region is known as the base (B) and has majority charge carriers of opposite polarity to those in the surrounding material. The two outer regions are known as the emitter (E) and the collector (C). Under the proper operating conditions the emitter will emit or inject majority charge carriers into the base region, and because the base is very thin, most will ultimately reach the collector. The emitter is highly doped to reduce resistance. The collector is lightly doped to reduce the junction capacitance of the collector-base junction. The schematic circuit symbols for bipolar transistors are shown in Figure 1.3. The arrows on the emitter indicate the current direction, where IE = IB + IC. The collector is usually at a higher voltage than the emitter. The emitter-base junction is forward biased while the collector-base junction is reversed biased. 1.2.2 Digital control 1.2.2.1 Transfer function A transfer function defines the relationship between the inputs to a system and its outputs. The transfer function is typically written in the frequency (or s) domain, rather than the time domain. The Laplace transform is used to map the time domain representation into the frequency domain representation" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003893_iceee.2010.5661528-Figure2.1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003893_iceee.2010.5661528-Figure2.1-1.png", + "caption": "Figure 2.1 The model of stator and winding", + "texts": [ + " Therefore, firstly, the three dimensions electromagnetic field of AFSMPMSM is calculated, and the component of its harmonic wave is also analyzed, and the number of low step electromagnetic force wave is found. The electromagnet field of a 5kW AFSMPMSM is calculated by using FEM. And the mark of magnetic force is imposed on the stator to calculate the electromagnetic force. This motor\u2019s structure is double stators and single rotor. Because the magnet path of this motor is symmetrical, so the model is one half of this motor for simplifying calculation. The three dimensional model of stator, winding, and permanent magnet (PM) is shown in Fig.2.1. And the winding element figure which is This work was supported in part by Outstanding Young Teachers Project of Shanghai Municipal Education Commission(sdj09014 ,sdj09A111) 978-1-4244-7161-4/10/$26.00 \u00a92010 IEEE imposed the electric current density is showed in Fig.2.2. And the magnet density distribution under the same pole and different radius is shown in Fig.2.3. And the magnet density circumference distribution on the stator teeth is also showed in Fig.2.4. The number and magnitude of basic wave and some harmonic waves are shown in Table ,which change alone with time and position in period, and which cause the vibration and noise of motor" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002508_ijtc2007-44479-Figure5-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002508_ijtc2007-44479-Figure5-1.png", + "caption": "Figure 5: Stress distribution in the pad structure for 10% of clearance in upward direction for L/D=0.5 and eccentricity ratio =0.6", + "texts": [], + "surrounding_texts": [ + "EHL analysis of radialy adjustable partial arc bearing is proposed which includes deformation of the bearing surface. Simulation results show that the level of the structural deformation depends on the radial adjustments given in addition to the bearing clearance and the lubricant film thickness. The plots and figures above show that the results agree in general." + ] + }, + { + "image_filename": "designv11_69_0002456_1.25389-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002456_1.25389-Figure2-1.png", + "caption": "Fig. 2 Universal joint. The potentiometer that detects the rotation is visible.", + "texts": [ + ", in vertical wind tunnels) would be needed. However, for the purpose of investigating the possible chaotic nature of spin, which requires two independent spinmotions, the present configuration with only 3 degrees of freedom is expected to be satisfactory. Figure 1 reports a picture of the wind-tunnel setup. A tripod is set in the divergent so as to anchor the sting to the wind tunnel. Themodel had to be connected to the sting by a universal joint and a hinge around the yaw axis to allow the three rotations \u2019, , . The joint is shown in Fig. 2 along with the degrees of freedom ( denotes the yaw rotation). The model is anchored to the universal joint through four screws (see the holes on the picture) so that its lower side would be visible on Fig. 2 if bolted. Figure 3 reports the reference frame with the goal of showing that \u2019 is the rotation of plane with respect to plane , where is fixed to the ground reference frame. is the rotation of plane with respect to plane , and is the yaw rotation (i.e., in the direction normal to plane ). The main limitation of such a joint is that rotations \u2019 and are not complete. This was not a problem in the present experiments because the maximum angles for \u2019 and were never reached in the tests, as D ow nl oa de d by U N IV E R SI D A D D E S E V IL L A o n Fe br ua ry 2 3, 2 01 5 | h ttp :// ar c", + " In fact, the size of the joint would have fixed any other model dimension because the model had to be realized to contain the whole mechanism for the detection of the rotations [29]. Knowing that frictional damping had given many problems in windtunnel tests for wing rocking, ball bearings were the first choice. The smallest size of the standard commercial bearings, therefore, provided the overall joint dimensions. For the detection of themotion many systems were considered, i.e., magnetic, resistive, and photoelectrical devices. For the sake of simplicity and for their reduced size, small stopless potentiometers (visible in Fig. 2) were chosen. B. Model Design Themodelwas designed following the idea that a chaotic behavior is theoretically possible if there are two independent spin motions [30]. For our experiments, the first motion was assumed to be governed by thewing stall [31] and the second one by the asymmetry of the vortices originating from the nose of the forebody [32]. The leading-edge stall of the wing was ensured by the very bad stall properties of the five-digit NACA23012 airfoil. The asymmetry of the rollmoment induced by thewing stall provides autorotation for the initiation of spin [33]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002104_s10832-007-9133-3-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002104_s10832-007-9133-3-Figure1-1.png", + "caption": "Fig. 1 (a) Structure of linear stepping ultrasonic motor. (b) Prototype of linear stepping ultrasonic motor", + "texts": [ + " The standing-wave type self-correct stepping ultrasonic motors are based on standing wave driving and selfcorrection positioning. It is necessary to ensure the rotor (or slider) to move into the self-correction area of goal position each step. The complete open-loop control could result in miss steps. In this paper, the new type linear stepping ultrasonic motor, based on vibrators alternately driving and slider grooves positioning, is capable of bi-direction stepping movements without accumulative position errors driven by an open-loop control circuit. The linear stepping ultrasonic motor is shown as the Fig. 1, whose stator is composed of 2 \u039b-shape vibrators bonded on the base of the motor and whose slider is a bar rectangular section with grooves uniformly distributed along the length in a side surface. The ends of two mutually perpendicular legs of each \u039b-shape vibrator respectively fix piezoelectric ceramics. The \u039b-shape vibrator, whose dimension is shown as the Fig. 2, were made from bronze with elastic modulus 1.1\u00d7 1011 N/m2, density 8,800 kg/m3, and Poisson ratio 0.33. Each piezoelement made from PZT8 has dimension 8\u00d7 3" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000976_cp:20040396-FigureI-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000976_cp:20040396-FigureI-1.png", + "caption": "Fig. I Typical flux distributions for pole-arc to pole-pitch ratio =0.7", + "texts": [ + " Printed and published by the IEE, Michael Faraday House, Six Hills Way, Stevenage, SGI 2AY The resultant cogging torque comprises components associated with each rotor magnet and stator slot. The superposition method, which was originally proposed in [2], is evaluated by determining the contribution of each magnet in the 4-pole, 6-slot motor to the resultant cogging torque, for pole-arc to pole-pitch ratios of 0.7 and 1 .O. The cogging torque waveform which results when the motoi is equipped with a single magnet having a pole-arc to pole. pitch ratio =0.7, Fig. la, as calculated by the Maxwell stress method, is shown in Fig. 2a. The contribution of a second magnet, Fig. Ib, is then simply obtained by phase shifting the cogging torque waveform which was deduced for the single magnet by 90\" mechanical, the resultant cogging torque waveform being obtained by summing the two waveforms, Fig. 2a. This approach can similarly he applied when the motor is equipped with the 4-pole rotor, as shown in Fig. IC, the cogging torque deduced from the finite element analysis of the 4-pole motor also being shown in Fig.2~. As can be seen, excellent agreement is achieved. From the foregoing, it would appear that the superposition method which was proposed in [2] is of high accuracy. However, in general, the resultant cogging torque waveform cannot be obtained in this way. This is highlighted in Figs. 3 and 4, in which the same procedure has been applied to deduce the cogging torque when the magnet pole-arc to polepitch ratio is 1" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001213_msf.503-504.179-Figure4-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001213_msf.503-504.179-Figure4-1.png", + "caption": "Fig. 4: Schematic representation of 3D-ECAP die inserts for 90\u00b0 (a) and 120\u00b0 channel (b).", + "texts": [ + " This would complicate the shape of a multi-faceted die insert used for realising this process in practice. All rights reserved. No part of contents of this paper may be reproduced or transmitted in any form or by any means without the written permission of Trans Tech Publications, www.ttp.net. (ID: 129.93.16.3, University of Nebraska-Lincoln, Lincoln, USA-08/04/15,01:56:11) The channel configurations illustrated in Fig.1b and Fig. 3 were realised in tooling, which comprised a multi-segment die insert and prestressing outer rings. Fig. 4 shows the schematic views of the die inserts used in both cases. To reduce the size of die inserts, the inlet of the channel was out of axis. This required appropriate measures to be taken to compensate for eccentric loads. The amount of equivalent plastic strain generated in the material passing one turn of an equal channel depends on the half angle \u03d5 between channel passages [7]. 3/cot2 \u03d5=\u03b5 (1) For more turns, this value has to be multiplied by the number of turns (Table 1). Thus for 90\u00b0 channel, the equivalent plastic strain produced due to one turn of the channel is \u03b5=1" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002229_ssp.141-143.55-Figure3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002229_ssp.141-143.55-Figure3-1.png", + "caption": "Fig. 3: Tool design for the production of a \u201csecurity flange\u201d", + "texts": [ + " Thus, the correlation of the parameter temperature and the liquid fraction is of determining importance for the respective material. The liquid fraction phase and/or remaining solid phase can be determined with the aid of a Differential Thermo Analysis (DTA). Measurement results of such an analysis for TiAl6V4 alloy is shown in Figure 2. Because of its simple and robust layout combined with filigree pins, the workpiece shape \"security flange\" seems ideal for preliminary tests where the forming capability of TiAl6V4 is to be observed (figure 3). Additionally, new results concerning the behavior of lubricants also can be obtained by such trials. The used material for de tool cavity was a molybdenum MHC based material, which can withstand forming temperatures higher than 1550\u00b0C. Due to the formation of a titan oxide layer on the material surface and the bad corrosion resistance of the MHC material at high temperatures of the tool (~650\u00b0C), the entire forming process was reduced with argon. In order to prevent adhesion of the work piece to the tool, it is necessary to use a special lubricant" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001897_bf01974774-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001897_bf01974774-Figure1-1.png", + "caption": "Fig. 1. Schematic representation of thrust mea-", + "texts": [ + " I t appears clear then, that the applicability of eq. [1.2] would depend on the level of concentration of the materials concerned and also on the range of shear rates. Experimental Apparatus The instrument used to measure axial thrust in the present work was essentially the same as that used by Shertzer and Metzner (9), with a slight modification. A Zenith gear pump (a constant-volume metering pump) was used to supply polymer solutions from a reservoir into a capillary, which was suspended horizontally between two flexible steel plates. (See fig. 1.) The axial thrust of liquid jets was measured by determining the deflection of the plates, as recorded on a Servo Recorder (Heath Kid, Model EUW-20A) by means of a displacement transducer (Hewlett Packard, Mode124DCDT). The transducer has a sensitivity of +0.050 inches in full scale and was calibrated against a known horizontal force. In order to increase the sensitivity of the recorder (250 millivolts full scale), a voltage divider (1/100) was used, and a Zener reference supply (Heath Kid, Model EUA-20-27) was used to stabilize the servo of the recorder", + " The argument was based on the theoretical considerations due to Middleman and Gavis (3) and some experimental observations made on the die swell behavior of Newtonian liquid jets (3, 8). Fig. 11 shows plots of (rll)R,5 vS. Reynolds number for ET 597 solutions of various concentrations with two capillary diameters. I t can be seen tha t (Til)R,L increases with Reynolds number, and that, at a given shear rate, Reynolds numbers for concentrated solutions are much lower than those for dilute solutions. This is attributable to an increase in viscosity as the concentration is increased. Note that Reynolds numbers in fig. 1 1 are calculated from the expression NRe ~ ~Dn v2-n using the flow curve data in fig. 3. I t may be surmised from fig. 11 that Reynolds numbers in polymer melts are exceedingly small, in general. Nevertheless, previous studies by Han et al. (13, 16) show that the exit pressures, -- (~2~) R, L, correlate with die swell ratio in a number of tests made over the range of Reynolds numbers investigated, say 10 -5 ~ 10 -a. Based on the facts presented above, the authors believe tha t the surface tension and the exit effects might be of secondary importance, if not negligible, in determining the elastic properties of viscoelastic solutions of moderate to high concentrations and of polymer melts" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000212_mfi.1994.398393-Figure7-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000212_mfi.1994.398393-Figure7-1.png", + "caption": "Figure 7: An example of the map combination", + "texts": [ + "2 COMBINATION OF TWO MAP USING DEMPSTER-SHAFER RULE Each map has been expressed as the occupied prohability of cells, that is, denoted as P(s(ci) = occ]. D-S rule (Dempster-Shafer rule) is applied for combining two maps and increasing the reliability as follows. P (8 (Ci) = occ] (13) - P, [s (Cj) = ace] P, [S(CI;) = occ] 1 - b Here, xs = Pu(l - Po) + P,(1 - P,) , P, and P, are probability in visual sensor map and ultra-sonic sensor map, respectively. As an example, Fig. 5 combines with Fig. 6 become fused information as shown in Fig. 7. With the above combined Map (Fig. 7), robot positions can be estimated using the least square method or probability method. 4 SIMULATION RESULTS The proposed methods are tested on an environment as shown in Fig.8. The mobile robot move from ongin point to final place through intermediate points, P1 and P2. Here, origin point is (O ,O) , final goal is (85.85), and intermediat,e points are P1( 15,15) and P2(85,15), where the units of numbers are 0.1 meter. At first, we tested dead-reckoning method on this environnient. The results are shown in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002482_s1068798x08070058-Figure3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002482_s1068798x08070058-Figure3-1.png", + "caption": "Fig. 3. Ball\u2013channel contact: (1) contact arc (ABC); (2) contact area; (3) frictional path Lx for a single ball\u2013channel contact; D2\u2013D2, instantaneous axis of ball rotation relative to the channel.", + "texts": [ + " 7 2008 WEAR CALCULATIONS FOR RADIAL BALL BEARINGS 647 area and takes the following form, according to the Boussinesq solution (14) where \u03b7 is the reduced elastic modulus of the contacting materials; pxy is the contact pressure at the point (x, y, 0), which is at a distance r from the point where the elastic strain is determined; ao and bo are the semiaxes of the contact-area ellipse. The gaps between the ball and various channel profiles were presented in [10]. For a channel of circular profile (R2 \u2248 1.03R3), the gap between bodies 1 and 2 at a distance x from the Oz axis in the xOz plane (Fig. 2) is (15) where R2ch and R3 are the radii of the channel and the wall (Fig. 1c). During a single contact with the ball at point E (Fig. 3), which is a distance x from the plane of rotation yOz, the channel wear is as follows, according to Eq. (11) (16) where Lx = is the frictional path at the selected point of the channel. The duration of a single ball\u2013channel contact at this point is t = 2bx/V, where bx is half the length of an element of contact area in cross section x; V is the velocity of the ball center O relative to the channel in the Oy direction. The contact surface of the ball moves relative to point E of the channel in the y direction, at constant slip velocity The frictional path at this point is where xc is the coordinate of the instantaneous axis of ball rotation relative to the channel; R3 is the ball radius. The instantaneous axis D2\u2013D2 of ball rotation (Fig. 3) occupies an intermediate position between the lower point of the channel and the ends of the contact arc. The total torque created by frictional forces over the whole contact area relative to the axis of rotation of the freerolling ball must be zero. According to Heathcote, xc is ux \u03b7 pxy r ------ dxdy, bo\u2013 +bo \u222b ao\u2013 +ao \u222b= zx R3 R2ch\u2013 R2ch 2 x2\u2013 R3 2 x2\u2013 ,\u2013+= \u2206h2x G pxy HB -------\u239d \u23a0 \u239b \u239e m dy, 0 Lx \u222b= V xdt 0 t\u222b V x V 2 --- xc R3 -----\u239d \u23a0 \u239b \u239e 2 x R3 -----\u239d \u23a0 \u239b \u239e 2 \u2013 .\u2248 Lx V xt bx xc R3 -----\u239d \u23a0 \u239b \u239e 2 x R3 -----\u239d \u23a0 \u239b \u239e 2 \u2013 ,\u2248= numerically equal to the coordinate (along the Ox axis) of a quarter of the volume of the pressure semiellipsoid with semiaxes ao, bo, po, positioned between the yOz plane and the parallel plane passing through the point (xc, 0, 0)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001730_095440605x32048-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001730_095440605x32048-Figure1-1.png", + "caption": "Fig. 1 Bearing geometry and schematic illustration of roughness structures", + "texts": [ + " The thrust of the present article is to extend the stochastic roughness theory [3] to investigate into the isotropic roughness\u2013slip interaction in case of the steadystate performance characteristics in terms of load capacity, end flowrate, and frictional parameters of hydrodynamic porous journal bearings of finite width at different parameters of practical importance, i.e. bearing feeding parameter, slip coefficient, clearance ratio, and eccentricity ratio. Moreover, the present analysis includes the solution of the governing equation for three-dimensional flow in the porous bush, i.e. Darcy\u2019s law along with the equation of continuity. Therefore, this analysis is based on a more generalized approach to obtain results valid for any value of porous bush thickness. The porous bearing configuration with roughness is shown schematically in Fig. 1. For an isoviscous incompressible lubricant, the steady-state lubrication model appropriate to the isotropic roughness pattern of a porous journal bearing with the effect of velocity slip is represented by the following two equations. In the porous bush Ku @2 p0 @u2 \u00fe R H0 2 @2 p0 @ y2 \u00fe D L 2 Kz @2 p0 @ z2 \u00bc 0 (1) In the clearance region @ @u E{ H 3 (1\u00fe ju)} @ p @u \u00fe D L 2 @ @ z E{ H 3 (1\u00fe jz)} @ p @ z \u00bc 6 @ @u \u00bdE{ H(1\u00fe j0u)} \u00fe b @ p0 @ y y\u00bc0 (2) where Ki \u00bc ki ky , i \u00bc u, z Non-dimensional film thickness with surface roughness H \u00bc h(u)\u00fe hs(u, j) Here, h(u) denotes the nominal smooth part of the film geometry and hs is the part due to the surface roughness, as measured from the nominal level" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003526_icicip.2010.5565222-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003526_icicip.2010.5565222-Figure1-1.png", + "caption": "Fig. 1. Leader-follower formation", + "texts": [ + " PROBLEM FORMULATION We consider a multi-vehicle system that consists of N USVs, each of which is modeled as velocity-controlled vehicle with kinematics [19] \u03b7\u0307i(t) = \u23a1 \u23a3 cos\u03c8i(t) \u2212 sin\u03c8i(t) 0 sin\u03c8i(t) cos\u03c8i(t) 0 0 0 1 \u23a4 \u23a6 \u03bdi(t) (1) where \u03b7i(t) = [xi(t), yi(t), \u03c8i(t)] T \u2208 SE(2) represents a position vector in the inertial reference frame; (xi(t), yi(t)) \u2208 2 represents the position in Cartesian coordinates; \u03c8t \u2208 (\u2212\u03c0, \u03c0] represents the heading angle; \u03bdi(t) = [ui(t), \u03c5i(t), ri(t)] T \u2208 3 represents the velocity vector in the body-fixed reference frame; (ui(t), \u03c5i(t)) and ri(t) represent the linear and angular velocities, respectively. Each of USVs is equipped with a single screw propeller and a rudder. Fig.1 presents the basic geometric structure about two USVs moving in a leader-follower formation. The line-ofsight (LOS) range \u03c1 and angle \u03bb are defined as: \u03c1ij(t) = \u221a (xi(t)\u2212 xj(t))2 + (yi(t)\u2212 yj(t))2 (2) \u03bbij = atan2(yi(t)\u2212 yj(t), xi(t)\u2212 xj(t)) (3) where \u03c1ij(t) \u2208 > 0 and \u03bbij \u2208 (\u2212\u03c0, \u03c0]. The formation control presented in this paper are based on the measurement of LOS range and angle between the leader and the follower. The formation specification is defined as sij(t) = [\u03c1ij(t), \u03bbij(t)] T (4) where sij(t) \u2208 2" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001210_tac.1972.1100049-Figure3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001210_tac.1972.1100049-Figure3-1.png", + "caption": "Fig. 3. The controlling moment.", + "texts": [ + " By choosing i o = iMo/(kl + iP&), the initial periodic modulation (nutation) of the angular velocity vector, generated by the complex pole (kl + iPIz) / I , will be canceled. 111. APPLICATIONS FOR SOME ROTATING BODY CONTROL PROBLEXS Consider t,he stabilization of free gyros with viscous gimbal damp- ing. The object is to keep the symmetry axis of t.he gyro in a fixed direction in inertial coordinates (e = 00). A disturbance A0 in the posit.ion of 8 mll be assumed. A controlling moment M, proportional t o A0 and inclined a t an angle E to A0 is shorn in Fig. 3. This relationship may be d-cribed in the following block diagram (Fig. 4): (k = I z / I ) , vhich in turn corresponds to t.he tmmfer function In analyzing the stability of this system, the charact.eristic equation wil be rewritten H(s) = S2 + - S + ikPS + !!! a + i - b kl JfO I I I (8) where a = cos E and b = sin E . Then, We may expand (9) to obtain the continued fraction (10) (see Appendix B): T ( S ) - = - + , L a - SI iI &fob\\ 1 g(s) 121 - + 7 A Then, for kl > 0, the condit.ion for stability is It is necessary to pursue further investigation in order to obtain the best combination of l l l o and E " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002428_01495730701738280-Figure9-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002428_01495730701738280-Figure9-1.png", + "caption": "Figure 9 Initial crack growth direction for the case of d = 0 3.", + "texts": [ + " Corresponding to Figures 6a and 6b, Figures 7a and 7b show the variations of energy release rate G as a function of 1 \u2212 0 for 4 different values of 0 150 \u2264 0 \u2264 180 and another 4 different values of 0 340 \u2264 0 \u2264 370 correspond to 2 peak values in Figure 2 for the case of Sr = 0 1, f = 0 1 and d = 0 5. From Figure 7a, maximum value of G(320.3Pa \u00b7m) occurs at 1 \u2212 0 = 0 for 0 = 170 and from Figrue 7b, maximum value of G (318.3Pa \u00b7m) occurs at 1 \u2212 0 = 0 for 0 = 360 . These results are shown schematically in Figure 8a. In much the same way as mentioned before, Figures 8b and 8c show the maximum values of G for a couple of crack direction ( 1 and 0 for Sr = 0 1, f = 0 7 and Sr = 0 5, f = 0 7, respectively for the case of d = 0 5. Similarly, for relatively shallow inclusion (d = 0 3), Figure 9 show schematically the maximum values of G for a couple of crack direction ( 1 and 0 for three cases of (a) Sr = 0 1, f = 0 1, (b) Sr = 0 1, f = 0 7 and (c) Sr = 0 5, f = 0 7, respectively. From Figures 8 and 9, it can be seen that the initial crack growth is caused almost at a radial direction emanating from the inclusion within 1 \u2212 0 \u2264 10 , it shows 2 possible initial crack growth points, which is chosen the first largest G max point and the second largest G max point are located at the angle 0 (130 \u2264 0 \u2264 170 and 330 \u2264 0 \u2264 360 " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001607_his-2004-13-407-Figure7-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001607_his-2004-13-407-Figure7-1.png", + "caption": "Figure 7. Piano movers\u2019 problem", + "texts": [ + " 1983b), in which the robots must determine which path to take in order to move an object around an obstacle of this shape and layout. More precisely, the problem is simplified to enable the object to be transported by actual humanoid robots, and the aim is to transport the piano through an L-shaped passageway of a fixed width. The difficulties of this task obviously lie in finding the feasible path through a narrow passage while avoiding collision with the walls. Suppose that we use the path (of width ) and the piano (i.e., the target object of size ) shown in Fig.7. Then, the following relationship should be satisfied for all deg deg : (1) Note that the left side of the equation (1) is the or component of the point p3. It becomes maximized when deg . The maximum value is given as follows: ! \"$#%\" & (2) Hence, the Equation (2) gives the minimum width of the passageway that allows a rectangular object of width and length to pass through it. In our experiment, we used the HOAP-1 operating PC with two servers: Motion Server which executes operation commands, and Vision Server which processes visual data from the CCD camera" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000478_s0967-0661(00)00096-4-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000478_s0967-0661(00)00096-4-Figure2-1.png", + "caption": "Fig. 2. A free-#oating robot. I inertial frame, 0 manipulator's base frame.", + "texts": [ + " One of the controls increases its amplitude while the other decreases. The approach is applicable when controls have no restriction on their amplitudes being equal to 1. (2) only one control is scaled, either w i (Jt/m, m, 1) or w i (Jt/m, 1, m) and the time is prolonged. The scaling coe$cient should be calculated by optimizing a performance criterion which takes into account the cost of generation and distance reduction from the goal. To illustrate ideas put forward in this paper a free#oating robot presented in Fig. 2 was chosen as an object. Kinematic equations of the robot derived from Dule7 ba (1996), are as follows (l\"1 [m], a mass (m) takes any value): [x5 y5 HQ q5 1 q5 2 ]T\"Xu 1 #>u 2 , X\" 1 8C 2s 1H#s 12H# m 1 a 33 (sH#2s 1H#s 12H ) !2c 1H!c 12H! m 2 a 33 (cH#2c 1H#c 12H ) 8m 1 /a 33 1 0 D , >\" 1 8C s 12H# m 1 a 33 (sH#2s 1H#s 12H ) !Mc 12H# m 2 a 33 (cH#2c 1H#c 12H )N 8m 2 /a 33 0 1 D with m 1 \"!(19#6c 1 #12c 2 #3c 12 ), m 2 \"!(7# 6c 2 #3c 12 ), a 33 \"34#12c 1 #12c 2 #6c 12 . A standard robotic convention is used s 12H\"sin(q 1 # q 2 #H), etc" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001493_iros.2006.281826-Figure5-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001493_iros.2006.281826-Figure5-1.png", + "caption": "Fig. 5 - 7, show that hair rotation, coupled with the lever effect makes it easier to complete the hair detachment. However, a too-long hair might have other drawbacks, such as increased matting [10].", + "texts": [ + " The lever effect After observing various attachment-detachment sequences of Magnetic Hair we found out that there always seems to be a preferred path for detachment: a path of least effort. Note that when detaching the hairs, we are not concerned here on the exact motion equations, but rather with the fact that the minimum energy detachment path seems to start always with a rotation (Fig. 4 - 7). Fig. 4 shows an attachment-detachment sequence of a hand-sized model of a gecko's foot (a \"magnetic\" version). Each hair ends in a cylindrical magnet that simulates the Van der Waals interaction, but at a bigger scale. Fig. 5 Shows the detachment condition when the hair, originally adhered to a substrate, rotates over axis P. A hair subject to a external force (for example an actuator) rotates if MRELEASE > MAdh; where MAdh = moment over axis P due to Adhesion forces, and MRELEASE moment due to a external force applied on the hair. Fig. 6 is a force-curve obtained experimentally. It shows the pull force a magnet experiments when placed on an iron plate for different rotations. The dashed line is corresponds to a 2 mm diameter 0", + " The bending of the hair that occurs in the area right of \"Release mode\" is unrealistic in a close packed (carpet-like) arrangement of hairs. (Fig. 3) From Fig. 12 we conclude that hairs in Tack mode are softer and therefore can \"stick\" better. Hairs in release mode are stiffer and thus are easier to release. E. The Moment Distribution Another effect of the characteristic curvature of the hairs is the ability to distribute economically [9] a big load (as see in Fig. 2) into smaller loads to each hair (MAdh of Fig. 5). The importance of this function, (the complementary of the peeling effect of section C), comes to the fore in systems that don't ensure uniform load distribution. In Fig. 13b a load of 63Kg on a magnetic pad causes a momentum that is not distributed uniformly into the magnets. This causes an unwanted detachment (peeling) of the inferior part that in turn causes a spiraling (and dangerous) loss of adhesion of the system as a whole. The characteristic curved shape gecko hair seems naturallyfitted [9] to avoid this issue by conveying loads and tensions acting in one end of the hairs to the substrate efficiently" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002246_13506501jet421-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002246_13506501jet421-Figure1-1.png", + "caption": "Fig. 1 Squeeze film geometry between two different spheres", + "texts": [ + " To account for the couple stress effects arising from the lubricant blended with various additives, the non-Newtonian squeezefilm Reynolds-type equation between two different spheres is derived using the Stokes motion equations. A closed-form solution for the film pressure, the loadcarrying capacity, and the response time is obtained for engineering application. The influence of nonNewtonian couple stresses on the squeeze film characteristics under different sphere-to-sphere geometries is analysed by comparing with the corresponding Newtonian-lubricant case. Figure 1 shows the squeeze film configuration of two different spheres. The origin of the fixed cylindrical coordinates locates at half distance of the minimum film thickness. The upper surface S2 is approaching the lower fixed surface S1 with a squeezing velocity, Vsq = \u2212\u2202h/\u2202t . The two surfaces S1 and S2 forming the gap, provided R1, R2 x, can be approximated by parabolic relationships by Hamrock [11] h1 = hm 2 + r2 2R1 (1) h2 = hm 2 + r2 2R2 (2) where hm denotes the minimum film thickness along the vertical line of centres, h1 is the upper film thickness, and h2 is the lower film thickness measured from the horizontal r\u03b8-plane" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000702_t-pas.1975.31999-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000702_t-pas.1975.31999-Figure1-1.png", + "caption": "Fig. 1. The experimental machine.", + "texts": [], + "surrounding_texts": [ + "Introduct ion\nThe a v a i l a b i l i t y of pmier4ul connu t a t i o n a l f a c i l i t i e s 11% l e d t o a widespread i n t e r e s t i n the c a l c u l a t i o n o f the maqnetic f i e l d s w i t h i n electric machines. TO be of pract ical value such deterninat ions must t a k e in to accoun t nagne t i c s a tu ra t ion ani? n u s t he economical. There is now an extensive l i t e r a t u r e d e s c r i b i n g a v a r i e t y of i t e r a t i v e techniques for accoropl ishinq this which a r e based on e i ther the f in i te t l iFFerence [ 1 , 2 , 3 ] or f i n i t e l emen t t echn iques [ 4 ,SI , t hese re ferences b ing typ ica l . Iiowever , t h e experimental data by which the accuracy and economy of the var ious methods could he compared i s fa r nore l imi ted . Xi th the a in of reducing th i s def ic iency , a svnchronous machine has been equipped w i t h comprehensive instrumentation for the measurement O F n a g n e t i c f i e l d s and torcrue under a wide va r i e ty o f s t eady- s t a t e and t r a n s i e n t e x c i t a t i o n s . Here the apparatus i s b r i e f l y descr ibed and t h e results of nen-circui t f ie ld computat ions are qiven and compared with p e a s u e d d a t a . Complete d e t a i l s of the machine dimensions , i t s m a t e r i a l p r o p e r t i e s and t h e r e s u l t s of a verv extensive series of experiments are avai lable to thers involved i n program developrzent by a p p l i c a t i o n t o Professor T.H. Barton, Faculty of Enuineerinq, : ; k G i l l Univers i ty , I lontreal , Canada.\n'The program\nThis paper i s t h e f i r s t of a series i n khich it i s Intended t o r e p o r t the r e s u l t s 04 measurements' and computations of var ious aspec ts o f the nachine f ie ld . nor the p resent purpose , a t ten t ion i s concentrated anthe radial conponent of the a i r 7 a n f i e l C i n r e g i o n s s u f f i c i e n t l y f a r renovcii from the co re\nrbinery Committee of the LEEE Powcr Euginceriq sockty for presentation at the h p r T 75 1532, recommended md approved by the IEEE Rotating MaIEEE PES Wlntcr Mating, New York, N.Y., January 2631.1975. Manusaipt submitted August 30,1973; made available for printing Novrmber 18,1974.\nends for it t o be ax ia l ly un i fom. T h e Cinitm element method based on t r i a n q l a r e l e m e n t s has been mployed, the proqrarq heinn n modification of that developed by S i l v e s t e r , Cabayan and 3rovne [GI. T h i s prooran is based on the nagnet ic model l ing and cornnutationzl .methods descrihetl by S i l v e s t e r and Chari [4] , bu t i s more economical and i s e a s i e r t o use. I t e r a t ion o f t he so lu t ion towards i t s u l t ima te value i s contj.pued u n t i l t h e r e s i d u a l s a r e redaced t o 10 w h i l e t h e r a t i o of t h e maTnitude of t h e p o t e n t i a l t o t h e magnitude of ti15 r e s i d u a l i .e. n o m ($1 /nom (3) , i s 2.5 x 10 .\nfhe tes t machine\nThe test machine is a three-phase, four-pole, synchronous machine rated a t 133q rev/F.in, 254 v o l t s , 14.5 anps per phase, 10 hv, 0.8 pf. leading. The s a l i en t po le f i e l c? i s on t h e s t a to r ; each po le a r r i e s a 1068 t u r n exci ta t ion windinq and t h e r e i s a s r ru i r r e l cage darnper windinn i n t h e pole faces . The three-phase windinq i s on t h e 48 s l a t r o t o r which is skerlcd I1:7 one s l o t p i t c h . The core lenqth is 17.3 a, t h e r o t o r c l i m e t a r i s 20.85 m, t h e s t a t o r b o r e i s 21.41 cm. The airnan under the mid-section of the po les i s t h e r e f o r e 3.20 x.n. The p o l e t i p s a r e chanfered fron a point about 39 e l e c t r i c a l from the po le ax i s , t he a i rgap i nc reas ina l i n e a r l y t o 7.94 n.n. a t the t o p , which is rounded. Pigure 1 denicts the nachine mounted so a s t o f a c i l i t a t e f l u x n e a s u r e m e n t s .\nAs may be seen the machine is mouterl with i t s axis v e r t i c a l . The r o t o r is supported on a torque t ransducer and can be ro ta ted s lowly by a motorized worn dr ive. A l l experiments areth refore performed under mechanical ly s teady s ta te condi t ions. T h i s i s not regarded as a s i n n i f i c a n t l i m i t a t i o n s ince . the inf luence of the mechanical parameters on e l e c t r i c a l t r a n s i e n t s i s small .\nThe r a d i a l a i r q a p f i e l d i s measured by a Hal l probe which can be positioned anvwhere i n t h e a i r q a p by r o t a t i o n a l and v e r t i c a l movements. Rotat ional novement i s notorizeA, y e x t i c a l Jnovenent, i s nanual.\nThe angular pos i t ion of t h e r o t o r and the angular and v e r t i c a l p o s i t i o n s o f t h e Hal l -probe are inc?icatcd both direct ly on s c a l e s and as ana lope vo l t a rps de r ived from potentiometers.\nThe probe output is fed t o a qaussneter d i r e c t l y i n d i c a t i n g the f lux dens i ty and a l s o providing an analogue voltaqe proportional t o f lux dens i ty .", + "The torque t ransducer s imi la r ly g ives direct indica t ion of to rque toge ther w i t h an analogue voltage output.\nThe vo l t age ou tpu t s p ropor t iona l t o pos i t i on , f l ux dens i ty and torque when appl ied t o an XY p l o t t e r o r o t h e r e c o r d i n g d e v i c e permit the rap id acquis i t ion of a l a rue amount of data .\nResul ts\nA l a r g e number o\u20ac radial f lux dens i typosit ion curves have already been obtained. Circular t r ave r ses a t va r ious ax ia l pos i t i ons and a x i a l t r a v e r s e s a t key radial loca t ions have be n made under magnetic conditons ranqing from unsa tu ra t ed to h igh ly sa tu ra t ed .\nThe magnetic conditions are ind ica t ed by Figure 2 which s ows the mean f lux dens i ty over the pole arc a t a po in t midway down the core s tack . Exper imenta l resu l t s were obtained a t va r ious cons t an t l eve l s o f e x c i t a t i o n c u r r e n t by i n t e q r a t i n a t h e f luxmeter out put as the Uall probe was traversed round the airgap a t cons tan t speed. These r e s u l t s are shown by t h e f u l l l i n e . Computations were made f o r e x c i t a t i o n s of 2 amperes, the rated value, 4, 6 and 8 amperes, these abnormally higher values being chosen s i n c e w e wished t o emphasize the f fects of magnet ic sa tura t ion . Agreement between computed and measured r e s u l t s is good up t o about 150% of noma1 exc i t a t ion , i.e. 3A. Beyond this there is a proqress ive ly\nincreas ing d ivergence , the ca lcu la ted va lue being 8.3% h i g h a t 400% of noma1 exc i ta t ion . The la t ter yields very high values of flux d e n s i t y i n t h e t e e t h , we es t imate va lues as high as 2.16 webers/m i n c o n t r a s t t o t h e maximum v a l u e a tr a t e d e x c i t a t i o n of 1.6 webers/m. There is considerable , doubt as to the accuracy of the BH c h a r a c t e r i s t i c o?? t he i f o n a t these very h igh sa tura t ion leve ls so t h a t c o m p u t a t i o n a l e r r o r s i n t h i s r e g i o n a r e no t pa r t i cu la r ly d i s tu rb ing .\n\u2018The a i rgap f l u% dens i ty d i s t r ibu t ion w i t h normal e x c i t a t i o n , i.e. I = 2.OA, and the ro tor pos i t ioned so t h f t a too th ax is coincided w i t h each pole axis and each neutral zone, is shown i n F i g u r e 3. The measurements were made with the probe half way down the core s tack and t h u s uninfluenced by end e f f e c t s .\nComputed results a re g iven fo r po in t s c l o s e t o t h e r o t o r s u r f a c e , midway between r o t o r and s t a t o r and c l o s e t o t h e s t a t o r surface. To obta in these , the a i rqap w a s t r i angu la t ed i n h ree concen t r i c circular bands. The computed d i s t r i b u t i o n s a r e t y p i c a l o f the f in i te t r iangular e lement t echnique i n t h a t h e flux density remains constant over the width of an element and changes uddenly a t an element boundary. Techniques ar avai lable for smoothing these var ia t ions but it w a s f e l t t o be u s e f u l t o i n d i c a t e t h e i r f u l l e f f e c t . The inf luence of t h e r o t o r s l o t openings is most apparent as i s the diminution of this e f f e c t w i t h i n c r e a s i n g d i s t a n c e from the rotor. There is no means of con t ro l in s the r a d i a l p o s i t i o n of the Ilall probe and it may, i n fact , vary between the xtremes represented by the three computed d i s t r i b u t i o n s . The f i n i t e . H a l l c r y s t a l w i d t h of 1.2 mm i s s i g n i f i c a n t compared t o t h e s l o t opening of 2.5 mm and w i l l g ive subs t an t i a l smoothing of the rapid variations i n f l ux which occur a t s l o t edges. Bearing i n mind these two e f f ec t s , p robe r ad ia l pos i t i on and c rys t a l w id th , the genera l f i t between the", + "measured and computed d i s t r i b u t i o n s is good and, as noted in the d i scuss ion of F igure 2 , the average value over a po le p i t ch o f computed and measured d i s t r i b u t i o n s i s t h e same.\nFigure 4., which cgnpares calculated and measured f l u x d i s t r i b u t i o n s a t f i e l d c u r r e n t s of 2 , 4 , 6 and 8 amperes , is i n t e r e s t i n g i n t h a t it i l l u s t r a t e s t h e i n f l u e n c e s o f t h e damper winding s l o t s . There are f ive of these p e r p o l e , p o s i t i o n e d r e l a t i v e t o t h e s l o t s a s shown i n F i g u r e 5. They are b u r i e d i n t h e pole face being closed by s l o t b r i d g e s 1.5 mm deep. A t normal e x c i t a t i o n l e v e l s t h e s e b r idges a r e no t s a u ra t ed and have o inf luence on the a i rgap f lux d i s t r ibu t ion . AS t h e x c i t a t i o n l e v e l is increased the s l o t b r i d g e s s a t u r a t e and t h e i r e f f e c t becomes progressively more apparent. Considering\nFigure 4d, the pole axis damper s l o t , coinciding with a ro tor too th , causes a d i p i n f lux dens i ty . The n e x t s l o t , c o i n c i d i n g w i t h a r o t o r s l o t , augments the f lux dip caused by t h e l a t t e r . The t h i r d s l o t , l i k e the f irst , coincident with a ro tor too th , causes a d ip i n f lux dens i ty .\n3W\nL\nFig. 5. Quarter cross-sect ion of the machine taken between two adjacent pole axes.\nComents and conclusions\nThe computational programs andthe experimental equipment which avebeen developed for the work descr ibed here provide an extremely powerful and convenient too l for the inves t iga t ion of electric machine f i e l d s which it i s i n t e n d e d t o e x p l o i t i n a va r i e ty of ways. Most aspects of the computational process haveb enmechanized and the pre l iminary resu l t s g iven here i n d i c a t e t h a t the process is accurate up t o abnormally hiqh leve ls of magnet ic sa tura t ion .\nI f a detailed p i c t u r e of the f lux d i s t r i b u t i o n is requ i r ed , t he nrrmber o f f i n i t e elements must be increased , e .g . , the th ree layers ofairgap elements u ed here. The elements m u s t a l s o be carefully chosen with an eye t o the expec ted d is t r ibu t ions to main ta in efficiency of omputation. However, macros c o p i c e f f e c t s , e.g. , mean f l u x d e n s i t i e s , pole f luxes and winding inductances, can be cor rec t ly p red ic ted w i t h su rp r i s ing ly few elements." + ] + }, + { + "image_filename": "designv11_69_0003276_ihmsc.2010.76-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003276_ihmsc.2010.76-Figure2-1.png", + "caption": "Figure 2. The geometry of the morphing wings.", + "texts": [ + " While the projectile with the morphing wings can alter the aerodynamic characteristic, as a rather unique control mode, the dynamic model of missile with morphing wings has to be analyzed. 978-0-7695-4151-8/10 $26.00 \u00a9 2010 IEEE DOI 10.1109/IHMSC.2010.76 2890 In fact, the projectile with morphing wings is different from the ordinary projectile with fixed wings as the dynamic characteristic of the whole projectile with morphing wings during the wings altering. The morphing wings of the projectile are assumed as Fig. 2 shows a 2-D view of the model of morphing wings. In order to analyze the dynamics of the missile with the morphing wings, based on the muti-rigid body dynamics model method, the vehicle is regarded as one composed of the morphing wings and the body section. The moment of the projectile is stated as 1 2= +H H H , (1) where H is the gross momentum moment of the missile; 1H is the momentum moment of the morphing wings; 2H is the momentum moment of the body section. It is assumed that 1m is the mass of the morphing wings and 2m is the mass of the body section, the relationship formula 1 2m m m+ = is satisfied" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003271_02644400910959160-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003271_02644400910959160-Figure2-1.png", + "caption": "Figure 2. Critical geometries description; (a) DIN and JGMA standards, (b) AGMA standard", + "texts": [ + " There are many standard approaches attempted to describe the dimensions of the critical geometric parameters (h and t) in spur gears. As reported in Sraml and Flasker (2007) and Ciavarella and Demelio (1999), in both Deutsches Institut fu\u0308r Normung (DIN) and Japanese Gear Manufactures Association (JGMA) standards, the lines drawn at 30 from the center line are constructed such that they run tangentially to the fillets of the gear tooth, from which this construction and further calculations of t can be determined, see Figure 2(a). While AGMA standard has an intersecting Lewis parabola which is inscribed in a gear tooth profile. The vertex of the parabola lies at the intersection of the tooth profile centre line and the projection of the applied force, as shown in Figure 2(b). In addition to the critical geometries, the stresses on the fillet radius (rf) should be taken into consideration because it is a region that gives the maximum stress (Mehmet, Stress concentration factor 363 1999). In general, designers and manufactures recognized a fillet region in gears as the highest stress point. In this region, Kt is a main consideration of stress calculations in gears. Therefore, in addition to many theoretical equations and diagrams that are used for calculating the stress around that region, there are several practical methods, such as strain gauges, and photoelasticity used for the same purpose" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002827_6.2008-7126-Figure3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002827_6.2008-7126-Figure3-1.png", + "caption": "Figure 3. ELV Free-body Diagram of Pitch-Plane", + "texts": [ + " Important assumptions for all those examples are that the reference models must be stable and are developed by a Jacobian linearization of the system American Institute of Aeronautics and Astronautics 3 dynamics about trajectory equilibrium (trim) points. They also assume a special linear form of system dynamics, where the adaptive controller is assumed to augment a baseline linear controller. None of these assumptions are required for the method described in this paper. II. Equations of Motion The derived equations of motion are presented with relation to a defined coordinate system. For simplicity, only the pitch-plane of the ELV motion is studied (see Figure 3). Several assumptions are made that allow for simplification of the equations. The equations of motion of an ELV are complicated by the fact that the vehicle has time-varying mass and inertia. There can also be relative motion between various masses within the vehicle and the origin of the body axes, such as fuel sloshing, engine gimbal rotation, and vehicle flexibility. The derivations and modeling assumptions in this research follow those stated by Greensite1. The launch azimuth is assumed to be directly east" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003976_icma.2009.5244953-Figure4-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003976_icma.2009.5244953-Figure4-1.png", + "caption": "Fig. 4. Hook Toe on Grid Line", + "texts": [], + "surrounding_texts": [ + "Feasible walking trajectories have to satisfy many condi tions such as workspace, collisions, limitations of motors and hanging condit ion. These independent conditions are hard to analyze, which means that ordinal optimization methods are hard to implement. Then we propose to implement a genetic algorithm [6] to optimize the walking trajectories [5]. A GA represents the solution as a chromosome which is a set of genes. The serching methods are crossover and mutation. Crossover is exchanging some genes among two or more chromosomes. Mutation is replacing some genes of a chromosome. These operations prevent converging one local solution and need no evaluation analysis such as derivation. So it is easier to evaluate by factors in parallel. A GA is one of the best methods for our study because of its easy implementation. A. Implementation of GA We implement a GA to optimize gait parameters. Height and Offset are the integral multiples of the grid interval and Width is the odd multiple of the grid interval. So, it is reasenable to represent them by integrals or odds. And it is accurate enough to represent Position and Distance by milimiters. Finally, we define the binary as the gene and the 27-bit string as the chromosome as shown in Fig.7. We describe restraint conditions and evaluation factors for the GA. We define the restraint conditions as below; 1) Trajectories are within the workspace, 2) No collisions among legs and grid lines, 3) Joint torques are not over the motor limitation, 4) Stability margin angles are within \u00b1 90 degrees. We define the evaluation factors as below; 1) Achievement, 2) Energy consumption, 3) Maximum joint torque , 4) Average stability margin angle, 5) Maximum stability margin angle. The first evaluation factor indicates achievement ratio of one period walking satisfying the restraint conditions. A feasible solution has lOO%-value achievement. It is the only factor to be maximized. The others are to be minimized. Each evaluations are calculated independently in case of walking upward or downward and in case of walking rightward or leftward . Assume that the walking is static, calculations of the factors are based on the equiliblium. The equiliblium is a statically indeterminate problem. Hence we implement to ASTERISK a simple load distribution control based on position errors of legs, we assume that reaction forces acting on toes are norm-minimized. Next, we describe the evaluation of the solutions . We take the dominace ranking method [7] to evaluate the factors in parallel. Now we explain the dominance and Pareto optimality in case of minimization. We give a set of solutions D and evaluation functions Em(x)(m = 1, ... , M) , xE D . Then x is dominant to y when And x* is the Pareto optimum if none is dominant to x*. Pareto optima have rank 1. The set of Pareto optima is called Pareto frontier, and Pareto optima of the D excluding the original Pareto frontier have rank 2. The greater ranks are defined by the same way. The dominance ranking method proposes that operations of a GA are operated based on the ranks. But, our implementation is modified to give a priority to the achievements. We consider a solution dominant when it has a dominance with the achievements , even if it has no dominance with the other evaluations . We determine the dominance with the other evaluations only when we cannot determine the dominance with achievements. Also we give priority orders among solutions with the same rank based on their linear-conbined evaluations other than their achievements . At last, we describe the update of generations. We fill 1/4 of the next generation with the elites in three orders; one order is based on both sets of evaluations of walking upward or downward and walking rightward or leftward , another is based only on the set of evaluations of walking upward or downward, and the other is based on the set of evaluations of walking rightward or leftward. We take as the selection method the roulette wheel selection whose roulette wheels are based on the rank values. We make two roulette wheels; one is based only on the set of evaluations of walking upward or downward and named UD roulette , and the other is based only on the set of evaluations of walking rightward or leftward and named LR roulette. We take as the crossover method the uniform crossover between two solutions . We fill another 1/4 of the next generation with the children which have both parents selected by UD roulette and selected by LR roulette . We fill other 1/4 of the next generation with the children which have parents selected only by UD roulette . We fill the other 1/4 of the next generation with the children which have parents selected only by LR roulette. We give the high mutation probability. A mutation happens 5% of the time for each gene of the children. Fig.8 shows the next generation." + ] + }, + { + "image_filename": "designv11_69_0000282_tmag.2003.810347-Figure11-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000282_tmag.2003.810347-Figure11-1.png", + "caption": "Fig. 11. Contours of magnetic flux density (mechanical angle 10 ).", + "texts": [ + " We discuss the influence which the deterioration of magnetic characteristic of the claw pole has on the holding torque characteristic. The direct current of 235 AT is supplied to each coil in the same direction. In addition, this analysis is carried out using the analyzed model shown in Fig. 4, meshes shown in Fig. 5, the B-H curve shown in Fig. 6, the magnetization of permanent magnet shown in Fig. 3, and the analysis condition shown in Fig. 7. The contours of the magnetic flux density in the root of the claw pole on mechanical angle 10 are shown in Fig. 11. From Fig. 11, the same tendency in as the case of cogging torque is shown. The holding torque characteristic in Type 0 is shown in Fig. 12. This figure shows that one cycle is 30 . Although not illustrated, the holding torque characteristics of Types 1\u20134 are also the same waveform. A comparison of the amplitude of holding torque is shown in Fig. 13. The difference of the maximum and the minimum of holding torque is defined as the amplitude. The amplitude of holding torque in Type 0 is the larger than other models, and it becomes small in order of Type 1, Type 2, Type 3, and Type 4" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000512_810105-Figure11-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000512_810105-Figure11-1.png", + "caption": "Fig. 11 - Effect of original point setting error", + "texts": [ + " made by cutting. The transverse profiles of both gears take similar form by lapping as seen in Fig. 10 (b). From Fig. 10 (c) where the tooth bearing on the gear tooth is shown by the black area we can see that the bearing is in agreement with that estimated from the measured results of the profiles. The effects of the accuracy of the original point setting on the measured results are discussed in the following. 810105 7 When the original point setting has an error in x-direction by AX as shown in Fig. 11, the angular error A0ox of the test gear at the reference position is given by A0\u201e no* AX (31) tloy Xo which causes the tooth error A0ox>Xi\u00abniy at the measuring point. Since the deviation of the measuring position by AX causes another error -nixAX, then the total tooth error (e)Ax caused by deviation of the original point comes to Similarly the profile error caused by deviation of the original point in y and z directions, A Y and Az, can be discribed as follows. (e)AY = Cf^niy - niy)AY (e)Az= < ^ ^ n i y ~ n i z ) A Z (33) (34) The errors of the lengthwise profile and transverse profile, when the original point deviates by 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001784_11539902_140-Figure3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001784_11539902_140-Figure3-1.png", + "caption": "Fig. 3. Desired and control locus and obstacle avoidance", + "texts": [ + " Since V\u0307S \u2264 0, VS \u2208 \u221e, which implies that \u0303kji, \u03c3\u0303kji, w\u0303kj \u2208 \u221e, if the Jacobian is full rank, hNFS \u2208 \u221e and kji, \u03c3kji, wkj \u2208 \u221e, so, \u0302kji, \u03c3\u0302kji, w\u0302kj \u2208 \u221e and h\u0302NFS \u2208 \u221e. Considering that hres \u2208 \u221e, so K\u03b5 \u2208 \u221e. Then, from Eq. 21, s\u0307 (t) \u2208 \u221e. Since s (t) \u2208 2 and s\u0307 (t) \u2208 \u221e, s (t) \u2192 0 as t \u2192 +\u221e, which is followed by e\u0307 (t) \u2192 0. End of the proof. The simulation is performed on a real robot composed of a 3-wheeled mobile platform and a 4-DOF modular manipulator, as shown in Fig. 1(a). In order to verify the algorithm, the robot is required to follow a spacial trajectory in Fig. 3(a), which has been planned to ensure the robot far away form singularities or joint limits. Two ball-like task-consistent obstacles with radius of 0.2m are considered, one is on the motion plane of the mobile platform and the other is on the way of the modular manipulator as shown in Fig. 3(b). The simulation time is selected as 20 seconds. Each element of h is approximated by a NFS. The gain matrices and constants are selected as follows: KP = diag {100} , KI = diag {10} , K\u03b5 = diag {50} , \u0393 kji = 0.1, \u0393\u03c3kji = 0.1, \u0393wkj = 0.1, \u039b = diag {2.0} and Nr = 200. The cut-off distance is selected as dc = 0.5m, and the coefficient is determined by k\u03c6 = 1.0. The desired and the controlled locus are shown in Fig. 3(a). Two obstacles are avoided by controlling self-motions of the mobile modular manipulator in Figure 3(b). The tracking position and velocity errors are given by Fig. 4. It can be observed that the proposed algorithm is effective in both avoiding obstacles and controling the end-effector to follow a desired spacial trajectory simultaneously. A mobile modular manipulator composed by a 3-wheeled nonholonomic mobile platform and a n-DOF onboard modular manipulator is investigated in this paper. Firstly, an integrated dynamic modeling method is presented. Secondly, a new obstacle avoidance algorithm using self-motions is proposed, which can avoid obstacles without affecting the end-effector planning task" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003592_1.4001258-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003592_1.4001258-Figure2-1.png", + "caption": "Fig. 2 CFD mesh", + "texts": [ + "org/about-asme/terms-of-use o s F t 7 i t fl s s c E s f c t c T W D M 0 Downloaded Fr \u2022 In the case of unsteady-state, exhaust mass flow rate can be approximated by a step function, where each inlet is either opened or closed. In order to carry out the CFD-calculation, a multipurpose gemetry containing the fluid parts exhaust gas, coolant and the olid part Steel-AISI 1010 has been created in Pro-Engineer\u00ae. or the FE analysis, only the solid part of the model is used. The geometry is then imported to Gambit\u00ae for meshing. A 3D etrahedral mesh, shown in Fig. 2, consists of approximately 1.1 106 elements. These are divided in 350,000 solid elements and 35,000 fluid elements. The boundary layer mesh was attached to mplement prism layers with a fine resolution at the interface beween the fluid and solid part. This allows a proper simulation of ow in the wall boundary layer and of the heat transfer to the olid wall 10 . After the mesh is generated and imported into FLUENT \u00ae, solver etting, material property, boundary conditions, and operation onditions must be specified properly" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000329_robot.1994.351067-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000329_robot.1994.351067-Figure1-1.png", + "caption": "Fig. 1. 4 dof redundant robot", + "texts": [ + " (3) Becauseself-motion ismntinuousinaneighborhd of m=O, joint accelerations do not exceed their bounds when self-motion is stopped. (4) Becausesystemoscillationisavoided,the weighting factors can be easily selected for multiple criteria. We describe an example to demonstrate the efficiency of this method for solving inverse kinematics of a redundant robot. q(0) = [O,O, 0, OJT rad/sec. The initial states for this example are asumed to be q(0) = [1.45,1.41,0.00, - 0.5OlT rad, We consider a four degreeof-freedom robot described by Fig. 1. The workspace is described by the DenavitHartenberg parameters listed in Table 1. and we desire the End-effector trajectory: straight line, motion: Desired time: td = 4.00 sec Sampling period: End-effector velocity: At = 0.02 sec X = [ ~ . ~ O , - O . ~ O , O . ~ O ] ~ m/sec Joint constraints (i=l, 2,3,4) kill 1.57 rad k i l l 1.57 rad/= (2) Singuiarity avoidance: 2 2 c2 =cos 92 +cos 94, c20 =0.01 Computer simulation provides the following results: (1) The joint velocity only consists of its minimumnorm solution, i" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003558_ipec.2010.5544492-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003558_ipec.2010.5544492-Figure1-1.png", + "caption": "Fig. 1. Definition of estimated frame.", + "texts": [ + " Sanada* * Osaka Prefecture University, 1-1 Gakuen-cho, Naka-ku, Osaka, Japan 978-1-4244-5393-1/10/$26.00 \u00a92010 IEEE \u0394++ + \u2212+ = q d aq d qd qd refq refd D D V i i pLRL LpLR v v \u03c9\u03c9 \u03c9 0 _ _ (3) +\u2212+ \u2212\u2212\u2212 \u2212 = )( )( )( )sin()cos( )sin()cos( sincos 3 2 3 2 3 2 3 2 3 2 w v u T q d if if if D D \u03c0\u03c0 \u03c0\u03c0 \u03b8\u03b8 \u03b8\u03b8 \u03b8\u03b8 (4) where vd_ref and vq_ref are the reference voltages in the d-q frame, R is resistance including armature resistance and the on-resistance of the switching device in the inverter, Dd and Dq are state variables dependant upon the rotor position \u03b8 and the phase currents iu, iv, iw. Fig. 1 shows the relation of the reference frames. The d-q reference frame, which is synchronously rotating at electrical angular speed \u03c9, is usually used for an IPMSM control system. However due to lack of a position sensor in a sensorless control system, the d-q reference frame is not be available. Hence, the estimated rotating frame, the \u03b3\u2212\u03b4 frame, which lags the d-q reference frame by position error \u03b8e is used. The IPMSM model in the \u03b3\u2212\u03b4 frame derived from (1) is described as follows. + + \u2212+ = \u03b4 \u03b3 \u03b4 \u03b3 \u03b4 \u03b3 \u03c9 \u03c9 e e i i pLRL LpLR v v dq qd (5) \u2212 \u2212+ \u2212 = \u03b3 \u03b4 \u03b4 \u03b3 \u03c9\u03c9 \u03b8 \u03b8 i i LE e e d e e ex )\u02c6( cos sin (6) )])((}){([ qqdadqdex piLLiLLE \u2212\u2212+\u2212= \u03c9 (7) where v\u03b3 and v\u03b4 are the \u03b3- and \u03b4-axis terminal voltages, i\u03b3 and i\u03b4 are the \u03b3- and \u03b4-axis armature currents, e\u03b3 and e\u03b4 are the \u03b3- and \u03b4-axis extended EMF and \u03c9\u0302 is the estimated electrical angular speed" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003686_msec2009-84049-Figure6-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003686_msec2009-84049-Figure6-1.png", + "caption": "Figure 6. Sequential 3D Off-axis HPDL deposition profile and temperature distribution with cross sections. Laser Power: 2100 W, beam spot: 12\u00d70.5 mm2 with 6.35 mm defocus down. Stellite 6 powder flow rate: 35 g/min, scanning speed 3.33 mm/s", + "texts": [ + " The powder flow with the temperature profile has been predicted as in Figure 4. As seen in Figure 4, the powder stream, which is injected from the flat nozzle, flows towards the laser focus area. Due to gravity, the powder stream spreads when x distance increases. In the laser irradiation zone, particles are quickly heated up from ambient temperature to around maximum 2000 K, indicating some particles are melted. After passing through the laser beam, the powders cool down due to the convection to the surrounding gases. As seen from Figure 6, the powder temperatures tend to be uniform in the middle of the powder stream along z direction, which is reasonable because of the uniform laser beam intensity there. It is worth to note that the particle trajectory is critical to determining the particle temperature history because it 4 Copyright \u00a9 2009 by ASME Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 09/21/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use determines the laser particle interaction time and the laser intensity that the particle experiences, thus influencing the powder temperature distribution [18]", + " Main material properties for powder and substrate [8-9, 29] Property, Symbol (Unit) Stellite 6 Mild steel(1018) Density, \u03c1s or \u03c1l (kg/ m3) 8380 7800 Specific heat, cps or cpl (J/kg K) 421 610 Solid conductivity, ks (W/m K) 14.8 50.0 Liquid conductivity, kl (W/m K) 48.8 35.0 Latent heat, Lm (J/kg) 2.92e+5 2.46e+5 Liquidus temperature, Tl (K) 1630 1803 Solidus temperature, Ts (K) 1533 1766 Dynamic liquid viscosity, \u03bcl (Pa s) 5.96e-3 6.10e-3 Surface tension coefficient (N/m K) -1.12e-4 -4.90e-4 A sequential three dimensional track evolution in a 1.8 s period during the process is illustrated in Figure 6. As seen in Figure 6, a distinct, wide, rectangular molten pool is generated at the front of the track with the moving beam. The melt solidifies very quickly due to the high attendant cooling rate and forms a track after the laser moves away. A wide track with a height about 0.4 mm and little dilution is being formed. Corresponding cross sections along the scanning direction are also displayed. It can be seen that the molten pool peak temperature can be as high as around 1795 K at t=1.8 s and the fluid motion velocity is about 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001132_0470870508.ch20-Figure20.3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001132_0470870508.ch20-Figure20.3-1.png", + "caption": "Figure 20.3 Dipole and monopole charges are associated with the Gyricon bead. On application of a field, the bead traverses the cavity and rotates. On reaching the wall it adheres", + "texts": [ + " The beads are called bichromal because they consist of two colors. Associated with each color is an electric charge; one hemisphere has a different magnitude of charge than the other, or a different polarity. Each bead has a dipole moment that is proportional to the charge difference between the two hemispheres. It also has a monopole moment that is proportional to the net charge on the bead. Both are important to the electro-optical behavior of the bead. A single bead in its cavity is illustrated in Figure 20.3. Initially, the bead is adhered to the bottom of the cavity, with its white side facing up. As the electric field is applied, both the monopole and the dipole moments of the bead are acted upon. Rotation of the bead is initially prevented by the adhesion of the bead to the cavity wall. However, the field acts on the monopole moment of the bead, pulling it away from its adhesion to the cavity wall. Once this has happened, the dipole moment of the bead causes the bead to rotate, as the bead crosses the cavity" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001218_detc2006-99153-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001218_detc2006-99153-Figure1-1.png", + "caption": "Figure 1. Reuleaux\u2019s method.", + "texts": [ + " An over determined version of Reuleaux method is formulated and a solution of the problem is found using the least square method. At the end, some examples are provided to verify the solutions methods. Reuleaux\u2019s method [11] use geometric construction to find the instantaneous pole of a planar displacement using two homologous (separate position of the same) points of a rigid body. The instantaneous pole is found to be the intersection point of the bisecting lines of the lines joining each of the homologous points P1,P\u20321 and P2,P\u20322 as shown in fig. 1. The total angle of rotation of the rigid body around the pole can be also determined from the construction and thus the circular motion of the body is reconstructed. The angle of rotation will be the angle between the line connecting the pole to any point of the body at the first position and the line connecting the pole to the same point af- 2 nloaded From: http://proceedings.asmedigitalcollection.asme.org/pdfaccess.ashx?url= ter rotation. Reuleaux\u2019s method can be used for the kinematic registration problem [10]", + " Midlines can be written as m\u0302a = (ma,m0 a) = 1 2 (l\u0302a + l\u0302\u2032b) m\u0302b = (mb,m0 b) = 1 2 (l\u0302b + l\u0302\u2032b) Next, we find the lines perpendicular to both the midlines m\u0302a and m\u0302b and the common perpendiculars s\u0302a and s\u0302b (see fig. Copyright c\u00a9 2006 by ASME l=/data/conferences/idetc/cie2006/71356/ on 02/27/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use Downlo 5). Again, equation (14) is used to find the Plu\u0308cker coordinates of these lines. c\u0302a = (ca,c0 a) = m\u0302a\u00d7 s\u0302a c\u0302b = (cb,c0 b) = m\u0302b\u00d7 s\u0302b The lines c\u0302a and c\u0302b correspond to the bisecting lines in the two dimensional version of Reuleaux\u2019s method (see fig 1). The line coordinates of the screw axis that represent the motion from the first to the second position of the given lines will be the normalized common perpendicular between c\u0302a and c\u0302b namely: s\u0302c = |c\u0302a\u00d7 c\u0302b| = (s,s0) (2) The screw parameters of a helical motion are a line and a pitch. we have found the Plu\u0308cker line coordinate of the screw from equation (2). We will use the same construction method to find the pitch of the screw. In the two dimensional of Reuleaux\u2019s method. The angle of rotation of the rigid body around the pole is the angle between the line connecting any point of the body at the first position to the pole and the line that connects the corresponding point after rotation to the pole (see fig 1). Following the same construction for the three dimensional generalization of the Reuleaux\u2019s method, we can easily see that the angle of rotation of the rigid body around the screw axis will be the angle between the common perpendicular of any line of the rigid body before displacement and the screw and the common perpendicular of the corresponding line after displacement and the screw axis (see fig. 6). 4 aded From: http://proceedings.asmedigitalcollection.asme.org/pdfaccess.ashx?url Let g\u0302 be the common perpendicular between any line of the body at the first position and the screw axis and g\u0302\u2032 be the common perpendicular between the corresponding line after the helical motion and the screw axis" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000355_robot.2001.932765-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000355_robot.2001.932765-Figure2-1.png", + "caption": "Figure 2: Relittion b/w I z l p . ~ and the upper bound of v", + "texts": [ + " Thercfore, the following inequality is always satisfied by Weyl's Theorein in [2] Since A,,,i,(kZZ\") is zero, the minimum eigenvalue of QI( is not sinaller than the minimum value among diagonal ciitries of Q. 0 Above analysis can be explained easily as follows: In the case of trajectory tracking control for robot manipulators, we start, the siniulation/experiInerit with zero error z = 0 after adjusting initial conditions. The value of Lyapiinov fiiiiction is zero at the start time, however, the error increases to some extent because V(0 ,O) may have any positive constant smaller than c3y2 as shown in Figure 2. This Figure depicts the upper bound of V vs. 121 of the eqiiation (27). Since the suggested perforniance limit 1zlp.L is tlie convergent point as we can see in Figiirc 2, the Euclidian norm of error tends to stay at this poilit,. This analysis can naturally illustrate the gain t uniiig. The PID gain tuning has been an important subject, however, it has not been much investigated till now. Recently, tlie noticeable tuning method was suggested as the name of \"square law\" by Park et a1[9]. They showed that the square law is a good tuning method through their experiments for the industrial robot manipulator" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002899_1.5061071-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002899_1.5061071-Figure1-1.png", + "caption": "Figure 1 : Pure Zirconia 3D Objects manufactured by SLS/SLM Technology (a-CAD design, b-real manufactured part)", + "texts": [ + " The dark colour of the parts results from a lack of oxygen, above 200\u00b0C Zirconia turned back into white. Finally the developed technology was applied to manufacture parts to demonstrate the geometry reproduction capability and to measure the density. According to the process parameters discussed above, Design of Experiments was applied to indicate the major parameters. Those parameters are directly linked with the laser power density distribution. Tests have been performed according to Table II. Accurate geometry and 3D pure zirconia objects can be obtained using this SLS/SLM Process (Fig 1, Fig 2). The density remains low, only 56% (Fig 3). One may note that further sintering in conventional furnace cannot increase the object density. This states the fact that laser melts part of the ceramic and freezes the structure (Fig. 4). To improve ZrO2 component density, one may improve the powder bed density. Further studies need to be done regarding the way to process the powder bed, before laser action (powder layering sequence). Indeed the theoretical density of a powder bed composed of single size particles before laser action is 74%" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002302_13506501jet152-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002302_13506501jet152-Figure1-1.png", + "caption": "Fig. 1 (a) Diagram of the drive and (b) solid model of the drive", + "texts": [ + "eywords: toroidal drive, friction, friction coefficient, sliding ratio The toroidal drive was proposed by Kuehnle [1] in 1966. The drive consists of four basic elements, Fig 1: (a) the central input worm; (b) radically positioned planets; (c) a stator of toroidal shape; and (d) a rotor, which forms the central output shaft upon which planets are mounted. The planets have balls or rollers instead of teeth. Each planet meshes with the toroidal grooves in the stator. The rotor is the output. The drive can transmit large torque in a small size and is suitable for top end technical fields such as aviation and space flight, etc. Kuehnle et al. [2] invented a special machine tool that can produce finished stator", + " According to energy principle, following relationship should be given p0m \u00bc pama \u00fe ml(p0 pa) From this equation, the friction coefficient between mesh teeth can be obtained as m \u00bc ma pa p0 \u00fe ml 1 pa p0 Based on the analysis, the friction coefficient m between mesh teeth for the drive is m \u00bc ma (l 4 0:4) m \u00bc ma pa p0 \u00fe ml 1 pa p0 (0:4 4 l 4 3) m \u00bc ml (3 4 l) (2) where ma is the asperity friction coefficient and ml the EHL friction coefficient. From equation (2), it is known that the friction coefficient m between mesh teeth for the drive is decided by the asperity friction coefficient ma, EHL friction coefficient ml, and the ratio l. Friction between the planet and the stator or the worm is comprised of rolling friction and sliding one. Sliding ratio between both surfaces is defined as ratio of the relative sliding distance between them to the total relative moving distance between them. As shown in Fig. 1(a), a planet tooth meshes with stator at angle f1. Let df1 denote an angular increment of the planet at angle f1. If symbol i01 denotes speed ratio between planet and stator in the coordinate system attached to the rotor, df denotes angular increment of the stator about axis o2 o, dlt denote the length of the moving arc of the planet tooth in the circumferential direction of the stator, the arc length dlt can be calculated as dlt \u00bc (a\u00fe R cosf1)df \u00bc i01R a R \u00fe cosf1 df1 (3) where a is centre distance and R the radius of the planet rolling circle" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001577_piee.1971.0022-Figure3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001577_piee.1971.0022-Figure3-1.png", + "caption": "Fig. 3 Variation of angle a with Pf, Qb = 0, V = 1 0 a Machine directly connected to busbar b With transmission line of impedance = 0-05 +y'0-15p.u. c With transmission line of impedance = 0 1 0 +./0-30p.u. d With transmission line of impedance = 0-20 +./0-60p.il.", + "texts": [], + "surrounding_texts": [ + "If constant flux linkages are assumed during the first swing, eqns. 2 and 4 give the transient torque/angle characteristics of the unregulated d.q. and conventional machines. Typical curves are shown in Fig. 2. It has already been pointed Equal-area criterion a Typical transient-torque/angle characteristic of conventional machine b As a, but for a d.q. machine A2 = area above xy line for conventional machine /I3 = area above xy line for d.q. machine A, = A2 = A-i S\u00abi = maximum swing of load angle of conventional machine for given clearing angle Sec 8\u00ab2 = as \u00a7\u00ab!, but for a d.q. machine Se3 = maximum possible swing of Se in conventional machine without loss of stability S,!4 = as S83, but for a d.q. machine out in Reference 6 that the d.q. machine can swing to greater angles than the conventional machine and still remain stable. This is because the transient torque/angle characteristic has been shifted from the origin of the steady-state characteristic by an angle a. The conventional machine also has a slight built-in advantage because of its transient saliency (represented by the sin 28 term in eqn. 2), but this is small compared with the advantage gained by the backswing a. The value of a depends on the initial conditions. Figs. 3 and 4 show the 144 PROC. IEE, Vol. 118, No. 1, JANUARY 1971 variations of a for different Pb and Qb, respectively, and also for different transmission-line impedances. The angle a increases at high values of Pb and at leading values of Qb, but is reduced as the transmission-line reactance increases. Hence it can be concluded that the advantage derived in the increase of transient stability of a d.q. machine is greater when the machine is operating near full load, and also when it is absorbing reactive power. Fig. 5 shows the variations of E], $\\ and 8e for unregulated d.q. and conventional machines during a transient swing due to a 3-phase fault, and includes the results both with and without the assumption of constant flux linkages. As would be Fig. 4 Variation of angle oc with Qb pb = 1 0 , v = 1 0 a Machine directly connected to busbar b Transmission-line impedance = 0 0 5 -1-yO- I5p.u. c Transmission-line impedance = 0 1 0 + y 0 - 3 0 p . u . (I Transmission-line impedance = 0-20 +y '0 -60p .u . Transients in unregulated machine Pb = I 0, Qb = 0, clearing time = 0- 1 s Transmission-line impedance before fault = 0 05 +y'0-15 Transmission-line impedance after fault = 0-10 +y0-30 a El in d.q. machine, assuming constant flux linkages c E] in conventional machine, assuming constant flux linkages b and d. As (a) and (c), respectively, but without assuming constant flux linkages e \\ in d.q. machine with constant flux linkages / As (e), but without assuming constant flux linkages g So in d.q. machine, wiihout assuming constant flux linkages h As (g), but for a conventional, machine k As (&'), but assuming constant flux linkages m As (k), but for conventional machine PROC. 1EE, Vol. 118, No. 1, JANUARY 1971 expected, the rotor-angle excursions are less when constant flux linkages are assumed. The d.q. machine has a greater rotor-angle excursion than the conventional machine, and the oscillation is less damped.9 This is because, for the conditions considered, a positive area above the line xy (in Fig. 2), equal to the negative area below the line, can only be found with higher swings for a d.q. machine. However, as the clearing time is increased, this pattern is reversed, and the conventional machine pulls out of step earlier than the d.q. machine. This can also be observed from a study of if and ,- during the transient period. Figs. 6 and 7 illustrate the variations of if, he, S and ,- in unregulated conventional and d.q. machines, for clearing times of 0-20 and 0-21 s, respectively. For the latter time, the conventional machine becomes unstable. It can be seen that, in a d.q. machine, for clearing times near the critical value, <\u00a3,- swings below cf>e, thus bringing (S + <\u00a3,) nearer to 90\u00b0. This results in higher torques near the peak of the rotor-angle swing (eqn. 3), which improves the transient stability. Alternatively, for given values of sin (5 + (/\u00bb,) or Te, the Transients in unregulated machine Pb = 1 0, Qb = 0, clearing time = 0-20s Transmission-line impedance before and after fault = 0 0 5 + . / 0 - 15p.u. d.q. machine conventional machine Transients in unregulated machines Pb = 1 0, Qb = 0, clearing time = 0-21 s Transmission-line impedance before and after fault = 0-05 +j0- 15p.u. d.q. machine conventional machine 145 160- O-2 O-4 O-6 time, s Fig. 8" + ] + }, + { + "image_filename": "designv11_69_0002698_iros.2007.4399277-Figure5-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002698_iros.2007.4399277-Figure5-1.png", + "caption": "Fig. 5. Inverse kinematics of avoidance of a gap on a thin wall.", + "texts": [ + " Avoidance of a straight line (cylinder) When the link LM is avoiding a straight line of which position is sr , the link LM must be located apart from the line by the distance of more than c (in case of cylindrical obstacle, it is c + cB) by rotating joint Jk (see Fig. 4). Assuming the joint angles besides k\u03c6 are known, k\u03c6 realizing contact limit is solved by coordinate transformation matrices. Similarly to the case of avoiding a point (a ball), there are four patterns of solutions. B.3. Avoidance of a gap on a thin wall When the link LM is passing through a gap on a thin wall, one lateral DOF is restricted by the gap. Other DOF of the link are free (see Fig. 5). The link LM is located between the gap by rotating the joint Jk, which is located before the wall. There are two patterns of solutions of k\u03c6 , which are right and left rotations. B.4. Avoidance of a gap on a thick wall When the link LM is passing through a gap on a thick wall, two DOF, which are lateral movement and lateral rotation, are restricted by the gap (see Fig. 6). Therefore, two additional DOF are necessary. The joint JM is located in front of the gap by rotating the joint Jk, which is located before the wall" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001455_s00542-006-0293-x-Figure9-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001455_s00542-006-0293-x-Figure9-1.png", + "caption": "Fig. 9 Foil bearing model for web-roller interface", + "texts": [ + "0 (m/s), and the web stops completely when the roller velocity reaches to Ur = 5.0 (m/s). Fig. 3 Surface roughness distribution Fig. 4 Probability density of test web Two theoretical models for predicting the slip onset velocity, Model 1 and Model 2 are formulated as follows. At the first step of modeling, it is necessary to estimate the entrained air film thickness and traction coefficient between the web and roller. The entrained air film thickness will be obtained based on the foil bearing model as shown in Fig. 9, in which the entrained air film thickness and air film pressure are determined by solving simultaneously the following equations (Hashimoto and Okajima 2006). @ @x h3p @p @x \u00fe @ @z h3p @p @z \u00fe 12k tw p\u00f0p ps\u00de \u00bc 6gU @\u00f0ph\u00de @x \u00f02\u00de T R T @2w @x2 \u00bc 1 L ZL=2 L=2 p pa\u00f0 \u00dedz \u00f03\u00de h \u00bc w\u00fe d \u00f04\u00de d \u00bc 1 2R x\u00fe RB 2 2 x\\ RB 2 0 RB 2 x RB 2 1 2R x RB 2 2 x[ RB 2 8>>< >>: \u00f05\u00de The boundary conditions for the modified Reynolds equation (2) are given as follows: p\u00f0xs; z\u00de \u00bc pa; p\u00f0xe; z\u00de \u00bc pa \u00f06a\u00de p x; L 2 \u00bc pa; p x; L 2 \u00bc pa \u00f06b\u00de On the other hand, the boundary conditions for the web equilibrium equation (3) are given by: w\u00f0xs\u00de \u00bc 0; w\u00f0xe\u00de \u00bc 0 \u00f07a\u00de @2w @x2 x\u00bcxs \u00bc 0; @2w @x2 x\u00bcxe \u00bc 0 \u00f07b\u00de where xs and xe indicate the coordinates of inlet and outlet boundaries of web wrap region, and these are the unknown variables" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002819_6.2007-3813-Figure30-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002819_6.2007-3813-Figure30-1.png", + "caption": "Figure 30. The effect of pin location on forces and moments.", + "texts": [ + " While the forces generated on the body are an order of magnitude greater than when the pins are located inside the cavity, the moments on the body are of the same order of magnitude. High pressures are generated at the nose in front of the pin, but there are large areas of low pressure due to the separation behind the pin which result in offsetting the higher pressures at the nose. When the pins are located near the tail of the projectile, the net force and the magnitude of the moment increase relative to pin locations at the nose and the midbody as shown in Figure 30. For the tail location, there is little of the projectile body remaining that experiences the separated flow behind the pins and thus the net force is higher. One can also see that for both the nose and the tail locations that the sparseness of the pressure taps in these regions will affect the overall integrated force. While it was not the original intent to capture high detail data in these areas, it is felt that the data still provides a good representation of the changes in the surface pressures even if the forces are off to some degree. Figure 30 also illustrates another important point. Comparing the peak pressures produced upstream of the pin for each pin location, it is found that the increase in surface pressures are on the order of 3 psi for pin locations at the Page 15 nose and the tail while the surface pressure increase is closer to 0.3 psi for the pin located inside the cavity at Tap 14. This tenfold increase in pressure is due to the fact that the pin inside the cavity is shielded from the freestream. Based on these measurements and those in Ref" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003735_s0025654410020020-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003735_s0025654410020020-Figure1-1.png", + "caption": "Fig. 1.", + "texts": [ + " In [2], to make the use of these functions in problems of dynamics more convenient, he constructed Pade\u0301 approximations corresponding to them. This approach allowed him to develop an essentially new two-dimensional model of sliding and spinning friction and further construct several models of combined dry friction based on the use of Pade\u0301 approximations [3]. A distinguishing feature of these models is that all of them are constructed under the assumption that the Coulomb law in differential form holds for a small surface area element in the interior of the contact spot (Fig. 1). But experiments demonstrate that the graph of the actual characteristic of the friction law has the form shown in Fig. 2 [4] and can be described by the function F = f sign v \u2212 av + bv3, (1.1) where v is the relative sliding velocity of the bodies, f is the friction coefficient, and a and b are parameters determined by the coordinates of the minimum of the characteristic in Fig. 2. *e-mail: kireenk@ipmnet.ru 166 The use of the Coulomb law in generalized form (1.1) in the case of combined kinematics, where the bodies in friction simultaneously participate in sliding and spinning motions, results in the following: the differentials of the friction force dF and the friction torque dMC with respect to the center of the contact disk (Fig. 1) are determined by the formulas dF = \u2212f\u03c3 V |V| (1 + \u03bc1|V|3 \u2212 \u03bc2|V|) dS, dMC = \u2212f\u03c3 r\u00d7 V |V| (1 + \u03bc1|V|3 \u2212 \u03bc2|V|) dS, V = (v \u2212 \u03c9y, \u03c9x), r = (x, y), (1.2) where \u03c9 is the angular spinning velocity, \u03c3 is the normal contact stress distribution, and the coefficients \u03bc1 and \u03bc2 can be determined experimentally. MECHANICS OF SOLIDS Vol. 45 No. 2 2010 2. COUPLED SLIDING AND SPINNING FRICTION MODELS 2.1. Integral Model By integrating the expressions (1.2) over the contact spot, we see that the force F and the torque MC have the form F = \u2212f \u222b\u222b G \u03c3(x, y) V |V| dx dy \u2212 f \u222b\u222b G \u03c3(x, y)V(\u03bc1V2 \u2212 \u03bc2) dx dy, MC = \u2212f \u222b\u222b G \u03c3(x, y) r\u00d7 V |V| dx dy \u2212 f \u222b\u222b G \u03c3(x, y)r \u00d7 V(\u03bc1V2 \u2212 \u03bc2) dx dy, G = {(x, y) : x2 + y2 R2}. (2.1) Under the assumption that the normal contact stress distribution for circular contact sites has the property of central symmetry, it is convenient to write the integrals (2.1) in polar coordinates {r, \u03d5} with pole at the center of the contact spot (Fig. 1). As a result of the change of variables x = r cos \u03d5, y = r sin \u03d5, r \u2208 [0, R], \u03d5 \u2208 [0, 2\u03c0], in formulas (2.1), the magnitudes of the friction force and torque have the form F = f 2\u03c0\u222b 0 R\u222b 0 r\u03c3(r)(v \u2212 \u03c9r sin \u03d5) dr d\u03d5\u221a v2 \u2212 2v\u03c9r sin \u03d5 + \u03c92r2 + 2\u03c0f [ (\u03bc1v 3 \u2212 \u03bc2v) R\u222b 0 r\u03c3(r) dr + 2\u03bc1v\u03c92 R\u222b 0 r3\u03c3(r) dr ] , MC = f 2\u03c0\u222b 0 R\u222b 0 r2\u03c3(r)(r\u03c9 \u2212 v sin \u03d5) dr d\u03d5\u221a v2 \u2212 2v\u03c9r sin \u03d5 + \u03c92r2 + 2\u03c0f [ (2\u03bc1v 2 \u2212 \u03bc2)\u03c9 R\u222b 0 r3\u03c3(r) dr + \u03bc1\u03c9 3 R\u222b 0 r5\u03c3(r) dr ] . (2.2) Thus, just as in the case of the classical Coulomb law in differential form, the normal component of the friction force is zero" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000291_ias.1995.530343-Figure6-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000291_ias.1995.530343-Figure6-1.png", + "caption": "Fig. 6: Outline and winding layout of a cross-section of the motor under consideration", + "texts": [ + " It is obvious that all flux through the Game links both end ring and end windings. Therefore, it is part of the mutual inductance and not part of the leakage components. To conclude, separating the mutual and leakage components can only be done using a 3D approach. IV. THREE DIMENSIONAL Em RING INDUCTANCE CALCULATION A. Three dimensional model The motor modelled consists of 48 stator slots and 40 rotor slots. The stator has a two layer winding short-pitched by two slots. The motor has open rotor slots. Figure 6 shows the cross-section of one pole together with the winding layout. The 3D model is built using an extrusion based mesh generator [9]. Due to symmetry, only one fourth of one end region has to be modelled. The 3D model consists of a material mesh and a set of coil meshes required for the currents. Both meshes are generated separately allowing a different extrusion direction for material mesh and coil meshes. The material mesh can be built by rotating a base plan similar to figure 1 around the center line of the shaft or by shifting a base plan similar to figure 6 in the axial direction. All outlines required in the material mesh have to be present in the base plane. Figure 7 shows the material mesh when the extrusion is performed in axial direction. From this figure it can be seen that a part of the iron core 11s modelled as well. 51 7 The stator end winding is not incorporated in the material mesh. The end winding is modelled as a set of current driven coils in air. This is feasible since current redistribution due to skin effect is negligible in the stranded stator end winding", + " Therefore, 11 rotor coils are required for the current excitation. The stator winding is represented by 22 current driven coils. Only 2 end winding coils are completely inside the model, the other 20 coils are cut off at the boundaries of the model (figure 8). Figure 9 shows some of the end winding coils used for the end winding excitation Figure 9 shows the two end winding coils that are completely inside the model (coil 1 and coil 2) together with three other coils that are cut off at the boundary of the model. When referring to the cross-section of figure 6, coil 1 occupies the upper half of the first stator slot (the slot in the upper left comer) and the lower half of slot eleven, coil 2 occupies the upper half of the second slot and the lower half of slot twelf. In the real motor, each of the stator coils contains four turns. The currents for both rotor and stator coils are obtained from a two dimensional finite element analysis. This analysis is performed at various slip values. Figures 10 to 12 show the currents in the three stator phases, the currents in the rotor bars and ring segments as phasors for the different slip values. The numbering of the bars 51 8 is given in figure 6 . Ring segment 2 is the ring segment between bars 1 and 2. For clarity, the stator currents are enlarged 10 times. The currents in figure 12 are calculated with reduced supply voltage in order to obtain the rated stator current at standstill. From the figures 10 to 12, an increasing phase shift between the stator mmf and the rotor mmf is noticed. At standstill (figure 12) the stator mmf is apprciximately 180\" shitted with respect to the rotor mmf. The three dimensional problem is defined as time-harmonic, neglecting saturation in the iron core", + " Furthermore, the end ring leakage referred to the stator, Lor is found to be of the same magnitude as the end winding leakage inductance Lps. This already is an indication that the influence of the end ring leakage is not negligible in all load situations. The influence of the end ring leakage inductance, calculated by the 3D approach, is examined using a two dimensional finite element analysis by including or neglecting it in the analysis. A . Combinedfinite element - circuit model The outline of the finite element model is shown in figure 6. As the number of slots per pole is an integer, it is sufficient to model only one pole, and apply the appropriate boundary conditions. The analysis includes saturation using an iterative proces of both static (non-linear) and time-harmonic (linear) solutions [4]. Instead of using an effective magnetisation characteristic to obtain the reluctivities, the reluctivities are calculated based on two static solutions. From a time-harmonic solution, real and imaginary part of both stator- and rotor currents are extracted" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002200_09544054jem699-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002200_09544054jem699-Figure2-1.png", + "caption": "Fig. 2 Different designs for a hydraulic lever", + "texts": [ + "comDownloaded from working environment. Therefore, only the following two categories of DPs will be considered in this paper: (a) structure parameters (SPs), whose values can be specified in the design process; (b) time-domain parameters (TPs), whose values can only be specified within a range. Mathematically, DPs are defined as DPs \u00bc fSPs,TPsg \u00bc SPs; @SP @t [ TPs; @TP @t 6\u00bc 0 \u00f02\u00de SPs are constant with respect to time, and TPs are the functions of time. For example, in the hydraulic lever system shown in Fig. 2(a), the FRs for its cylinder component are as follows: FR1, at a certain position; FR2, supply a force. There are DPs associated with these two FRs: DP1, the support point position; DP2, the pressure in the cylinder. Of these, DP1 is an SP and its value can be specified in the design process, while DP2 is a TP and its value can only be specified within a range during the design process. In fact, the classification of DPs into SP or TP categories will also depend on the system\u2019s FRs and the design creativity. In the above example, DP1, the cylinder\u2019s support point, can be considered as an SP in a regular design. However, if the cylinder\u2019s support point needs to vary, then FRl is required of a variable position within a range. Accordingly, DPl can also be classified as a TP if driven by another cylinder as shown in Fig. 2(b). Here DPl2 [a, b], where a and b represent the lower limit and upper limit respectively of the cylinder\u2019s support point. The factual values of DPl need to be specified during system operation and must be steady. In this way a new design is invented. The benefit of transforming support points from SP to TP is that the optimal position of the lever\u2019s arm of force can be maintained in operation and thus the system efficiency is increased. The classification of DPs is a creative action in system design, in addition to analysing the system\u2019s FRs and scheme DPs [8]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002812_978-1-4020-8600-7_27-Figure3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002812_978-1-4020-8600-7_27-Figure3-1.png", + "caption": "Fig. 3 Experimental parallel manipulator.", + "texts": [ + " (b) Return to step (a) until all the possible orientations of the platform at the first point P0 have been explored. Then compare all the previously saved functions fcost and choose the trajectory having the greatest fcost. For validation purposes, our approach was applied on an experimental manipulator 3-RRR. This manipulator has been developed by the research groups of ITLag and IRCCyN. The values of its geometric parameters are: l1 = l2 = 26 cm, r = 29 cm and R = 36 cm. The prototype is shown in Figure 3. It was required to guide along the path shown in Figure 4 an axially symmetric tool fixed at the centroid of the mobile platform. We applied the proposed algorithm to find the orientations to be used by the mobile platform during the task. The prototype was positioned in the first working mode (WM1). Figure 5a displays the optimal path found in the feasibility map. Figure 5b displays the optimal values of \u03ba\u22121(A\u0304) along the prescribed path. Finally, Figure 6 shows a sequence of 4 configurations of the manipulator during the achievement of the task" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001047_ias.2003.1257638-Figure8-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001047_ias.2003.1257638-Figure8-1.png", + "caption": "Figure 8. Finite element analysis: flux distribution for rated s\u03c9e. (broken bars number 15 and 16).", + "texts": [ + " This is accomplished by setting the harmonic frequency in the Finite Element solver to the rated slip frequency. Figure 7 shows a comparison of the rotor bars current for the machine (referred to the stator) with symmetric rotor and asymmetric rotors. As it was presented in the simulation results, the currents on the bars adjacent to the faulty ones are increased. The absence of current in the faulty bars breaks the shielding effect of the rotor cage and the flux penetrates deeply in the rotor at this region as it is shown in Fig. 8. Since the flux pattern changes only at the fault region, the induced currents change only on the bars adjacent to the faulty ones, when compared to the symmetric rotor case. (a) broken bars number 15 and 16; (b) bar number 15 at fault. V. CONCLUSIONS In this paper, it has been presented a method suitable for computer simulation of induction machines with rotor asymmetries. The method is based on the classical fourth order transient model for symmetrical induction machines, with additional computation limited to the transformation of the rotor current vector to a rotor fixed reference frame and the application of linear transformations the extended rotor current vector" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000529_s0043-1648(02)00034-0-Figure13-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000529_s0043-1648(02)00034-0-Figure13-1.png", + "caption": "Fig. 13. Schematic illustration of the small roller from test no. 15 [13], on which 2 large and 11 small flaking failures were observed.", + "texts": [ + " 11, the values of |\u03c4 zx |max are also larger than it is possible to achieve due to the assumption of only elastic contact, and such high values can only occur theoretically. Flaking failures caused an excess vibration level in our double-roller testing, Fig. 1. Usually, one or more large flaking failures were observed to have occurred on the defect-free small roller surface after such testing. These large flaking failures, accompanied by both LM and OLM cracks, caused an increase in vibration. Also of note, several small flaking failures developed concurrently in most cases. Fig. 13 shows the schematic illustration of the small roller after test no. 15 [13], in which the symbols \u201c \u201d and \u201c \u201d indicate large and small flaking failures, respectively. Fig. 14 shows the top and cross-sectional views of one of the small flaking failures observed on the test no. 15 small roller, demonstrating that only OLM cracking may occur initially. The two flaking failures seen in Fig. 15 were also detected on the test no. 15 small roller. Fig. 15(a) has a small flaking failure with a short LM crack developing, that is a fraction of the length of the OLM cracks shown in this figure and in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003940_978-1-84996-432-6_19-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003940_978-1-84996-432-6_19-Figure1-1.png", + "caption": "Fig. 1. Roller hemming process", + "texts": [ + " The numerical simulation for this scenario is very complex, given the need to ensure the continuity of the forces of a nodal element to the next one throughout the whole process, for which a high number of iterations, and therefore a large time is required. In the automotive industry, hemming is used to join two sheet metal panels by bending the flange of the outer panel over the inner one. Currently in the industry there are basically two methods of mechanical hemming: the conventional mechanical hemming and the roller hemming (Figure. 1). The roller hemming process is generally carried out in three steps. The orientation (angle) of the roller changes in between the hemming steps. The tensile state created during roller hemming process (non bending plane strain) with a component in the axis of folding, favors a smaller elongation of the grain in the deformed area and thereby delay the onset of shear bands and fractures compared with plain strain bending state achieved using the conventional hemming. To perform and develop the FEM methodology three different geometries were selected, straight, concave and convex" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001222_isie.2006.296103-Figure7-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001222_isie.2006.296103-Figure7-1.png", + "caption": "Fig. 7. Mechanical setup", + "texts": [ + " In this case the utilization of special filter techniques according to [2] is not necessary, because the integral related to the uncertain detection of t1 is negligible. This can be seen by the measurement results in paragraph IV even for small backlash values like 2g = 0,5\u00b0 . Thus, it is sufficient to find the maximum value QM by making use of simple search algorithms. As the measurement results will show, the measurement of the position has to be carried out with a high resolution. For that reason the sinusoidal analogue signals of a 2048 pulse incremental encoder are used. The resolution reached is approximately 0,5 Mio. PPR. Fig. 7 depicts the mechanical setup. A permanent magnet synchronous machine with a rated torque of 15 Nm drives the mechanics. The machine is fed by an inverter with field oriented control. In order to enable thorough investigations in a wide range of backlash values a special mechanical backlash element has been designed and realized (see Fig. 8). The realized module is a claw-type construction [6]. It provides the possibility to change the backlash angle manually to certain values within a wide range" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002334_j.ics.2006.12.012-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002334_j.ics.2006.12.012-Figure2-1.png", + "caption": "Fig. 2. Module system: (a) power supply module, (b) assemble power supply module, (c) motor drivers module system, (d) assemble motor drivers system.", + "texts": [ + " The design concepts are as follows: (i) simple modular architecture with a laptop PC for easy assembly and maintenance (ii) omni-directional motion (iii) omni-vision system (iv) strong kicking mechanism (v) cableless system as much as possible. The specifications of the robot are shown in Table 1. The robot has been designed using 3D-CAD software (Fig. 1b). The driving mechanism consists of 3 sets of a motor, a motor driver, a planetary gear, and an omni-directional wheel. The various electric devices and circuits are designed as functional modules. Each functional module, such as the motor driver or USB serial controller, is developed as a \u201cpackage\u201d (Fig. 2). Therefore, we can remove a broken part and replace the part easily, and carry and assemble the robot without a large amount of labor. The behavior of the robot is programmed on a laptop PC, and the behavior commands are received by a referee box PC via onboard wireless LAN. The laptop PC, installed on the robot, sends motor control commands (target velocities) to the motor drivers through USB/ RS232C serial converters. The system architecture and the power system are illustrated in Fig. 3a, b" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003293_s0263574710000421-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003293_s0263574710000421-Figure1-1.png", + "caption": "Fig. 1. The proposed hyper-redundant manipulator.", + "texts": [ + " This subject has been approached by means of a continuous backbone curve which contains essential macroscopic geometric features of the desired motions.30\u201332 Finally, a numerical example which consists of solving the forward displacement, velocity, and acceleration analyses of an SHRM built with four modules is included. The proposed hyper-redundant manipulator consists of an optional number of identical redundant parallel manipulators with autonomous motions assembled in series connection; see Fig. 1, where n is called the output platform, k and k \u2212 1 are two consecutive platforms, and 0 is the fixed platform. The base module of the proposed SHRM consists of two platforms, for instance labeled 1 and 0, connected each other by means of a spherical parallel manipulator (SPM) and a redundant planar parallel manipulator (RPPM). Evidently, the position and orientation of body 1 with respect to body 0 are controlled independently by means of the SPM and the RPPM, respectively. Hence, the base module can be considered as a decoupled robot", + " Finally, once the reduced acceleration state of the output platform with respect to the fixed platform is computed, six-dimensional vector 0 An O = [0\u03c9\u0307n; 0an O \u2212 0\u03c9n \u00d7 0vn O]T, the linear acceleration of the center of the output platform is obtained combining elementary kinematics and the properties of a helicoidal vector field as follows: 0an o = D ( 0 An O ) + 0\u03c9\u0307n \u00d7 ro/O + 0\u03c9n \u00d7 ( 0vn o ) . (32) In this section the kinematics of an SHRM built with four modules is presented. The parameters, using SI units, for the base module and its home position are listed in Table I. The home position of the base module is repeated for each module by increasing in nmodh(nmod = 2, 3, 4) times the coordinate along the Y-axis. Keeping this in mind, the resulting home position of the SHRM is depicted in Fig. 1. Furthermore, the actuable joints are affected by periodical functions of the form C j i sin(t), where C j i is the coefficient associated with the ith generalized coordinate in the jth module; these coefficients are given in Table II. With these data, the exercise consists of finding the time history of the angular and linear kinematic properties of the center of the output platform with respect to the fixed platform, expressed in the reference frame OXYZ. As expected, due to the periodical functions, threedimensional undulatory motions are generated in the SHRM, animations in mpg format are available for this example, and the most representative numerical results obtained for it are given in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001873_iecon.2005.1568952-Figure4-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001873_iecon.2005.1568952-Figure4-1.png", + "caption": "Fig. 4 The process parameters of the proposed RT system", + "texts": [ + " The aim of the layer thickness in the Z direction is 1mm. S/N = \u239f \u239f \u23a0 \u239e \u239c \u239c \u239d \u239b 2 2 10 \u03b4 ylog10 (6) \u2211 = = n i iy n y 1 1 (7) 1 )( 1 2 2 \u2212 \u2212 = \u2211 = n yy n i i \u03b4 (8) where y is the mean of samples, 2\u03b4 is the variance of samples, y is the samples, and n is the number of samples. There are many process parameters of laser cladding need to be considered, such as laser power, scanning speed, nozzle offset ( X\u2206 ), powder feed rate, laser spot size, and tool path offset. Several process parameters of the proposed RT system are shown in Fig. 4. X\u2206 is the distance between the center of the spot size and the intersection point of the axis of powder stream and substrate surface. X\u2206 is defined to be positive when the intersection point of the axis of powder stream and laser beam was beneath the center in the spot size, otherwise it was negative. Table I shows the variables and levels selected for this study. An L27 orthogonal array with 6 columns and 27 rows is chosen to account for the processes parameters and their levels. The physical properties of nickel-based alloy we adopted are good for the proposed RT system" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001118_0471758159.ch5-Figure5.36-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001118_0471758159.ch5-Figure5.36-1.png", + "caption": "FIGURE 5.36 Modeling the effect of a common-mode choke on (a) the currents of a twowire line, (b) the differential-mode components, and (c) the common-mode components.", + "texts": [ + " For example, we will find that microamperes of commonmode current will produce the same level of radiated electric field as tens of milliamperes of differential-mode current! Common-mode currents are not intended to be present on the conductors of an electronic system, but nevertheless are present in all practical systems. Because of their considerable potential for producing radiated electric fields, we must determine a method for reducing them. One of the most effective methods for reducing common-mode currents is with common-mode chokes. A pair of wires carrying currents I\u03021 and I\u03022 are wound around a ferromagnetic core as shown in Fig. 5.36a. Note the directions of the windings. The equivalent circuit is also shown. Here we assume that the windings are identical, such that L1 \u00bc L2 \u00bc L. In order to investigate the effect of the core on blocking the common-mode current, we calculate the impedance of one winding: Z\u03021 \u00bc V\u03021 I\u03021 \u00bc pLI\u03021 \u00fe pMI\u03022 I\u03021 (5:31) Now let us investigate the contribution to the series impedance due to each component of the current. First let us consider common-mode currents in which I\u03021 \u00bc I\u0302C and I\u03022 \u00bc I\u0302C. Substituting into (5", + "31) gives Z\u0302CM \u00bc p(L\u00feM) (5:32) The contribution to the series impedance due to differential-mode currents where I\u03021 \u00bc I\u0302D and I\u03022 \u00bc I\u0302D is Z\u0302DM \u00bc p(L M) (5:33) If the windings are symmetric and all the flux remains in the core, i.e., the flux of one winding completely links the other winding, then L \u00bc M and Z\u0302DM \u00bc 0! Thus in the ideal case where L \u00bc M a common-mode choke has no effect on differential-mode currents, but selectively places an inductance (impedance) 2L in series with the two conductors to common-mode currents. These notions are illustrated in Fig. 5.36. In addition to selectively placing inductors L\u00feM in series with the commonmode currents, use of ferrite cores places a frequency-dependent resistance, R( f ), in series with the common-mode currents as well. This resistance becomes dominant at the higher frequencies as was the case for a ferrite bead in the previous section. Hence common-mode currents not only are blocked but also have their energy dissipated in the R( f ). Thus common-mode chokes can be effective in blocking and dissipating common-mode currents" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000031_s0307-904x(81)80025-x-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000031_s0307-904x(81)80025-x-Figure1-1.png", + "caption": "Figure 1", + "texts": [ + " An energetic approach permits an overview view of the object being investigation. Moreover, energetic functionals can be appropriately extended by the addition of further terms. Linearization makes it possible to use classical variational methods, which yield simple solutions in a closed form. We divide the wall of a real tyre into two domain's: an elastic layer of tread (protector) and the remaining rein- forced part, i.e. carcass, sidewalls and belt. We consider the reinforced part first (see Figure 1). V o l u m e o f m o d e l Let the axis of revolution be z and let the cylindrical coordinates be r, ~, z. The belt is considered to be fully flexible in the radial direction and almost rigid in the circumferential direction. We identify the path of cord in the carcass with the meridian curve, f. Let the lower boundary of the belt cross-section b be approximated by the circle: b2(r) + (r - a + rg) 2 = r~ The air compressed in the tyre cavity always seeks to expand. At the same time, however, the length of the meridian curve (cord) is fixed. Thus, with respect to the obvious symmetry, the meridian curve f of the free zone (sidewall) is given by solving the following isoperimetric problem :7 r! r ! (2.f f(r)drlf ro ro air volume meridian length Considering the boundary conditions: f (r , ) = b(rO f ( r O = b'(rO its solution can be obtained in the following form: r f(r; Ol, X) = zl + l\" tan0(r; 01, X) dr (1) r! where rl, zl can be expressed by means of a, rg, 01 (Figure 1) and: sin0(r; 0~, X) = (1 - r2/r~)/(2X) - O~ The parameter X i's defined as follows: X = 1/Ix(r 0 rl I, wher~ x(r l ) = [d sin0(r; 01, X)/dr]r=r, denotes the curvature of the meridian curve (.f) at the point (&, zt). 0307-904X/811060422-06/$02.00 The integral (I) may be expressed by means of.elementary functions (for X = 0.5/(I + sin00) or elliptic integrals of the first and second kinds F, E. In the most frequent case X < 0.5[(1 + sin01) is: f(r; O1, X) = zl + 2r,x/~/k [E(k, cO(r)) - E ( k , cO(r1)) - (1 - k:/2)(F(k, cO(r)) - F(k, cO(rl)))] where: k = 2[~/l/X + 2 - 2 sin01 and cO(t) = arcsin (x/l - tU/k) The search of the function fwhen a group of five quantities a, L, rg, rB, zB (Figure 1) is given will be shortly called problem (B) (basic problem of meridian theory) and symbolically written as follows: (a, L, rg, rB, ZB) ~ (0,, X) (B) The numerical solution of the problem (B), which leads to two nonlinear simultaneous equations, has been treated elsewhere.7,~ For calculating the energy of the model the volume: r I a V=4rr frf(r)dr+frb(r)dr r B r 1 is needed. The second integral is an elementary one. The first can be transferred by integration by parts as follows: r I r~ f r f(r)dr= r~zl--r~zB-- f r:tanO dr 2 rB rB However, the integral on the right hand side is of the same kind as the analytic expression (1), i", + " To take account of this and to avoid the difficult problem of determining the stiffness coefficient ku(w), we simply put k u ( w ) = 1.5p(w) co: + ku(O) and ku(co) = 1.Sp(co) co2 + ku(0) and Ep(co) = Ep(O)(1 + co:1400) On the other hand, the belt tension increments produce distinct changes o f the lateral and torsional stiffnesses. Tiffs is demonstrated for the 175/70 R 13 tyre in Figure 6. O n t y r e g e o m e t r y Let the equator radius o f the 185 SR 14 tyre vary and let the belt width be leg = 2Lx (Figure 1). The corresponding courses o f the lateral stiffness C L = [dF(w)/dw]w=o and the average radial stiffness C R = Flu, where F = 4.9 kN and u is the corresponding radial deflection, are shown in Fig~tre 7. Both functions CL and C R attain their maxima at aL ~ 310 mm and aR ~ 295 mm. The optimum geometry (with respect to these two stiffnesses only!) is characterized by the compromise radius ao, aR < ao < aL, that may be constructed, for example, by means o f game theory3 a R e f e r e n c e s 1 Frank, F" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001783_wcica.2006.1713764-Figure3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001783_wcica.2006.1713764-Figure3-1.png", + "caption": "Fig. 3 Coordinates and Frames", + "texts": [ + " During flight test, the information about helicopter is periodically sampled and sent via the RS-232 link to the flight control computer, which in turn is calculated and sent by appropriate pulse width modulation signals to the servos. In additional, the data is also sent to the earth work station by wireless network to complete more complex assignments and supervise. There are two coordinates that we often use: the reference inertial coordinate and body coordinate attached to the helicopter, as shown in Fig. 3 where cm is the center of the mass. The transformation between the two coordinates is given by a homogeneous transformation matrix IBM which represents the relative orientation between the two coordinates. The matrix can be expressed as following [18] \u2212\u22c5 \u2212 \u22c5 \u2212 = \u03c6\u03c6 \u03c6\u03c6 \u03b8\u03b8 \u03b8\u03b8 \u03c8\u03c8 \u03c8\u03c8 cossin0 sincos0 001 cos0sin 010 sin0cos 100 0cossin 0sincos IBM , (1) where \u03c6 , \u03b8 and \u03c8 are the Euler angles (fuselage attitude angles) of the helicopter with respect to the inertial frame. Due to orthogonality of the coordinate transformation, the inversion of the rotation matrix is equal to the transpose of the matrix", + " From the above equations, the neural network architecture can be expressed in matrix form as )(XWU NN \u03c3= , (27) Based on the analysis of the inversion by Lyapunov stability method [17], a bounded weight adaptive update law is given by [15], [16], [17] ][ WerW w \u03bb\u03c3 +\u0393\u2212= , (28) where TT Pber )(= , nnP 22 \u00d7\u211c\u2208 is the positive definite solution to the Lyapunov Equation 0=++ QPAPAT . Usually, Q is chosen as IQ \u03b1= ( +\u2208 R\u03b1 ). w\u0393 are network learning rate, \u03bb is the e-modification gain for adaptive control theory. 0>\u0393w and 0>\u03bb guarantees that tracking error e and neural network weight W are uniformly ultimately bounded. Additional assumption is that all signals in the closed loop system including plant states are bounded. All details of the proof of boundedness and relations of all variables can be found in [17]. . FLIGHT TEST RESULTS The hybrid control system shown in Fig. 3 was implemented on the Raptor 60 helicopter. The pitch channel is most complicated of the three channels: pitch, roll and yaw. To be simplified, only the control of pitch channel is given. pK and DK are chosen based on a natural frequency of 1.0 srad / and a damping ratio of 0.6. The network learning rate w\u0393 is chosen as I3 , and gain 3.0=\u03bb . During the experiment, the helicopter is taken off on the platform by human maneuver, and controlled to work in hover condition and maintain a stable height" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003812_003-Figure3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003812_003-Figure3-1.png", + "caption": "Figure 3. Shot put kinematics.", + "texts": [ + "26 kg for men, 4 kg for women) the aerodynamic effects (such as drag) can be neglected. An interesting exercise is to calculate the optimum angle of release in order to maximize the range. Of course, if the launch was made from ground level, it is well known that the optimum angle would be 45\u25e6. The fact that the launch is made from a height above the ground though will alter the value of this angle. Let it be assumed that the launch angle is \u03b8 and that the velocity is decomposed into a horizontal and a vertical component (see figure 3). We then have x(t) = u cos \u03b8 t z(t) = hl + u sin \u03b8 t \u2212 1 2 gt2 (12) where x(t) is the horizontal distance travelled, z(t) is the vertical distance from the ground and hl is the launch height from the ground. When the shot hits the ground after time tmax, the horizontal distance travelled is equal to the range R. R = utmax cos \u03b8 0 = hl + utmax sin \u03b8 \u2212 1 2 gt2 max. From the above equations we obtain R = u2 sin 2\u03b8 2g \u23a1 \u23a31 + \u221a( 1 + 2ghl u2 sin2 \u03b8 )\u23a4 \u23a6 . (13) 596 P H Y S I C S E D U C A T I O N November 2010 As an exercise in differentiation it can be proved (see [10]), that the optimum angle for maximum range is given by sin \u03b8max = \u221a 1 2(1 + ghl u2 ) " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001363_agronj1972.00021962006400040028x-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001363_agronj1972.00021962006400040028x-Figure2-1.png", + "caption": "Fig. 2. Construction detail of Thatchmeter 11. All dimensions are in centimeters.", + "texts": [], + "surrounding_texts": [ + "The object of this study was to develop a method for measuring the variation in compressibility of thatch on bermudagrass experimental and golf course greens and to evaluate the data as an indicator of grass growth and thatch development. A \"thatchmeter\" was designed to rapidly determine differences in compression between a bearing pressure of 7.3 g/cm2 and one of 570 g/cma, the former less than that of a golf ball, and the latter, estimated from the weight of an average man. Regressions of compressibility on rate of grass growth and on thatch weight and depth were statistically significant (.001). Compressibility averaged between 6.1 and 11.5 mm per green for 24 golf greens under daily play. A 10-reading average per plot or green can be obtained in 5 mill and requires little professional supervision. The value of the procedure as a research tool is apparent, but specific ranges for acceptable playing quality of golf greens, between excessive hardness and excessive thatch, must be established individually for various grass varieties and environments.\nAdditional key word: Thatchmeter.\n'T1 HATCH has been described by the United States A Golf Association (3) as an accumulation at the soil surface of dead but undecomposed stems and leaves. At the other extreme, Thompson (2) states that thatch is a tightly intermingled layer of living stems, leaves, and stolons that develop between the green vegetation and the soil surface. Musser (1) considers thatch to be dead and dying organic materials in this zone.\nThe term \"thatch,\" as used in this report, refers to that portion of living and dead grass stems, stolons, and organic debris that lies above the mineral soil line. The dividing point between thatch and soil often\n1 Contribution of the Department of Soil Science, University of Florida, Gainesville 32601. Journal Series 4178. Received Nov. 8, 1971.\n2 Soil Chemist.\nis visually indefinite, especially where aerification or topdressing wth soil has been practiced or where rhizomes permeate and massively heave the immediate surface of the soil.\nExcessive thatch development on golf greens is a major detriment to their playing quality. It also makes extended fertility trials difficult to evaluate. A reliable method for rapid estimation of thatch development could be of considerable value in timing of maintenance practices involving periodic removal of thatch, or of topdressing with soil to stabilize it.\nThe fact that the degree of springiness of turf has been recognized as an indicator of thatch development led to the concept of measuring thatch by its compressibility. The object of this work was to develop and to evaluate such a procedure.\nMATERIALS AND METHODS Thatchmeter I\nThis instrument (not illustrated) consisted of a light base carrying a lever that could be variably loaded over a vertical cylinder independent of the base. The base had a contact area of 826 cm2 and a bearing pressure of 5.2 g/cm2. The depression difference between the light base load and the heavy cylinder load was read by a pointer installed above the weighted lever in a manner that magnified depression difference by 10. Thatchmeter I was used to establish curves of compressibility for two areas of bermudagrass (Cynodon dactylon var. Tifdwarf), one Verticut for thatch removal the previous fall, and the other unmodified.\nThatchmeter II As a result of experience with model I it was decided to\nconstruct a system employing two fixed loads, one that would not exceed that of a golf ball, and the other approximating the compression applied by a man. A golf ball was observed to depress bermudagrass (freshly mowed to 6-mm height) by 6.5 mm. An 83-kg man's weight was calculated to exert a pressure of 570 g/cm2, based on an estimated bearing area of 145 cm2 for the ball and heel of one shoe.", + "502 AGRONOMY JOURNAL, VOL. 64, JULY-AUGUST 1972\nModel 11, shown in Fig. 1 and 2, has a base of 524 cma with bearing pressure of 7.3 g/cma. The base, less the high-pressure cylinder, depressed freshly mowed grass by only 5 mm. The high-pressure cylinder end is 7.92 cm in area, loaded to 570 g/(m2, with 10 x multiplication of compression readings. The bare consists of three layers of exterior plywood, glued together. Otherwise, soft-wood stock is used. The line activating the pu ley lever (L) is 9-kg-test plastic-coated woven fishing-leader wile. I t is clamped to the pulley under a screw head to permit adjustment of the pulley lever to a horizontal position at apprctximately 1-cm projection of the compression cylinder. Pulley ant1 pointer (P) are clamped between a jam nut on a smoothshank bolt and a wing nut that can be loosened for 0 adjustmei t of the pointer with the instrument on a flat, hard surface. The reading scale is flexible tape with divisions 1 cm apzrt, glued to the top surface of the arc.\nThe compression weight (W) consists of molded lead with a 9-cin section of 3.175-cm outside diameter brass lead-filled drain pipe extending 8 cm. Total weight of the lead plus pipe is 4,514 g.\nEtultiple readings were made with Thatchmeter I1 on 40 individual plots of a slow-release N fertilization test maintained by twice-weekly mowings at 8 mm. The area had been verticut 60 days piior to establishment of the fertilizer test and 102 days before making the compression readings. All clippings were col ected and accumulated, and total yields by plot were compared to compression values.\nA second type of evaluation was made by taking individual coripression readings and associated sod cores 3.6 cm in diameter at 33 locations on greens plots representing a range of conditioils. The compression readings were individually plotted against the observed depth of thatch above the mineral soil line.\nA third evaluation was made by taking 31 individual sod cores immediately following compression readings, truncating the core at the depth of visible compression achieved by firm pressure between thumb and forefinger (approximately 600 g/cma) and then washing the compressible portion free of < 2-mm material. The compression readings were plotted against the 65-C dry weight of washed thatch material.\nBecause it was apparent that in practice several readings of the thatchmeter should be averaged for a given golf green or turf plot, 10 areas with predetermined approximate com,pression values ranging between 4 and 13 mm were sampled in detail for weight of thatch to the soil line. Eight-core composites of sod, along with specifically related compression readings, were taken of each area. The cores were washed free of < 2-mm material and the average compression readings plotted against the composite weight of thatch material for a given area.\nTo determine the characteristics of variation that might be expected under practical playing conditions 10 comprtasion readings were made on each of 21 regular greens and 3 practice greens of 2 local golf courses.\nRESULTS AND DISCUSSION\nThe ratios between compression readings made with Thatchmeter I for areas of high vs low thatch at comparable compression weights (Fig. 3) showed that sensitivity of the procedure apparently begins to drop significantly at about 600 g/cm2.\nData obtained by means of Thatchmeter I1 showed a highly significant (.001) and positive regression of grass blade growth continuously removed by mowing, on development of sub-harvest-level of thatch (Fig;. 4). They suggest the value of the procedure for periodic observation of relative growth response to various treatments, without the continuous and exacting job of harvesting for yield weights.", + "VOLK: MEASURE OF GROWTH ON BERMUDAGRASS GREENS 505\nData in Fig. 5, 6, and 7 show that there are significant (all .001) and positive l\u2019egressions of mm compressibility on measured thickness of thatch; on the weight of thatch > 2 mm in the compressible layer; and on the weight of thatch > 2 mm in the entire thatch depth to the mineral soil line. Figure 7 is of particular interest, because it introduces the real significance of average data from multiple readings, where compression readings a~e regressed on relatively clean thatch material taken to the most readily recognized lower boundary of thatch. It suggests that further evaluation of the procedure for routine usage in thatch control on a practical basis should involve determination of these two factors, i.e., compressibility and weight of thatch > 2 mm above soil line.\nThe potential for the practical use of thatch compressibility as an aid in maintenance of closely mowed turf is suggested by the data in Table 1. Means of 10 readings per green varied from 6.1 to 11.5 ram. Greens that still showed evidence of verticutting averaged 6.7 mm compression. Standard errors recorded in Table 1 indicate that 10 readings per green gave a logical mean value.\nline." + ] + }, + { + "image_filename": "designv11_69_0000143_iros.2000.895222-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000143_iros.2000.895222-Figure2-1.png", + "caption": "Figure 2: Planar 3-link manipulator and an obstacle", + "texts": [ + " If not, then set k = k + 1 and return to Step 3. 4 Simulation Example In this section, the proposed algorithm is demonstrated for computing the collision-free quasi-optimal trajectory of the 3-link planar free-joint manipulator whose first and second joint is actuated and third joint is not actuated. We consider here the problem that finds a trajectory from the initial configuration 8 = ( 0 1 , 8 2 , 8 3 ) = (Oo ,Oo ,Oo) to the desired configuration 8 = (90\",45\",45\") at t f = 3.0 [sec] in the presence of a circular obstacle (see Fig.2). Note that it has already been proved that the free-joint manipulator is small time locally controllable [12]. This means that the free-joint manipulator can follow any paths that can be realized by the corresponding 3-link planar fully-actuated manipulator with any accuracy. Accordingly, if the collision-free trajectories of the 3- link fully-actuated manipulator are exist, the collisionfree trajectories of the 3-link free-joint manipulator are also exist. The dynamics of the manipulator is represented by where M E R3x3 is the inertia matrix, c E R3 is the centrifugal and Coriolis vector, and q , 7 2 are the driving torques of the first and second joint respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003807_amr.139-141.1079-Figure5-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003807_amr.139-141.1079-Figure5-1.png", + "caption": "Fig. 5 Distribution of contact stressfor direct busbar roller", + "texts": [ + " To guarantee the convergence of the computation and get accurate enough results, we will increase the load step by step. Analysis of contact stress on the design of different rollers with varied convexity The design of different rollers with varied convexity gives significant influence to the quantification and distribution of the contact stress between the raceway and roller. It can also change the elastic deformation. Furthermore, it is playing a critical role in determining the load capability and the life of the roller bearing. Analysis of the Contact Stress according to Direct Busbar Roller. Fig.5 used finite element method, for direct busbar roller, the distribution of contact stress between inside ferrule and roller. For direct busbar roller, the curve of distribution of contact stress between inside ferrule kinematic pair and roller in the direction of roller is shown in Fig.6. From Fig. 5 and Fig. 6, we can see it clearly that by taking the design of direct busbar the stress distribution of the contact area between roller and the inside ferrule is nonuniform when we consider the roller with load. The stresses blow up at two end points, give singular distribution, and provide so called end effect. This kind of effect is adverse to the distribution of stress on the contact area, the stress situation under the contact area, the load capability, and the life of the roller bearing. At this time, the maximal contact stress should be 3920MPa, appears at the end point of the tapered roller busbar", + " (2)When R<5000mm, the maximal contact stress takes place in the middle of tapered roller busbar, when R decreases, the maximal contact stress increases. (3)The optimal R is around 5000mm~7000mm, in this interval, we have the lowest maximal contact stress and the stress is uniformly distributed. Comparison of the Bearing Life before and after Optimization. Still consider tapered roller bearing 3811/750/HC as example, according to the traditional direct busbar tapered roller design, we can see the obvious stress concentration at end points from Fig.5, with maximal contact stress 3920MPa. The point with maximal stress is accidentally in he minimum wall thickness of the inner ferrule, and then it follows the destroy of the contact area. By experiment, the life of this bearing is only 32 days. After optimizing as we mentioned above, by choosing ARC convex busbar tapered roller design, we can see from Fig. 12, the maximal contact stress 3000MPa, and the stress distribution is uniform. By experiment, the average life of the bearings after optimization is only 65 days, and the longest one is 102 days" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002451_sice.2007.4421349-Figure4-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002451_sice.2007.4421349-Figure4-1.png", + "caption": "Fig. 4 BLDC Motor and linear type Hall-effect sensor", + "texts": [ + " 3, T method is useful in high speed range. However, this method has a delay in slow speed range because the frequency of the pulses is proportional to the velocity and the velocity information can hardly be obtained in the low velocity range. So this method can not be used for electro-mechanical fin actuator. As mentioned in chapter 2.1, it is inevitable to reduce the delay in low velocity range for electro-mechanical fin actuator. Hence, linear type Hall-effect sensor is used for reducing delay. Fig. 4 shows BLDC motor and linear type Hall-effect sensor used in experiment. As shown in Fig. 5, unlike the square wave form of latched type Hall-effect sensor, the linear type Hall-effect sensor outputs the voltages in the forms of sine waves by rotating of a motor. Among the Hall signals(hA, hB, hC) in the forms of sine, the coordinate values of phase A(x1, y1) and that of phase B(x2, y2) are obtained by using the two signals(hA, hB). As shown in Fig. 6, Suppose that phase A is put as the coordinate on y axis, and phase B represented a 120-degree phase difference compared to phase A electrically" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000060_b104749n-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000060_b104749n-Figure1-1.png", + "caption": "Fig. 1 Biosensor assembly. a, Amperometric Clark electrode; b, dielectric; c, Ag/AgCl anode; d, Pt cathode; e, PTFE cap; f, gas-permeable membrane; g, PTFE O-ring; h, dialysis membrane; i, k-carrageenan membrane with immobilized enzyme; l, filling solution.", + "texts": [ + " An amperometric electrode able to measure dissolved oxygen concentration is screwed on to a PTFE cap open at one end, with the other end closed by a gas permeable membrane. The PTFE cap allows the electrode to be used also when dipped into organic solvents. Before fixing the PTFE cap on to the amperometric oxygen electrode it is filled with 2 ml of internal solution. The internal solution is prepared by weighing 2.849 g of KH2PO4 and 3.925 g of KCl, dissolving them in 500 ml of distilled water and diluting with Na2HPO4 until a pH of 6.6 is obtained. As shown in Fig. 1, the disk with the enzyme, prepared as described in the preceding section, is sandwiched between the gas-permeable membrane and a dialysis membrane. The whole system, consisting of the two membranes and the gel-like kcarageenan membrane with the enzyme immobilized on it, as described above, is then secured to the PTFE cap by means of an O-ring made of the same material. The biosensor is then dipped under constant stirring into a measurement cell thermostated at 20 \u00b0C containing the selected solvent and then coupled to a potentiostat" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001218_detc2006-99153-Figure3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001218_detc2006-99153-Figure3-1.png", + "caption": "Figure 3. Screw motion.", + "texts": [ + "org/about-asme/terms-of-use Down and the moment vector of the line is written as l0 = p1\u00d7 l The six-tuple coordinates (l, l0) = (l1, l2, l3; l0 1 , l0 2 , l0 3) are the normalized Plu\u0308cker coordinates of a line. Using dual numbers [17, 18], they can be written as l\u0302 = l+ \u03b5 l0 where \u03b52 = 0 is the dual unit. The direction vector l and moment vector l0 satisfy the Plu\u0308cker relation l \u00b7 l0 = 0 The condition for two intersecting lines is l\u0302a = (la, l0a) and l\u0302b = (lb, l0b) la \u00b7 l0b + l0a \u00b7 la = 0 A screw is a geometric element characterized by a line and a pitch. This element is used to describe the helical motion of a rigid body between two positions (see fig. 3). The screw parameters are line coordinates of the screw l\u0302 = (l, l0), angle of rotation around the screw \u03b8, and translation distance along the screw d as shown in fig 3. The pitch of a screw p is defined as the distance of the body translated along the screw divided by the angle of the body rotated around the screw, namely: p = d \u03b8 (1) Screw s\u0302 can be written as a six dimensional vector s\u0302 = (s,s0) = (l, l0 + p l) 3 loaded From: http://proceedings.asmedigitalcollection.asme.org/pdfaccess.ashx?ur Considering a pair of skew lines in space, the Plu\u0308cker coordinates of the common perpendicular between these two lines are derived in the Appedix. Let l\u0302a = (la, l0a) and l\u0302b = (lb, l0b) be a pair of lines given in their Plu\u0308cker coordinates that belong to the rigid body before a displacement" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000006_jsvi.2002.5135-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000006_jsvi.2002.5135-Figure2-1.png", + "caption": "Figure 2. Finite segment model of beam like structure: (a) continuous flexible system; (b) discrete model by overlapped finite segments and elastic beam element.", + "texts": [ + " Section 3.2 introduces the mathematical model of the locking impacts in the joints and section 3.3 establishes the dynamic equations of motion of the flexible deployment system. The modelling of the flexible beams includes two parts, namely the inertial modelling with the finite segment and the elastic modelling with the finite element. 3.1.1. The inertia modelling In the finite segment approach, the flexible beam is discretized into a number of rigid segments connected by beam elements, as shown in Figure 2. The virtual work equation for the discrete flexible beam s can be expressed as dqTs Ms .qs Qs\u00bd \u00bc 0; \u00f01\u00de where Ms is the mass matrix, qs is the generalized co-ordinate vector and Qs is the corresponding generalized force vector, which are defined as follows: qs \u00bc qTs1; q T s2; . . . ; q T sn ; Ms \u00bc diag Ms1;Ms2; . . . ; Msn\u00f0 \u00de; Qs \u00bc QT s1;Q T s2; . . . ; Q T sn ; \u00f02\u00de where qsi \u00f0i \u00bc 1; 2 ; n\u00de is the generalized co-ordinate of the ith segment. Msi is the generalized mass matrix of the ith segment Msi \u00bc diag\u00f0msiI3; J\u2019si\u00de; i \u00bc 1; 2; ; n; \u00f03\u00de where msi is the mass of the segment i; J\u2019si the inertia tensor of the segment about its center of mass and I3 is the identity matrix of 3 3" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003976_icma.2009.5244953-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003976_icma.2009.5244953-Figure2-1.png", + "caption": "Fig. 2. ASTERISK Hanging onto Wall", + "texts": [ + " For the grid wall locomotion on various situations, the paper proposes an optimization method for the omni directional gait of ASTERISK by using Genetic Algorithm (GA). Grid-like structure consists of many bars assembled in a matrix in a plane. Each foot of the robot has a hook for hanging on the bar. The stable hooking and the effective locomotion are evaluated so that the optimal gate is planned . The acquired gate plans are implemented to ASTERISK and the experimental results show the feasibility of our proposed grid wall walking. In this study, we separate ASTERISK's six legs into two tripods. We name one of them \"APod\" and the other one \"VPod\". Fig.2 shows ASTERISK on a grid wall. While hanging, three toes of a tripod shape an isosceles triangle. The circle-shaped toes belong to APod, and the star-shaped ones belong to VPod. We constraint APod and VPod to have the same shape. We name the height of the triangle \"Height\", and the width \"Width\". To measure the positions of toes, we set a right-hand system on the center of the body, whose x axis is directed to the groin of leg I, and whose z-axis to the wall. We represent the positions of APod and VPod by At last, we describe the walking gait" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001979_50037-5-Figure7.1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001979_50037-5-Figure7.1-1.png", + "caption": "Figure 7.1 Illustration of the problem.", + "texts": [ + " They sum to approximately linear pressure distribution and the fact that the pressure falls to zero at the edge of the contact ensures that the surfaces do not interfere outside the contact area. The three-dimensional (3D) equivalent of overlapping triangular elements is overlapping hexagonal pyramids on an equilateral triangular grid. An authoritative treatment of contact problems can be found in the monograph by Johnson [1]. 7.2 Example problems 7.2.1 Pin-in-hole interference fit One end of a steel pin is rigidly fixed to the solid plate while its other end is force fitted to the steel arm. The configuration is shown in Figure 7.1. This is a 3D analysis but because of the inherent symmetry of the model, analysis will be carried out for a quarter-symmetry model only. There are two objectives of the analysis. The first is to observe the force fit stresses of the pin, which is pushed into the arm\u2019s hole with geometric interference. The second is to find out stresses, contact pressures, and reaction forces due to a torque applied to the arm (force acting at the arm\u2019s end) and causing rotation of the arm. Stresses resulting from shearing of the pin and bending of the pin will be neglected purposefully" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001199_0005-2736(71)90066-6-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001199_0005-2736(71)90066-6-Figure1-1.png", + "caption": "Fig. 1. Transport of L-[14C] arginlne in wild type and Pm-AB in the presence and absence of cyclobeximide (actidione) (100 #g/ml). L-Arginlne is at an external concentration of 0.1 mM and a specific activity of 0.01 #C/0.1 #mole.", + "texts": [ + " *Presented in part at the 43rd and 44th Annual Meeting of the S.E. Branch of the American Society of Microbiology, 1969. Biochim. Biophys. Acta, 241 (1971) 677-681 Techniques employed. The methods and materials employed in culture maintenance, transport experiments, and extraction and chromatography of label have been previously described 7 . This study was undertaken to examine the apparent increase in the velocity o f L-arginine transport, in P m - AB # 10 (Pro- AB), as a function of time of incubation in the presence of L-arginine 7 . As shown in Fig. 1, the increase in rate o f L-arginine transport can be blocked by the addition of cycioheximide. This antibiotic has been shown to be a potent inhibitor of protein synthesis in Neurospora. In order to more closely examine the velocity o f L-arginine transport, cells were preincubated with or without 0.1 mM L-arginine in 1X Vogel's salts a . At appropriate intervals, aliquots were removed, f'dtered, washed and resuspended in 1X Vogel's salts plus L-[14C] arginine (0.1 mM). Thereafter, short-term uptakes (< 16 min) provided an estimate of the velocity (in #moles/min per mg dry wt" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002083_detc2007-34911-Figure9-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002083_detc2007-34911-Figure9-1.png", + "caption": "Figure 9 Road profile.", + "texts": [ + " However, it is clear that such a shape function is unable to accurately describe the exact mode shapes of a tire in contact, and this result indicates the importance of detailed modeling of tire/road contact interface. In the numerical tire model developed in this investigation, since tire/road contact is modeled using multiple contacts (contact nodes) defined along the elastic belt elements, accurate vibration characteristics of the tire can be obtained as demonstrated in Table 2. Dynamic Response In this example, the dynamic response of the tire to an uneven road is discussed. The road profile is given as shown in Fig. 9. The height of step (B) is assumed to be 30 mm, while the amplitude ( rH ) and wavelength ( rL ) of the sinusoidal road profile are, respectively, assumed to be 10 mm and 20 mm. The downward vertical force of 3000 N is applied at the center of the rim and the forward velocity is assumed to be 5 m/s (18 km/h). Figure 10 shows the vertical Downloaded From: http://proceedings.asmedigitalcollection.asme.org/pdfaccess.ashx?u displacement of the rim. As can be seen from this figure, the vertical vibration excited by the step is quickly damped out, while the small vibration resulting from the sinusoidal road unevenness is continuously transmitted to the rim" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003893_iceee.2010.5661528-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003893_iceee.2010.5661528-Figure1-1.png", + "caption": "Figure 1. 1. Whole model of motor Figure 1.2. The inner model of motor", + "texts": [ + " However, the FEM overcome the disadvantage of different from the actual and ideal model. The noise could be calculated precisely and motor\u2019s motion and structure also could be simulated actually by precise model concerning any certain motor.[5][6] In this paper, an AFSMPMSM\u2019s vibration and noise in no load situation are calculated by using FEM. And it is the theme of finding the source of noise, and the base is built to control the noise. Some parameters of this AFSMPMSM is followed by Table .The whole finite element model of motor is shown as following Fig.1.1, and motor\u2019s inner model is shown in Fig.1.2. TABLE I. SOME PARAMETERS OF 5KW AFSMPMSM Rated power 5kW Num of pole 22 Num of stator teeth 24 Frequency 82.5Hz Phase 6 II. THE ANALYSIS AND CALCULATION OF The gap exist the basic wave field and a series harmonic wave fields. And the tangency electromagnet force waves are produced by interaction of these fields. Besides, the axial electromagnet force waves, which are changeable with time and space, also are made by these fields. And stator became distortion and the vibration and noise are also produced because of these axial electromagnetic force waves acting at stator\u2019s core" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000696_j.mechatronics.2004.06.009-Figure3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000696_j.mechatronics.2004.06.009-Figure3-1.png", + "caption": "Fig. 3. Field acting on control points on a manipulator link.", + "texts": [ + " There are two points that can easily justify these empirical values. First, these constant values are determined considering the worst case scenario that the current beam analysis algorithm can achieve with very long link lengths and through narrow passages. Second, there is no need to reset these values. Once they are set as above, they work for other cases. As the link lengths get shorter, lu, lv and glim would be relatively short and there would be more space to manoeuvre since they are proportional to the link lengths. Fig. 3 shows a link surrounded by three obstacles. The field lines are shown on the link which needs to manoeuvre around the obstacle 1. There is free space around the link except for the front where it is very close to the obstacle 1. The field lines acted on near the tip of the link give a false idea of safety since they suggest that the link should rotate counter-clockwise. The control point 3 is about to change direction as the link approaches the obstacle 1. The control points 1 and 2 suggest that the link should rotate clockwise direction" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000662_s00366-004-0269-3-Figure8-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000662_s00366-004-0269-3-Figure8-1.png", + "caption": "Fig. 8 a Linkage of Fig. 7 where link 3 is fixed; b The corresponding image velocity diagram; c The final calculation of the relative velocity in the link of the original mechanism", + "texts": [ + " In the following example the transformation of the linkage makes it decomposable to Assur groups [10] of a lower class, thus reducing the complexity of the analysis procedure. Consider the linkage of Fig. 7, known in the literature as a Stephenson linkage type III [6]. The linkage has one degree of mobility and its structure is composed of the driving link 1 and a tetrad 2,3,4,5,6,7,8,9. Performing step 1 of algorithm 3 to this linkage, one may choose link 3 to be the link s. Applying step 2 yields the linkage shown in Ffig. 8. As a result of the transformation, the joints connected to link 3 become fixed supports, while the joints that were originally fixed supports become the joints connecting a new link, 0, to the rest of the mechanism. The linkage of Fig. 8 is a simple linkage that can be separated into a driving link and two dyads. The analysis of such a linkage is rather straightforward and canFig. 5 The first step in the transformation process be done for example by the image velocity method, as is shown in Fig. 8b. Now, one can employ the solution obtained in Fig. 8b and Eq. 6 to evaluate the linear velocities of the links in the original linkage. Figure 8c shows an example calculation of the transmission ratio between the angular velocities of link 6 and the driving link of the original linkage. Transferring the method to trusses As was explained in the section \u2018\u2018The duality relation between linkages and trusses\u2019\u2019, in 2001 a duality relation between plane linkages and plane trusses [8] has been established, yielding the result that for each truss there is a corresponding linkage having dual kinematical properties. This idea has opened up a new avenue of research and practical applications since due to this relation, knowledge and methods available for one of these systems can be transformed and employed in the other [2]", + " Thus, in accordance with the results of the section Transferring the method to linkages, the transformed Willis method in trusses enables to replace a compound truss with a simple one using the proposed technique. To clarify this idea, the compound truss given in Fig. 10 is analysed using the transformed Willis method. The truss of Fig. 10 is compound, thus it is reasonable to attempt solving it using the proposed procedure. The linkage dual to this truss is shown in Fig. 11a. The linkage of Fig. 11a, which actually is the same as the one treated in the section \u2018\u2018An example for the analysis of linkage by means of the transformed Willis method\u2019\u2019 (Fig. 8), is composed of a driving link 1\u2019 and a tetrad, thus according to the previous section it would be efficient to fix link 3\u2019 and change the driving link to 4\u2019, as shown in Fig. 11b. The transformed truss, dual to the transformed linkage of Fig. 11b, appears in Fig. 12a. The truss of Fig. 12 is a simple truss and thus it can be solved by one of the efficient methods available for solution of such trusses. One of such methods is the wellknown graphical method, called Maxwell-Cremona diagram [12], shown in Fig. 12b. One can see that consistently with the results reported in [2], the Maxwell Cremona diagram of the truss of Fig. 12a is identical to the image velocity diagram of its dual mechanism, as appears in Fig. 8. The solution of the transformed truss can now be substituted into Eq. 7 to yield the ratio between the weighted value of the external force and the weighted forces in the rods of the original truss. The algebraic manipulations needed to find the force in rod 6 of the original truss (Fig. 10) appear in Fig. 12c. The paper has introduced an approach for transforming methods between engineering fields through graph representations. It employs the fact that planetary gear trains and linkages have been represented by the same graph representation, to transform the Willis method from planetary gear trains to the terminology of linkages" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002781_wcica.2008.4592987-Figure4-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002781_wcica.2008.4592987-Figure4-1.png", + "caption": "Fig. 4 Wheel slip model on a rigid terrain", + "texts": [ + " The z direction of contact frame Ci is perpendicular to the tangent plane of the terrain at the contact point, while the x direction is parallel to the tangent plane of the terrain. The angle of the z direction of the contact frame Ci with respect to the z direction of the driving wheel frame Ai is i, which is defined as the wheel-terrain geometric contact angle. To consider wheel slip model, label the wheel-terrain contact point coordinate frame as Ci(t\u2022 \u2022t) at time (t\u2022 \u2022t), Ci(t) or Ci at time t. The wheel motion from Ci(t\u2022 \u2022t) to Ci(t) is defined by the wheel slip model involving rolling slip i plus wheel rolling displacement r i, lateral slip i and turning slip i. Figure 4 shows the geometric description of wheel slips. With the above definition of all coordinate frames, the transformation matrixes between each coordinate frame can be easily deduced. At any time t, the mobile robot has an instantaneous coordinate frame R attached to its body that moves with the robot, and its configuration vector u, is defined as (x y z x y z) T, with respect to a fixed worldcoordinate frame W, where (x y z)T is the position, and ( x y z) T is the orientation, with roll x, pitch y, and heading z" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003226_s1068366610010010-Figure8-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003226_s1068366610010010-Figure8-1.png", + "caption": "Fig. 8. Pressure distribution in cross section y = 0 at F = 8.75 \u00d7 10\u20136; S = 4.6 \u00d7 10\u20135 and different values of parameter Sr: 1\u2014Sr = 1.16 \u00d7 10\u20136; 2\u20144 \u00d7 10\u20137; 3\u20142.43 \u00d7 10\u20137; 4\u20141.67 \u00d7 10\u20137; 5\u20141.17 \u00d7 10\u20137.", + "texts": [ + " When the values S are small, the coefficient f declines in response to increased S, reaching a minimum, and then increases afterwards. The results in Fig. 7 show that the graphs h(x) slope more to the axis x as S grows; the dependences h(x) then resemble more the similar dependences in the UHD contact. As Sr declines (when the relaxation time ts is increased), the pressure distribution function becomes more symmetric in respect of the axes x; i.e., it approaches to the function of the pressure distribu tion in the UHD contact. This follows from Fig. 8, showing the pressure distribution functions in the middle cross section at F = 8.75 \u00d7 10\u20136; S = 4.6 \u00d7 10\u20135 and different values of the parameter Sr. These changes in the function p(x, y) result in con siderable growth of the coefficient fr in response to growth of Sr. This is confirmed by the dependences of the friction coefficient and its components in Fig. 9 on the parameter Sr at F = 8.75 \u00d7 10\u20136; S = 4.6 \u00d7 10\u20135. The coefficient ft grows weakly in response to increased Sr. This growth is due to the changes in the dependences \u03c1h on the coordinate x and y as Sr changes" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003562_iecon.2010.5675259-Figure6-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003562_iecon.2010.5675259-Figure6-1.png", + "caption": "Fig. 6. Experimental system", + "texts": [ + " The force controller of is defined as \ud835\udc36\ud835\udc53 = \u2223\ud835\udc39 \ud835\udc52\ud835\udc65\ud835\udc61 \ud835\udc63\ud835\udc52\ud835\udc5f \u2223. (12) This system estimates the force of the movement direction \ud835\udc39 \ud835\udc52\ud835\udc65\ud835\udc61 \ud835\udc5a\ud835\udc5c\ud835\udc63. The control system change the speed of the human operator depending on the force of the movement direction. Thus, force limiter \ud835\udc39\ud835\udc59\ud835\udc56\ud835\udc5a is defined. The speed scaling \ud835\udefc is calculated as \ud835\udefc = \u23a7\u23a8 \u23a9 \ud835\udc39\ud835\udc59\ud835\udc56\ud835\udc5a \ud835\udc39 \ud835\udc52\ud835\udc65\ud835\udc61 \ud835\udc5a\ud835\udc5c\ud835\udc63 (\ud835\udc39 \ud835\udc52\ud835\udc65\ud835\udc61 \ud835\udc5a\ud835\udc5c\ud835\udc63 > \ud835\udc39\ud835\udc59\ud835\udc56\ud835\udc5a) 1 (\ud835\udc39\ud835\udc59\ud835\udc56\ud835\udc5a) > \ud835\udc39 \ud835\udc52\ud835\udc65\ud835\udc61 \ud835\udc5a\ud835\udc5c\ud835\udc63 > \u2212\ud835\udc39\ud835\udc59\ud835\udc56\ud835\udc5a) \u2212 \ud835\udc39\ud835\udc59\ud835\udc56\ud835\udc5a \ud835\udc39 \ud835\udc52\ud835\udc65\ud835\udc61 \ud835\udc5a\ud835\udc5c\ud835\udc63 (\ud835\udc39 \ud835\udc52\ud835\udc65\ud835\udc61 \ud835\udc5a\ud835\udc5c\ud835\udc63 < \u2212\ud835\udc39\ud835\udc59\ud835\udc56\ud835\udc5a) (13) In this paper, loading sampling time \ud835\udc47 \ud835\udc59\ud835\udc5c\ud835\udc4e\ud835\udc51 \ud835\udc60 is scaring by \ud835\udefc \ud835\udc47 \ud835\udc59\ud835\udc5c\ud835\udc4e\ud835\udc51 \ud835\udc60 = \ud835\udc47\ud835\udc60 \ud835\udefc (14) where \ud835\udc47\ud835\udc60 is saved sampling time. Fig. 6 shows the experimental system. The experimental system is constructed by 2-DOF linear motor table. The experimental parameter is shown in Table I. The experimental procedure are described as follows, \u2219 Storing of the motion data: the human motion data is stored by the motion-saving system. In this experiment, the operator operate in a circular motion using experiment system. \u2219 Reproducing of the motion data: the motion loading system reproduce the stored motion data. \u2219 Variable time-space compliance: the motion decomposed to motion of movement direction and vertical direction" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003243_1.3195039-Figure3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003243_1.3195039-Figure3-1.png", + "caption": "Fig. 3 Difference mesh of the calculation sliders", + "texts": [ + " Then, the floating force may be calculated by the following equation: Fgas = pdxdy 10 In numerical analysis, the following convergence criterion is defined: Err n = Fgas n \u2212 Fgas n/2 Fgas n 11 In order to testify the validity of PNPFD method, numerical comparisons with the FPFD method are carried out for the two different kinds of experimental sliders two guides and multiguides , as illustrated in Fig. 2. Here, slider 1 is the two guides slider from the experiment of Tagawa 8 , and slider 2 is the multiguides slider from the experiment of Huang 9 . Also, calculation parameters are shown in Table 1. The flying attitude values in Table 1 are gained from the experiments of Tagawa and Huang. Figure 3 shows the difference grids of the two testing sliders. For FPFD, the difference directions are in the x- and y-coordinate directions. It is seen that the recess edges of slider 1 are all parallel to x or y, and that its recess edges are parallel to FPFD difference directions. But in slider 2, some recess directions are not parallel to the FPFD difference directions in the xOy coordinate system. 4 Results and Discussion 4.1 Pressure Distribution and Floating Force. Pressure profiles are given in Fig", + " Here, the pressure distributions of the two different types of sliders with two different finite difference methods FPFD and PNPFD. It is shown that the difference in pressure distribution induced by FPFD and PNPFD for slider 1 can be ignored, but for slider 2, this difference is apparent. It can be seen that there is obvious pressure saw-tooth at the trailing pad of slider 2, as shown in Fig. 4 c , gained from FPFD. The reason is that the difference direction follows the direction of length and width generally in the FPFD method. But, at the trailing pad of slider 2, as shown in Fig. 3 b , the direction of the texture is neither parallel to that of length nor to that of width. The PNPFD introduced here can solve this problem, and the pressure distribution is smooth at the trailing pad regime, as shown in Fig. 4 d . Effect of mesh grid density on the gas floating force is shown in Transactions of the ASME of Use: http://www.asme.org/about-asme/terms-of-use F w m c t t i n s t f J Downloaded Fr ig. 5. It is shown that the gas floating force increases apparently ith an increase in the mesh grid density" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000297_icsmc.2001.973518-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000297_icsmc.2001.973518-Figure2-1.png", + "caption": "Figure 2: Coordinate frames of a biped walking robot.", + "texts": [ + " 163 1 One complete walking cycle is divided into two phases: single support phase and double support phase. During the double support, one foot is on the ground and the other foot is in the swing motion. As soon as the swing foot reaches the ground, the biped robot is in the double support phase. The ZMP should be changed smoothly according to two support phases for dynamic stability. In this study, the ZMP pattern is set arbitrarily and smoothly within the support polygon before the biped robot begins walking. 3. Stabilization Control 3.1 Moment Equations A biped model is shown in Figure 2. A world coordinate frame 3 is fixed on the floor and a moving coordinate frame f is attached on the center of the waist to consider the relative motion of each particle. In modeling the biped robot, five assumptions are defined as follows: (1) The biped robot consists of a set of particles. (2) The foothold of the biped robot is rigid and not moved by any force and moment. (3) The contact region between the foot and the floor surface is a set of contact points. (4) The coefficients of friction for rotation around the X, Y and Z-axes are nearly zero at the contact point between the feet and the floor surface", + " The experimental schemes are as follows: (1) the unit patterns of the lower-limbs, waist and head are planned according to the step direction, (2) the compensatory unit patterns of the trunk and waist are derived by the stabilization control method that cancels the moments generated by the motion of the lowerlimbs, (3) the continuous walking pattern is combined on the basis of the step direction and is kept in the computer memory of WABIAN-RII. In walking experiments, we consider three continuous walking patterns as shown in Table 1, Table 2 and Figure 2. One walking pattern has different step length and time while another has the same step length and time. The experiment of S-type walking is conducted with the same walking time and length. sults clarify t,hat the stability of the continuous walking patterns is realized by the compensatory motion of the trunk and waist. Figure 5 shows a part of scenes of the S-type walking. The S-type walking with 10 steps is realized with the step time of 1.28[sec/step], the step width of 0.15[m/step] , and the turning angle 15[deg/step]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002018_004051756903900209-Figure3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002018_004051756903900209-Figure3-1.png", + "caption": "Fig. 3. Meridional tire geometry at 24-lb/in! inflation pressure, showing transducer locations.", + "texts": [ + " Through each at WESTERN OREGON UNIVERSITY on June 3, 2015trj.sagepub.comDownloaded from Tire Parameters The data reported in this study were obtained from three, identically constructed 8.25-14 two-ply rayon tires. Each ply had a cord count of 19 ends/in. be fore expansion. In the finished tire, the angle between the cord at the crown and the circu1l1feren_ tial line at tread center was 36\u00b0. The location of the transducers and the meridional geometry of the tires at 24 lb/in.: inflation pressure is shown in Figure 3. Tire I had four transducers located 90 0 apart circumferentially in the first or inner ply at the crown (R = 13.2 in.) ; Tire II had four trans ducers located 900 apart in the second or ollter ply at the sidewall (R = 11.8 in.) ; Tire III had a trans- hole, a separate cord is threaded in the manner that the eye of a needle is threaded, and the cord is doubled hack on itself. Thus, in service, load is transferred to the transducer by two segments of one cord at each end of the bar. This method of load transfer, with two cords in series with the aluminum bar, means that 2 lb force on the transducer cor responds to 1 Ib tension on the cord" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003354_optim.2010.5510527-Figure9-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003354_optim.2010.5510527-Figure9-1.png", + "caption": "Fig. 9. zoom of stator currents at t2 = 320ms", + "texts": [], + "surrounding_texts": [ + "Because a PMSM with an appropriate motor inverter was not at hand to verify the derived PMSM control, a simulation was utilized. In this section the correct emulation of an Active Front End \u2013 identical with a motor inverter and with similar control scheme \u2013 is verified by a comparison of simulation result and measurement. Figure 4 shows the measured transient response of an Active Front End. The three-phase side of the converter is connected to the 50- Hz lab grid by an inductance L\u03c3 = 250\u00b5H. Thereby the lineto-line voltage of the grid has a peak voltage of 540V reduced from the nominal voltage of u\u030212N = \u221a 2 \u00b7 400V = 566V by the winding ratio wp ws = 1.048 of the used transformer. On the converter DC side a resistor was connected in parallel to the DC-link causing a load current of 150A, yielding an active power of 90 kW. The transient time of three periods is due to the robust control parameters chosen for commissioning. An advantage of these parameters is the well-damped transient response with marginal overshoot of the DC-link voltage. In the same manner the AC currents offer a well-damped behavior leading to steady-state amplitude within two half periods. The optimal 30\u25e6-phase shift between line-to-line voltage and negative line current can be clearly identified. Thus, the proper operation of the chosen control is proved by the transient response upon a 90-kW-load step. In order to verify the used simulation approach, the real parameters of the test bench have been used in combination with the set of first-order differential equations. The same control algorithm and the same load step are used for the verification (Fig. 5). Besides the mentioned peak-to-peak voltage of u\u0302pg = 566V and the coupling inductance of L\u03c3pg = 250\u00b5H, a parasitic resistance of Rpg = 1m\u2126 is assumed and the winding ratio of wp ws = 1.048 is neglected. The parameters of the DC-link components are as follows: a DC-link capacitor of CDC = 4.2mF and its parallel parasitic resistance of RDC = 12 k\u2126. It can be clearly seen that simulation and measurement results fit very good. The voltage drops of the DC-link capacitor equal and the resulting transient-response times forced by the used controller have the same duration. In the same manner the only differences between amplitude, phase angle and dynamic shape of the simulated and the measured currents are noise and peaks caused by the measurement hardware. Based on this perfect match verifying the high accuracy of the chosen simulation approach, the forthcoming analysis of the control performance is done using the simulation." + ] + }, + { + "image_filename": "designv11_69_0000370_iros.1991.174444-Figure6-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000370_iros.1991.174444-Figure6-1.png", + "caption": "Fig. 6. Projection of trajectories and self-motion manifolds for Task 1 onto the 82-83 plane", + "texts": [], + "surrounding_texts": [ + "x-axis position of the center of the circle. The task is to rotate the circle of r unit radius, centered at (e, , 0). in unit time, in counterclockwise, thus the initial position at to = 0 is ( c , - r , 0). And the farthest position from the base of the manipulator and the initial position is at t = 112, therefore, we denote that as mid-position of a cyclic task, which is (c,+r, 0).\nConsider a task denoted as Task 1 where r = 1 and c, = 3 units. As shown in Fig. 1, Task 1 is placed between W-sheer 2 and W-sheet 4(a). There is one self-motion manifold comsponding to the initial position, i.e., (2. 0) in W-sheet 2, and one to the mid-position of the task, i.e., (4, 0) in W-sheer 4(a). Fig. 4 shows the MCL and self-motion manifolds for Task I . In this example, there exist twelve optimal configurations as shown in Fig. 4. But, the manipulator posture for the three-link manipulator is symmetric with respect to the origin of the O r e 3 plane, so that we may only consider the case of e3 2 0 that is an upper-arm posture with respect to link 2. The optimal configuration comsponds to extrema1 point on the curve of H , i.e., N'h = 0. So, the optimal configurations are referred to Point,, k = 1. . . 6. as denoted in Fig. 4. As mentioned before, the MCL consists of LMINL and LMAXL. Therefore, one can easily expect that the conservative joint motions may be produced from the initial configurations in Point,, k = 1.3.5, which are local maxima. We also note that Point,, k = 2.4,6, are local minima.\nRadius 1\nTo show the discussed limitations, we carry out the Task 1 using the exrended Jucobiun method in which each points are used as initial configurations. In Fig. 5 , the joint trajectories from Point, and Point4 face with singular configurations during performing the task. These initial configurations are local minima and reside in LMINL as shown in Fig. 4. And also the joint trajectory form Point6 can not completely perform the task because Point6 is placed at Group IV and a local minimum. Therefore, the invertible workspace of Group IV is limited by a circle of radius 3.5 as shown in Fig. 3. These trajectories obtained from the local minima amve at the singular configurations or can not carry out the task because of the limitation of the invertible workspace.\nOn the other hand, joint trajectories in Figs. S(a) and (c) carry out Task 1 without singularities while maximizing the manipulubil-\nGroup 111 Group IV 6.090 (2.18. 1.52) X X Group I\n4\n6\n2 I 8.663 (1.95, 1.35) 14.536 (1.97, 2.72) 17.812 (-3.139, 1.43)(\n14.155 (1.49, 1.02) 8.368 (-1.37. 2.70) X\n15.360 (0.86, 0.75) X X\n3 I 11.597 (1.73, 1.18)16.185 (-1.73, 3.12)l 10.030 (-2.75, 1.1O)l\nare the global maxima regardless of positions. Also, the invertible workspace of LMAXL of Group I reachs from the base of the manipulator to the boundary of the workspace. Therefore, we can obtain the inverse kinematic solution in two stages. One is to select a proper initial configuration on the self-motion manifolds\n- /?O -", + "for the initial position to achieve a globally optimized configuration [U]. The other ,is to solve inverse kinematic algorithms, such as (1) and (41, or extended Jacobian method, with a global maximum as the initial configuration. Therefore, it will trace the global maxima of MCL and results in conservative joint trajectories without singularities for almost entire workspace.\nTo inspect the procedure, consider a cyclic task, denoted as Task 2, which is r = 3.5 and c, = 3.7. Hence, the initial position is (0.2. 0) and the mid-position is (7.2, 0) at I = ID. It is an extremely large task in comparison with the geomctry of the manipulator and this task occupies almost entire workspace between Wsheets 1 and 4@). In Fig. 7, two small dotted circles and one inner", + "dotted circle are self-motion manifolds corresponding to the initial position and mid-position of Task 2. respectively.\nAs shown in Fig. 7, the joint trajectory for Task 2 retains arm posture and resides in only first quadrant in the plane. This joint trajectory is optimal and maximizing the manipulability measure without excessive joint motions and singularities. I t can be thought of as a generalization of concept of postures in the nonredundant manipulator into the redundant manipulator. The locus in the first quadrant corresponds to down-elbow posture and the other in the third quadrant is up-elbow posture. These results are same as the down-elbow and up-elbow postures of a nonredundant two-link manipulator. Therefore, by observing the global maxima of the MCL and the invertible workspace, we can determine a promising singular-free joint trajectories for a task occupying almost entire workspace.\n4. Conclusions\nWe suggested the manipulability constraint locus which is the set of configurations satisfying the necessary constraint for optimizing manipulability measure. An MCL based on the topological property of the optimality constraint characterizes the performance of a subtask when the manipulability measure is used. Through cyclic tasks, we explained how an inverse kinematic algorithm solves a singular-free joint trajectory for the given cyclic task by using the MCL, and also why the inverse kinematic algorithm can blunder into the singular configurations during performing the task by using LMINL.\nGenerally, we discussed the property of the inverse kinematic algorithm to perform a given cyclic task as well as satisfying the optimality constraint, and also showed that it is analogous to the poses for a nonredundant manipulator. And we investigated the global property of the MCL to provide the joint trajectory with considering the invertible workspace. Also, using the global maxima on self-motion manifolds for each of W-sheets. we suggest an\nalgorithm which generates a conservative joint trajectory for a task occupying almost entire workspace.\nTherefore, in order to optimize some objective functions in the invertible workspace, the MCL can be used to investigate some fundamental relations between the properties of the inverse kinematic algorithms and its ability to resolve such problems.\nReferences\nD. E. Whiney, \"Resolved motion rate control of manipulator and human prostheses,\" IEEE Trans. on Man-Machine Systems, MMS-10-2. pp. 47-53. 1969. C. A. Klein and C. H. Huang, \"Review of pseudoinverse control for use with kinematically redundant manipulators.\" IEEE Trans. on Sys., Man, and Cybernetics, SMC-13-3, pp. 245- 250. 1983. J. Baillieul. \"Kinematic programming alternatives for redundant manipulators,\" Proc. IEEE Int. Conf. on Robotics and Automation. pp. 722-728, 1985. D. R. Baker and C. W. Wampler 11, \"On the inverse kinematics of redundant manipulators,\" Int. J. Robotics Research,\nP. H. Chang, \"A closed-form solution for the control of manipulators with kinematically redundancy,\" Proc. IEEE Int. Con5 on Robotics and Automation, pp. 9-14, 1986. A. A. Maciejewski, \"Kinetic limitations of the use of redundancy in robotic manipulators,\" Proc. IEEE Int. Conf. on Robotics and Automation, pp. 1 13- 11 8. H89. J. Baillieul, \"Avoiding obstacles and resolving kinematic redundancy,\" Proc. IEEE Int. Conf. on Robotics and Automation, pp. 1698-1704, 1986. D. N. Nenchev, \"Redundancy resolution through local optimization: A review,\" Int. J. Robotic Systems, 6(6). pp. 769-798, 1989. J. W. Burdick. \"On the inverse kinematics of rcdundant manipulators: Characterimtion of the self-motion manifolds,\" Proc. IEEE Int. Conf on Robotics and Automation, pp. 264-270, 1989.\n7(2), pp. 2-21, 1988.\n[IO] A. Ghosal and B. Roth, \"Instantaneous properties of multidegrees-of-freedom motions - point trajectories,\" Trans. ASME J. Mech., Trans., and Automat., in Design, 109, pp.\n[ I l l C. W. Wampler 11, \"Winding number analysis of invertible workspaces for redundant manipulators,\" Inr. J. Robotics Research, 7(5), pp. 22-31. 1987.\n[I21 I. D. Walker and S . I. Marcus, \"Subtask performance by redundancy for redundant manipulators,\" IEEE J. Robotics and Automation, RA-4-3, pp. 350-354. 1988.\n[I31 T. Yoshikawa, \"Analysis and control of robot manipulators with redundancy.\" Robotics Research-1st Int. Symposium, eds. by M. Brady and R. Paul: pp. 735-731, 1984.\n[I41 A. LiCgeois. \"Automatic supcrvisory control of the configuration and behavior of multibody mechanisms.\" IEEE Trans. on Sys.. Man, and Cybernetics, SMC-7-12, pp. 868- 871, 1977.\n[I51 J. W. Burdick, B. C. Cetin. and J. Barhen, \"Efficient global redundant configuration resolution via sub-energy tunneling and terminal repelling,\" Proc. Int. Conf. on Robotics and Automation, pp. 939-944, 1991.\n107-1 15, 1987.\n- / 7 2 -" + ] + }, + { + "image_filename": "designv11_69_0003781_isciii.2009.5342279-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003781_isciii.2009.5342279-Figure2-1.png", + "caption": "Fig. 2. The outline of the proposed improvement of the driving system", + "texts": [ + " Utilizing the well known fact that the acceleration of the mass center point of a rigid body multiplied by its full mass is equal to the sum of the external forces acting on that system, and that the timederivative of momentum of the system computed with respect to the actual mass center point is equal to the momentum of the external forces (torque) with respect to this point the required active driving force components FAeA , FAfA , FBeB , FBfB , and FCeC , FCfC , as well as the hypothetical vertical constraint force components FAz , FBz , and FCz can be calculated. According to Fig. 1 if the small wheels do not have drives in the horizontal e directions no forces can be exerted. If the small wheels obtain drives (a rough idea for that is outlined in Fig. 2) then we can calculate with the existence of such components, too. It has to be noted that the laws of Classical Mechanics generally do not unambiguously determine these forces. It is physically possible to make the wheels working against each other: any pair of forces consisting of identical absolute value and opposite direction acting along the same line yield zero net torque for any point, and zero net force. To resolve this ambiguity in our case a Moore-Pennrose pseudo-inverse was applied for solving the following optimal task in the case of driven small wheels: the sum of the squares of the above components must be minimal under the constraints that the appropriate linear and rotational acceleration has to be achieved" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001712_nme.1428-Figure4-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001712_nme.1428-Figure4-1.png", + "caption": "Figure 4. Contact of rigid flat punch with rounded corners and the punch geometry.", + "texts": [ + " The calculated nodal values of the function g( i ) given in Table I compare well with the analytical result in Equation (62). Numerical results in Table I provide a clear demonstration that the proposed method is an efficient technique for the solution of singular integral equations of the second kind frequently encountered in contact mechanics. Copyright 2005 John Wiley & Sons, Ltd. Int. J. Numer. Meth. Engng 2005; 64:1236\u20131255 The problem of a fully sliding contact between the flat semi-infinite elastic substrate and a rigid flat punch with rounded corners, shown in Figure 4, is described by the equation \u2212f 1 \u2212 2 2(1 \u2212 ) (x) + 1 \u222b a1 \u2212a1 (t) t \u2212 x dt = \u2212 (1 \u2212 ) f (x) (63) where the derivative of gap function f (x) of the punch can be expressed as f (x) = \u2212x + r + e R , x \u2208 [\u2212a1, \u2212(r + e)] 0, x \u2208 [\u2212(r + e), (r \u2212 e)] \u2212x \u2212 (r \u2212 e) R , x \u2208 [(r \u2212 e), a1] (64) where r is the half-width of the flat end. The eccentricity e requires numerical determination from Equation (26). We carried out calculations for = 0.33, E = 2 (1 + ) = 1.15 \u00d7 1011 Pa, R = 50 mm, r = 4.15 mm, P = 1" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001120_bfb0035241-Figure8-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001120_bfb0035241-Figure8-1.png", + "caption": "Figure 8. Experimental Setup for Puma 560 Disturbance Sensitivity Tests", + "texts": [ + " They are impor tant to ensure that the actuators do not saturate over the desired range of inputs and expected disturbances. Experimentally, this controller performs well. A sample s tep respon~;e is shown in Figure 7. Response t ime is reasonable, and s teady-s ta te errors are smM1. 4.1. D i s t u r b a n c e R e j e c t i o n The results in Figure 7 show tha t our controller has good low-frequency rejection of process disturbances. We will now examine the effect of output disturbances on our PID torque controller. By positioning the manipulator properly, as shown in Figure 8, we can isolate the effect of environmental contact forces during unified position/force control to a torque on joint three. To introduce output disturbances in the torque control loop, we command the arm to maintain a horizontal force of l tN on the rigid aluminum post while sliding up at 0.05 meters per second. Motion will be controlled with a simple, operational-space PD controller. Figure 9 shows the PID controller's performance during compliant motion. The top plot represents the position error in the vertical direction, while the bottom figure shows the desired and sensed torques for joint three" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003799_kem.417-418.313-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003799_kem.417-418.313-Figure2-1.png", + "caption": "Fig. 2 Schematic of a continuous rail beam model on continuous elastic foundations", + "texts": [], + "surrounding_texts": [ + "Calculation of global bending stress is based on the accepted railroad engineering practice of treating the rail as a continuous beam on a continuous elastic foundation [2,3]. Fig. 1 shows the loads and foundations involved in the analysis. The rail is loaded by vertical load from vehicle weight and lateral from centrifugal forces and the rail is supported by vertical, kv, lateral, kL and torsional, k\u03c6, foundation. The longitudinal bending stress, \u03c3xxB can be calculated by the summation of vertical and lateral stresses. The Mv and ML are the bending and lateral moment, respectively. The \u03be is the wheel points and (x-\u03be) is the distance between wheel points and stress calculation points. LVxxB \u03c3\u03c3\u03c3 += (1) where zz L L yy V V I exM I fxM \u22c5 = \u22c5 = )( , )( \u03c3\u03c3 [ ]\u2211 \u2212\u2212\u2212\u2212= \u2212\u2212 i iviv x v i v xxe V xM iv \u03be\u03b2\u03be\u03b2 \u03b2 \u03be\u03b2 sin)(cos 4 )( [ ]\u2211 \u2212\u2212\u2212\u2212= \u2212\u2212 i iLiL x L i L xxe L xM iL \u03be\u03b2\u03be\u03b2 \u03b2 \u03be\u03b2 sin)(cos 4 )(" + ] + }, + { + "image_filename": "designv11_69_0002609_icma.2008.4798870-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002609_icma.2008.4798870-Figure2-1.png", + "caption": "Fig. 2. End positions lie within the same image (1. Case).", + "texts": [ + " To close the open chain between the points U and V within the same image, the vector connecting these points as well as the orientation of the internal frame of the calibration object w.r.t. the camera frame need to be known. This information can be retrieved 1The position of points on the calibration objects within the base frame, either calculated or measured, is from here onwards also referred to as \u201cend position\u201d. by application of the Gram-Schmidt algorithm, unless the recorded image does not contain at least three points that do not lie on the same line. The situation is shown in Fig. 2: The open chain between the two points on the calibration object w.r.t. the same base frame can be closed either via point U or point V . Assuming the joint configuration qj connects further to point V , then we need to extend (3) to include the connecting vector. Since the position uDH,s,e connects further to point V , the denotation is changed to v\u0302DH,s,e con with the subscript con for connecting vector: v\u0302DH,s,e con (\u03bd ,q j) = 0T DH,s,e n (\u03bd ,q j)nTc cTo ov\u0302 (5) where cTo represents the transformation matrix of the camera frame Sc into the internal frame So of the calibration object with the origin U " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000560_robot.1988.12083-Figure3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000560_robot.1988.12083-Figure3-1.png", + "caption": "Figure 3: Schematic diagram of a three-d.0.f. Cartesian type pantograph.", + "texts": [ + " It was shown9 that the magnification ratio between points F and A, and between points F and B are respectively represented as: The angle 1c. can be calculated as: (2) In order to obtained decoupled motion at point F . the orientations of axes U and V are specified as in Figure 4. There are several different ways of specifying the orientations of axes U and V and were shown.lo Cartesian Type Since only an ordinary pantograph can be used in a Cartesian type pantograph,1\u00b0 the skew angle should be either 0' or 180\". From Equation (l), 0 = 180\" gives R ' = R + 1 . Referring to Figure 3, the forward position equations are: where Xf, Yf and Zf are the coordinates of the hand reference point F ; U, W and V are the linear displacements of the input points A and B along the three actuator axes. U and V are measured from the intersection of these two axes. W is measured from the intersection of U- and W-axes. As is apparent, all the d.0.f. are fully decoupled. Cylindrical Type Referring to Figure 1, a cylindrical coordinate system (P,Q,&) is chosen as the reference frame to describe the motion of the hand reference point F ", + " Hence, we may conclude that a closed-chain, parallelogram manipulator (or an open-chain manipulator with pulley and chain system) has better mechanical efficiency than an open-chain manipulator and a pantograph type manipulator has the best mechanical efficiency among all. KINEMATICS OF S I X - D . O . F . PANTOGRAPH MANIPULATORS A six-d.0.f. pantograph type manipulator is obtained by attaching a three-roll wrist at the end link of either a Cartesian or cylindrical type pantograph. The wrist center should be coincident with the point F in order to have simpler kinematic relationship. For a six-d.o.f., Cartesian type pantograph manipulator (referring to Figure 3 ) , the position, velocity and acceleration analyses of the first three axes (U, V and W) are very simple since these three degrees of freedom are fully decoupled. The analyses of the last three axes ( 6 4 , 0 5 and B e ) , however, are not that simple due to the following reasons: The wrist experiences a pitching motion when either one or both of the first two axes (U and V) move. That is, the angle a2 changes. Also, the wrist experiences both a pitching and yawing when the third axis W moves. This is, both a2 and 4 change" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001626_s11018-006-0152-2-Figure4-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001626_s11018-006-0152-2-Figure4-1.png", + "caption": "Fig. 4. Set-up of vibrator and points of measurement of the absolute oscillations in a screw-cutting lathe for the purpose of estimating the balance of elastic displacements: 1\u201327) measurement points; I\u2013VII) centers of gravity.", + "texts": [ + ", a body which moves in space, however, any natural deformations in the body will be negligible; \u2022 a flexible beam (long frames, posts); \u2022 an immovable and nondeformable body. An example of an elastic form in the XOY plane is shown in Fig. 3. Let us illustrate the quasi-static method of determining the balance of elastic displacements, using as an example a screw-cutting lathe. The set-up of a piezoelectric vibrator and the points at which the absolute oscilations are measured for the purpose of determining the balance of elastic displacements, i.e., deformations of the elastic system as a function of the force applied to the cutting point, are shown in Fig. 4. A VshV-003 vibrator with DN-5 sensor was used to measure the oscillations. The amplitude-frequency characteris- tic of the elastic system of the lathe was first determined. The quasi-static behavior of the elastic system for the given lathe was maintained up to frequencies on the order of 80 Hz, hence excitation frequencies not greater than these values must be adopted for the measurements. From the point of view of increasing the measurement precision, it is best to work at higher frequencies, since interference from oscillations of the foundation is thereby reduced and the operation of the apparatus improved" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003384_10402000802369739-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003384_10402000802369739-Figure1-1.png", + "caption": "Fig. 1\u2014Schematic of test head.", + "texts": [ + " The maximum likelihood estimate of the shape parameter based on an uncensored sample of size n is the solution of the nonlinear equation: ( n\u2211 i=1 t \u03b2\u0302 i )( 1 \u03b2\u0302 + 1 n n\u2211 i=1 ln ti ) \u2212 n\u2211 i=1 t\u03b2\u0302i ln ti = 0 [2] Having solved this equation for \u03b2\u0302, the maximum likelihood estimator of the shape parameter, the ML estimate of the scale parameter is computed from: \u03b7\u0302 = [ 1 n n\u2211 i=1 ti \u03b2\u0302 ] 1 \u03b2\u0302 [3] The ML estimate of a quantile such as t0.10 (also referred to as L10 or B10) may be computed from the relation: t\u0302p = [ ln [ 1 1 \u2212 p ]] 1 \u03b2\u0302 \u00b7 \u03b7\u0302 [4] By definition, the probability of failing prior to tp is p. 223 T ri bo lo gy T ra ns ac tio ns 2 00 9. 52 :2 23 -2 30 . The test equipment used in this project consisted of two dualheaded machines of the ball-on-rod type. The basic description is given in Glover (3). As shown in the sketch in Fig. 1, each head consists of a rotating cylindrical test rod in contact with three balls under lubricated, pure rolling conditions. The three balls are radially loaded against the M-50 steel test rods by means of a load applied to two tapered bearing cups, which squeeze the balls into contact with the test rod. The contact is lubricated by a drip-fed aircraft turbine engine lubricant. The test rod diameter is 9.5 mm (3/8 in.) and its length is 76.2 mm (3 in.). The ball diameter is 12.7 mm (0.5 in.). The balls are made of SAE 52100 steel" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002870_s0001924000004802-Figure7-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002870_s0001924000004802-Figure7-1.png", + "caption": "Figure 7. Definition of indicator surface in phase plane.", + "texts": [ + " (10) ( , ) ( )[ ( ) ( ) ] ( ) for all 0 me f t g t u t d t e t t = + +\u23a7 \u23a8 < \u03b4 \u2265\u23a9 e&& . . . (12) SM uutu *I)( += . . . (13) } Therefore According to Equation (20), the last inequality implies V . < 0. Therefore the supervisory controller(23) stabilises the system in the sense of Lyapunov and guarantees that \u23d0e\u23d0 will decrease if \u23d0e\u23d0 \u2265 \u03b4. Since it is required to absolutely prevent the situation that \u23d0e\u23d0 \u2265 \u03b4, in the next step we define an indicator surface, I; where \u03bbS is specified by the designer. The so defined indicator surface has been shown in Fig. 7. It can easily be observed that to satisfy the condition \u23d0e\u23d0 < \u03b4, it is also necessary to control I such that Therefore the definition of indicator function can be modified as follows: Since I * in Equation (26) is a step function, the supervisory controller begins operation as soon as e hits the boundary \u23d0e\u23d0= \u03b4 or I hits the boundary \u23d0I\u23d0= \u03bbS\u03b4, and is idle while e and I satisfy the constraint sets \u23d0e\u23d0 < \u03b4 and \u23d0I\u23d0 < \u03bbS\u03b4, hence the system may oscillate across the boundary \u23d0I\u23d0 = \u03bbS\u03b4. One way to overcome this chattering problem is to allow I* to continuously changing from 0 to 1" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003210_j.mechmachtheory.2009.09.007-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003210_j.mechmachtheory.2009.09.007-Figure1-1.png", + "caption": "Fig. 1. The considered planar system.", + "texts": [ + " Usual recalls about theoretical and technical details will be given in Appendices. Let \u00f0O;~i;~j\u00de be a reference frame (which is a direct orthonormal basis), p an integer greater than or equal to 2, \u00f0li\u00de16i6p; p non negative numbers and \u00f0h\u00fei \u00de16i6p and \u00f0h i \u00de16i6p2p angles satisfying 8i 2 f1; . . . ;pg; p < h i < h\u00fei 6 p: \u00f01\u00de For all figures, hi\u2019s are chosen counter-clockwise. They are algebric angles; thus, representation could be either counterclockwise or clockwise. We define the workspace as the set of points Ap such as (see Fig. 1) A0 \u00bc 0; \u00f02a\u00ded ~j;0A1 ! \u00bc h1; \u00f02b\u00de 8i 2 f2; . . . ;pg; d Ai 2Ai 1 ! ;Ai 1Ai ! \u00bc hi; \u00f02c\u00de 8i 2 f1; . . . ;pg; Ai 1Ai \u00bc li; \u00f02d\u00de with the constraints 8i 2 f1; . . . ;pg; hi 2 h i ; h \u00fe i : \u00f02e\u00de We consider function Up from domain F \u00bc Yp i\u00bc1 h i ; h \u00fe i ; \u00f03\u00de to R2 and defined by 8\u00f0h1; . . . ; hp\u00de 2 F; Up\u00f0h1; . . . ; hp\u00de \u00bc Ap: \u00f04\u00de All the elements x \u00bc \u00f0h1; . . . ; hp\u00de of domain F satisfy (2e). Thus, according to constrained optimisation technics, consider the following definition: Definition 2.1. For all x \u00bc \u00f0h1; ", + " For simulation, a subject of 1.80 m of height is considered. Lengths of the upper limb were determined from anthropometric data [15]. Thus, segment lengths are presented as percent of total body height (0.108, 0.146 and 0.186 for the hand, forearm and upperarm, respectively, for the right arm). Angles correspond to minima and maxima of joints degrees of freedom of human upper limb, i.e. shoulder abduction/adduction ( 60 /120 ), elbow flexion/extension (0 /130 ), wrist abduction/adduction ( 10 /25 ). See Fig. 1, where p \u00bc 3 and O is the shoulder, A1 is the elbow and A2 is the wrist, i.e. segment 0A1 is the upperarm, A1A2 is the forearm and A2A3 is the hand. (1) Case 1 corresponds to free upperarm, fixed forearm and fixed hand displacements; (2) Case 2 corresponds to free upperarm, free forearm and fixed hand displacements; (3) Case 3 corresponds to free upperarm, free forearm and free hand displacements. Remark 5.1. Results of Appendix A only hold for p P n \u00bc 2. However, for theses simulations, we will use the value p \u00bc 1" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003932_aim.2010.5695862-Figure3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003932_aim.2010.5695862-Figure3-1.png", + "caption": "Fig. 3. Angles to the faulty insulator.", + "texts": [ + "b) The 1d in (6) represents the distance to the first microphone from the faulty insulator radiating the sound signal. The same procedures are applied for (5.b) and (5.c), and the distances to the second and the third microphones are 2 2 2 2 1 12 13 12 13 2 1 2 2 2l d d vel t t d D \u22c5 \u2212 \u22c5 \u0394 \u2212 \u0394 + \u22c5 \u22c5 \u0394 \u22c5 \u0394 = , (8.a) 2 2 2 2 1 12 13 12 13 3 1 2 2 4l d d vel t t d D \u22c5 + \u22c5 \u0394 + \u0394 \u2212 \u22c5 \u22c5 \u0394 \u22c5 \u0394 = . (8.b) With the measured distances, the angles to the sound source from the microphones can be also obtained. Fig. 3. illustrates the cosine laws to obtain the angle form the microphone to the sound source using the triangle, 1 2SM M\u0394 . The angle to the first microphone from the sound source, 1M\u03b8 , can be obtained as, 1 2 2 2 1 1 1 2 1 1 cos 2M d l d d l \u03b8 \u2212 + \u2212 = , (9.a) The same procedures are applied for the second and third microphones and the angles are obtained as, 2 2 2 2 1 2 1 1 2 1 cos 2M d l d d l \u03b8 \u2212 + \u2212 = , (10.b) 3 2 2 2 1 3 2 2 3 2 cos 2M d l d d l \u03b8 \u2212 + \u2212 = . (10.c) By these proposed algorithms, sound signal of bad insulator in two dimensional space could be estimated and carried out by TMS320f2812 bus type module", + " Three microphones are places at the center of the inspection robot, which is shown in fig. 4. To minimize the effect of the noise which is came from the DC motor or undesired moving of the robot hardware. The inspection robot is moving along the neutral wire on the top of the electric pole while it is inspecting the insulators and power transmission wires below the neutral wire. Therefore the angles to the insulators become less than 90 degree since the insulators are on the electric pole as shown in Fig. 4. As shown in Fig. 3, the triangle, 1 2SM M\u0394 can be divided into two triangles, and the coordinates of the three microphones are represented as 1 (0,0)M = , 2 1( ,0)M l= , and 3 2 3( ,0)M l l= + , respectively, with respect to 1M . When the coordinates of the sound source is ( , )S a b= , Pythagorean theorem provides two equations from the two triangles as, 2 2 2 1a b d+ = , (11.a) 2 2 2 1 2( )l a b d\u2212 + = . (11.b) Using these two equations, the coordinates of the sound source, ( , )S a b= , can be obtained as 22 2 2 2 2 2 21 2 1 1 2 1 1 1 1 , 2 2 d d l d d lS d l l \u2212 + \u2212 + = \u2212 " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001382_j.tsf.2005.07.024-Figure3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001382_j.tsf.2005.07.024-Figure3-1.png", + "caption": "Fig. 3. Cyclic voltammograms for cast films (a) 1/4C8N +Br (molar ratio, 1:19), (b) 1/4C7N +Br (molar ratio, 1:19), and (c) 1 in the absence of a matrix on a BPG electrode in a 0.5 M aqueous KCl solution at 25 -C under CO. Scan rate: 0.1 V/s.", + "texts": [ + " The results indicate that the matrix surfactant thin film acts as suitable electrode modifier for examining the CO binding properties of the diruthenium complex without any inactivation. The electrochemistry of complex 1 under CO atmosphere was also investigated at a modified electrode containing a film which with alkylammonium matrices 4C8N +Br and 4C7N +Br as well as in a cast film of the complex 1 itself without a matrix to find the most suitable environment for examining CO binding ability of complex 1 at the film state in aqueous media (Fig. 3a\u2013c). The potentials (E1/2 or Epc and Epa) measured for the four redox process under these experimental conditions are summarized in Table 1 which also includes electrochemical data of 1 at the modified electrode in the absence of matrix film. The complex 1 at the modified BPG electrode with 4C8N +Br or 4C7N +Br under N2 atmosphere undergoes a reversible Ru2 6+/Ru2 5+ oxidation, a reversible Ru2 5+/Ru2 4+ reduction, and an irreversible Ru2 4+/Ru2 3+ reduction [10], whose behavior almost identify with the film electrochemistry of 1 at the 4C8P +Br modified electrode. In the case of under CO atmosphere (Fig. 3a, b), E1/2 values for an oxidation of the complex (Ru2 6+/Ru2 5+) are the same under N2 or CO. Complex 1 undergoes three reductions corresponding to Ru2 5+/Ru2 4+, Ru2 4+/Ru2 3+, and Ru2 3+/Ru2 2+ redox processes which are similar to what is seen in the 1/ 4C8P +Br modified electrodes under CO. In contrast, a cast film of complex 1 in the absence of a matrix under CO undergoes only an oxidation which is located at E1/2=0.58 V (Fig. 3c). The results suggest that any of the three investigated surfactant matrices (4C8P +Br , 4C8N +Br , and 4C7N +Br ) provides a more suitable microenvironment for examining the CO binding ability of diruthenium complexes (such as 1) than cast films of the same complex at an electrode in the absence of a matrix. Aggregation of the complexes in the film is inhibited to some extent by adding surfactant matrices during the preparation of the films. This is why clear film electrochemistry and CO binding features were obtained in the matrix films" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002083_detc2007-34911-Figure4-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002083_detc2007-34911-Figure4-1.png", + "caption": "Figure 4. Modeling of sidewall.", + "texts": [ + " ownloaded From: http://proceedings.asmedigitalcollection.asme.org/pdfaccess.ashx?url In this section, the modeling of sidewall stiffness and contact forces between tread and road are discussed. The sidewall is modeled using the circumferential and radial springs and dampers defined between the belt and rim, while friction elements are attached along the flexible belt elements in order to account for the dynamic interaction between the tire and roads with short wavelength irregularities. Modeling of Sidewall As shown in Fig. 4, the sidewall flexibility of a tire is modeled using springs and dampers defined between the flexible belt and the rigid rim. These forces are generated due to the relative motion between the belt and rim in the circumferential and radial directions. It is assumed that flexible belt i and rigid rim j shown in Fig. 4 are assumed to be connected at point P that are, respectively, defined as i mP and j mP . The subscript m refers to the m-th spring-damper force application point. The global position vector of point i mP defined in element e on flexible belt i is given as ie ie ie ie P mm =r S T p (15) On the other hand, the global position vector of point j mP defined in rigid rim j is given as j j j j P Pm m = +r R A u (16) where jR is global position vector at the origin of body coordinate system attached to the center of the rim, jA is the orientation matrix, and j Pm u is the local position vector at point j mP that is given by 0[cos sin ]j T P m mm R \u03c6 \u03c6=u (17) where the rigid rim is assumed to be a circle of radius 0R and the angle m\u03c6 ( 0 2m\u03c6 \u03c0\u2264 \u2264 ) defines the location of force application point on the edge of the rigid rim", + " 16 and 17, the relative distance and velocity between the points P are defined as ,ij ie j ij ie j m P P m P Pm m m m = \u2212 = \u2212d r r d r r (18) Using the preceding equations, the components in the radial and circumferential directions are defined as , T Tij j ij ij j ij m m m m m m= =d A d d A d (19) where the matrix j mA is given as j j j m m=A A A (20) where the matrix j mA defines an orientation of the sidewall 4 Copyright \u00a9 2007 by ASME =/data/conferences/idetc/cie2007/71022/ on 05/12/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use coordinate system with respect to the body coordinate system. This coordinate system is used to define the radial and circumferential directions of the tire with respect to the body coordinate system as j j j m m m\u23a1 \u23a4= \u23a3 \u23a6A n t (21) Accordingly, the sidewall forces defined with respect to the sidewall coordinate system can be defined as 0( )ij ij ij ij ij m m m= \u2212 \u2212 \u2212F K d \u03b4 C d (22) where 0 0[ 0]T\u03b4=\u03b4 is as given in Fig. 4, and ijK and ijC , respectively, define the stiffness and damping matrices as [ ], [ ]ij ij r t r tdiag k k diag c c= =K C (23) where the subscript r refers to stiffness and damping in the radial direction, while t refers to those in the circumferential direction. Using Eq. 20, the sidewall force ij mF given by Eq. 22 can be defined with respect to the global coordinate system by ij j ij m m m=F A F . Accordingly, the generalized forces associated with the sidewall flexibility can be defined for flexible belt body i and rigid rim body j as follows: , T T Ti ie ie ij j j ij m m m m m m= = \u2212Q T S F Q B F (24) where [ ( / ) ]j j j j m Pm \u03b8= \u2202 \u2202B I A u " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002133_ramech.2008.4681514-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002133_ramech.2008.4681514-Figure1-1.png", + "caption": "Figure. 1 :The considered mobile robotic system manipulating an object", + "texts": [ + " Next, the system non-holonomic constraint will be derived, and using Natural Orthogonal Complement (NOC) Method an independent set of equations of motion is derived for the system. Finally, the MIC law is applied to manipulate an object with two cooperating manipulators tracking a given path. The obtained results reveal a coordinated motion of the object, manipulators and the base vehicle. II. SYSTEM DYNAMICS A. Basic Definitions and Calculations. A wheeled robotic system is considered on a flat surface as depicted in Fig. 1. Direct path method (DPM), [7], is used to express the base and the links center of mass position and orientation, and linear and angular velocities of the base and each link. To derive equations of motion, using Lagrange approach, it can be written: 978-1-4244-1676-9/08 /$25.00 (\u00a92008 IEEE RAM 20081124 The mass matrix (H), non-linear velocity vector (C) and gravity vector (G), are obtained in the following form: aiRb aRb a6b a Ob H = M + ib. I Jaqi aqj a4i a4j d aT aT dt a4i aqi +a=QZaqi i=l. N (1) +LNmk ck (m) ak (m) + Mk~ m=e k=C aqi as: E(m) art)Rb E (m) rCf) aRb Vector C could be written as: where T is the total system kinetic energy, U is the total system potential energy, N is the system degrees-of-freedom (DOF), qi , qi , and Q, are the i-th element of the vector of generalized coordinates, generalized speeds, and generalized forces, respectively, as defined below: q ={Rb X ,ql , q2 I (2) where Rb and 0 are position vector and the yaw angle of the base, and q1 and q2 are the first and second manipulator vectors ofjoint angles respectively, as below: Rb =(xGYG )T q2 (=(2) 9(2))T C = C1 q + C2 where: C ij=M b )+j 'b X,+Jb aq,aj, & b J& +e*( k X--a qs ) S= qs )m--glMk )q (3a) (3b) (3c) The terms of the total system kinetic energy (T) are explicitly detailed in [4] for a general unconstrained mobile robotic system", + " (29), the required controlling force GFe req could be (27) written: GFe req MMdes (Mdes Xdes + kd + kPe+FC)-Fw-FO (33) t the The required desired force obtained in (33) could be used to ields determine 1f for both manipulators, [11]: and itegy obile y the vel is 02xl 1 -c =~Flf e req F 2 e req (34) It can be shown that by application of the MIC law, all participating manipulators, the moving base and the manipulated object behave with the same desired impedance behavior, [14]. VI. SIMULATION RESULTS AND DISCUSSIONS The simulated system consists of two 6-DOF manipulators mounted on a wheeled mobile platform as shown in Fig. 1, while the moving base is driven with two differentially driver wheels. All geometric and mass properties of the mobile base, and each of the two identical manipulators are given in Tables (1)-(2). The first manipulator is equipped with a remote center compliance (RCC) which is initially free of tension or compression, and its stiffness and damping properties are chosen as, [15]: ke = 2.4x10 kg S2and be 5.5x10 kg s and the object parameters are mQbJ = 3 kg , r (1) = -r (2) (-0. 45,0,0) where r(i) is the position i-th end effector with respect to the object center of mass" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002690_robio.2007.4522277-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002690_robio.2007.4522277-Figure2-1.png", + "caption": "Fig. 2. Manipulating a 2D object by a single DOF robot finger with soft tip in 2D plane", + "texts": [ + " In addition, if there arose rolling between the finger-tip sphere and the object without slipping, then it would give rise to the equality \u2212r d dt \u03c6 = d dt Y (2) that means the zero relative velocity of the contact point between one along the finger sphere and another on the object surface. Equation (2) implies \u03b4Y + r\u03b4\u03c6 = 0, which decreases the DOF of the system. Thus, the overall DOF of the system becomes zero, that is, motion would stop and the finger-end should be stacked at the contact position. In the case of a robot finger with a soft tip as shown in Fig.2, there arises an area contact between the finger-tip and the object, which can be expressed as \u0394x = r + l\u2212 {(xm \u2212 x0) cos \u03b8 \u2212 (ym \u2212 y0) sin \u03b8} (3) where \u0394x denotes the maximum deformation of the fingertip material as shown in Fig.2. Equation (3) is no more a constraint, because it can be regarded that the reproducing force F of deformation arises in the normal direction to the object surface as a nonlinear function of \u0394x together with viscous-like force. Therefore, even if rolling of contact is taken into account, the net DOF of the system of Fig.2 becomes at least of one, that enables to stabilize rotational motion of the system at a state of force/torque balance. The details are discussed in the following by illustrating a more general case of use of a 2 DOF finger shown in Fig.3. III. IMMOBILIZATION BY A ROBOT FINGER WITH SOFT TIP The kinetic energy of the system of Fig.3 is expressed as K = 1 2 q\u0307TH(q)q\u0307 + 1 2 I\u03b8\u03072 (4) where q = (q1, q2)T, H(q) stands for the inertia matrix of the robot finger, I the inertia moment of the object around the z-axis at the fixed point Om", + "(21) can be equivalently expressed as d dt (K + \u0394P ) = \u2212c\u2016q\u0307\u20162 \u2212 \u03be(\u0394x)\u0394x\u03072 (26) This expresses Lyapunov\u2019 relation since K +\u0394P is positive definite. Thus, applying LaSalle\u2019s invariance theorem to the relation of eq.(26), we can prove that as t\u2192\u221e { q\u0307(t) \u2192 0, \u0394x\u0307(t) \u2192 0, \u03b8\u0307(t)\u2192 0 q(t) \u2192 qd, \u0394x(t) \u2192 \u0394xd, \u03b8(t) \u2192 \u03b8d (27) and, at the same time, f(t) \u2192 fd (= f(\u0394xd)) and \u03bb(t) \u2192 \u03bbd as t\u2192\u221e, where \u03bbd can be determined as \u03bbd = fdYd/l (28) Numerical simulation of maneuvering the overall fingerobject dyanamics on the single DOF model of Fig.2 (eqs.(16) and (17)) by using the proposed control input of eq.(18) was carried out. Physical parameters of this simulation are shown in Table I and parameters of control signals are shown in Table II. Figure 4 shows that convergences of key variables to constants shown in eq.(27) are established within 0.2 seconds. Furthermore, in Fig.4, the graph (l) shows that equation (28) is satisfied as t \u2192 \u221e, because \u03b1\u2032 and \u03b1 are defined as \u03b1\u2032 = tan\u22121 (\u03bb/f) and \u03b1 = tan\u22121 (Y/l) respectively. This result means that \u03bbd/fd coincides with Yd/l as shown in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002704_6.2007-6821-Figure3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002704_6.2007-6821-Figure3-1.png", + "caption": "Figure 3. Vehicle lateral translation. The ALTAV MkII translates via thrust differential (affecting pitch and roll, and thereby inducing a horizontal thrust) or by vectoring the thrust motors fore or aft.", + "texts": [], + "surrounding_texts": [ + "The Quanser ALTAV Mk II, shown during hover tests in Figure 1, is the extension of the positioning potential of the Mk I to a more effective vehicle platform for payload capacity, endurance and in particular efficient translation, manoeuvrability and handling under disturbance. The vehicle employs four vectoring thrusters placed along the equator of the vehicle envelope. The masses are specifically chosen to collocate the centre of gravity (CoG) with the centre of buoyancy (CoB) of the vehicle. In addition, the actuators are placed such that the centre of thrust when outputting maximum thrust is along the centreline of the vehicle. While the concept of four vectoring thrusting power-plants as used on the Quanser ALTAV is similar to vehicles such as the Piasecki helistat17 or advanced vehicle designs such as the Lockheed HAA (High Altitude Airship)15, there are several key distinctions. With most versions of BQR (Buoyant Quad-Rotor)17 and other lift assisted helicopter concepts, the power plants are placed below the centre-line of the airship envelope (or the centre of buoyancy). This is not desirable as it radically increases the thrust required to affect pitch or roll motions. Positioning is also very difficult as wind disturbances (in the horizontal plane) are actually positively reinforced by the vehicle. For these reasons, the centre of gravity of the vehicle should be as close as possible to the centres of buoyancy and area of the airship which is unlike the design of the helistat. Though this will make the vehicle less \u201cstable\u201d from a controller point of view, it will make it much more robust in terms of disturbances as has been demonstrated through previous generations of Quanser ALTAV systems. Thruster position relative to both the centre of gravity and the centre of area are also crucial. When accelerating, if the thruster positions are not in the same plane as the centre of gravity there will be an induced rotation of the vehicle which is obviously not desirable, though this can be attenuated through restrictions on accelerating. An even American Institute of Aeronautics and Astronautics 4 greater effect can result should the thrusters not be located in the same plane as the centre of area, or more specifically, the centre of drag. This can produce strong rotations resulting from the moment between the drag and the thrust which will not dampen during the course of operations. The Quanser ALTAV concept addresses these issues concurrently: the centre of gravity is placed near the centreline of a radially symmetric non-rigid envelope. The thrusters are placed along this centre-line using a system of mounting feet and guy-lines to ensure a solid mount using the resultant side-wall pressure of the airship itself. This allows for much higher velocities to be achieved and does not make the vehicle susceptible to positive reinforcement of disturbances. It should be noted that traditional airship types such as the Wasp or Hornet by Advanced Hybrid Airships USA can compensate for these effects and operate at high speeds through the use of aerodynamic control surfaces, but these types of vehicles are notoriously difficult to handle at launch and recovery owing to their inability to effectively operate at low speeds. The unique control and actuation concepts behind the Quanser ALTAV vehicles allows excellent control performance at both high and low speeds. In addition to the physical airframe configuration, the other main difference between vehicles such as the helistat and the Quanser ALTAV series is the thrust vectoring. While many airships do use thrust vectoring, they do not combine this capability with the precision hover capability which arises from the differential thrust capability demonstrated by the Quanser ALTAV." + ] + }, + { + "image_filename": "designv11_69_0003301_amc.2010.5464114-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003301_amc.2010.5464114-Figure2-1.png", + "caption": "Fig. 2. Concept of the control.", + "texts": [ + " In this way, each parson can feel each other as if the person, who is in distance side, is in front. In order to realize above system, the control scheme between the exoskeleton system and endoskeleton system is necessary. It is necessary that the control is implemented in common space because there has the different structure. The endoskeleton system is supposed that it is like human structure. Therefore, in this paper, the bilateral control based on human modal space is proposed. The conceptual diagram of this proposal is shown in Fig. 2. The position (angle) information and the force (torque) information acquired by the actuators are transformed from real actuator space into the human modal space by human model, respectively. Then the bilateral control is implemented in the same human modal space. Finally, the acceleration references are transformed into actuator space. In this section, the transform from the actuator space into the human modal space is shown. The structure is shown in Fig. 3. In Fig. 3, x \u2208 n, F \u2208 n, \u03b8 \u2208 n and T \u2208 n stand for position, force, angle and torque respectively, and superscript ext, res, ref and \u02c6 stand for external force (torque), response and reference value and estimated value respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002698_iros.2007.4399277-Figure3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002698_iros.2007.4399277-Figure3-1.png", + "caption": "Fig. 3. Inverse kinematics of avoidance of a ball.", + "texts": [ + " Inverse Kinematics of Avoiding Obstacle The method for solving inverse kinematics of avoidance of elementary obstacles is described in this section, where it is assumed as follows: (1) a robot arm is composed of only rotational joints, taking account that translational joints are rather inappropriate for avoiding arbitrary obstacles, (2) each link is a cylinder, the radius of which is c . There is no offset in the robot mechanism, (3) the radius of any joint is smaller than that of a cylinder of the link. The links and joints from the base are called as \u201cL0-J1-L1-J2-\u2026-Jn-Ln\u201d. B.1. Avoidance of a point (ball) When the link LM contacts with a ball of which radius and center are Bc and rB respectively, the link LM must be located apart from rB by the distance of more than Bc c+ by rotating the joint Jk (see Fig. 3). Assuming the joint angles besides k\u03c6 are known, k\u03c6 realizing contact limit is solved by coordinate transformation matrices. There are four patterns of solutions, which are inside and outside of the cylinder of link LM and right and left rotations of Jk. Therefore, it is necessary to select one solution satisfying the given condition among the four. B.2. Avoidance of a straight line (cylinder) When the link LM is avoiding a straight line of which position is sr , the link LM must be located apart from the line by the distance of more than c (in case of cylindrical obstacle, it is c + cB) by rotating joint Jk (see Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000006_jsvi.2002.5135-Figure3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000006_jsvi.2002.5135-Figure3-1.png", + "caption": "Figure 3. Two neighboring segments and the connecting beam element.", + "texts": [ + " Msi is the generalized mass matrix of the ith segment Msi \u00bc diag\u00f0msiI3; J\u2019si\u00de; i \u00bc 1; 2; ; n; \u00f03\u00de where msi is the mass of the segment i; J\u2019si the inertia tensor of the segment about its center of mass and I3 is the identity matrix of 3 3.Qsi \u00f0i \u00bc 1; 2; ; n\u00de is generalized force of the segment including the contributions from the connecting beam elements, which are introduced in the following section. 3.1.2. The elasticity modelling To account the elastic energy in the deformed flexible beam, the beam element is introduced. Consider two neighboring segments Bi and Bj in Figure 3, BiXiYi and BjXjYj are, respectively, the body frames with the origin at the mass center of the two segments. The element BiBj is the elastic beam from Bi to Bj; at both ends of which, the beam has displacements consisting of segments Bi and Bj; respectively, see Figure 3. The kinetic energy of the beam is expressed in terms of the rigid segments and the elastic energy is described by the element. For the sake of simplicity, the two segments with equal length are considered. Let Oe be the reference frame of the beam element. Its global position vector re and angles ye between the Xe and X axis are re \u00bc \u00f0ri \u00fe rj\u00de=2; ye \u00bc \u00f0yi \u00fe yj\u00de=2; \u00f04\u00de where yi and yj are, respectively, the angles Xi and Xj with respect to the X-axis. The global position vectors and angles of the points Bi and Bj with respect to the point Oe are ai \u00bc rpi re; aj \u00bc rpj re; ye i \u00bc yi ye; ye j \u00bc yj ye: \u00f05\u00de In the co-ordinate system of the beam element, the above vectors are ae i \u00bc RT e ai \u00bc xe i ye i ( ) ; ae j \u00bc RT e aj \u00bc xe j ye j ( ) ; \u00f06\u00de where Re is the transformation matrix from the beam element to global co-ordinate system. For the beam element BiBj shown in Figure 3, let l; E and I be, respectively, the length, elastic coefficients and area moment of inertia. Considering lateral deformation of the fictitious element in the co-ordinate of Oe; the following equations and bound conditions are satisfied [18]: EI d4ye dx \u00bc 0; ye \u00bc ye i ; dye dxe \u00bc ye i while xe \u00bc l=2; ye \u00bc ye j ; dye dxe \u00bc ye j while xe \u00bc l=2: \u00f07\u00de As in the conventional finite element method, the force\u2013displacement relations of the beam element can be denoted as Pe si Pe sj ( ) \u00bc Ke de si de sj ( ) ; \u00f08\u00de where de si \u00bc xe i ye i ye i 8>< >: 9>= >;; de sj \u00bc xe j ye j ye j 8>< >: 9>= >;; \u00f09\u00de Pe si \u00bc F e xi F e yi Me i 8>< >: 9>= >;; Pe sj \u00bc Fe xj F e yj Me j 8>< >: 9>= >;: \u00f010\u00de When the relative deformation of the beam element is small, it can be described by the linear theories" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003881_j.precisioneng.2010.10.003-Figure8-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003881_j.precisioneng.2010.10.003-Figure8-1.png", + "caption": "Fig. 8. Apparatus and principle of calibration.", + "texts": [ + " Moreover, the cost of a BPF is low compared with that of a lock-in amplifier. 4.2. Calibration In a previous study [6], the calibration was carried out by applying static torque to the output shaft of the spindle. However, because the magnitude of the generated electric charge is proportional to the strain rate in the piezofilm, it is difficult to measure static strain while keeping a constant load. Therefore, the following calibration procedure was developed. A cylindrical workpiece is fixed to a torque sensor (Kistler, Type 9329A) as shown in Fig. 8. Vb and the output of the torque sensor are recorded at the same time when the end mill moves circularly and cuts the cylinder side of the workpiece. The cutting force is denoted as Fc, where Rd is the radial depth of cut. The torque measured by the torque sensor at the workpiece is the product of F \u2032 c , which is the tangential component of Fc at the workpiece, and rw. can be obtained geometrically and changes according to Rd. Fc, Fc and are expressed as follows. Fc = Tc re (2) f c 30000 400 40030000 ency Hz -0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002701_iccas.2007.4407016-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002701_iccas.2007.4407016-Figure1-1.png", + "caption": "Fig. 1 Ground test rig of 4-degrees of freedom", + "texts": [ + " Some problems of flight control system were found and fixed during the ground test including a hidden interface conflict between pilot control box and flight control computer and RPM sensor failure caused by disconnection of power plug. The designed gains in the rotor governor and rate SAS were also tuned during the ground test. 2. GROUND TEST EQUIPMENT The most dominant flight control axes of helicopter during hover flight are the pitch, roll, yaw and heave (vertical). Therefore the 4 degrees of freedom test rig shown in Fig. 1 would be useful to evaluate the critical flight characteristics of hovering rotorcraft within reasonable motional limits. The inertia moments of ground test rig should be measured to evaluate the effect of ground test rig on the aircraft motion. The measured inertia moments of 40% scaled aircraft and ground test rig were shown in Table. 1. Table 1 The Moments of Inertia MOI (kg-m2) Scaled Aircraft @ Tilt = 90deg GT rig Ixx Iyy Izz 6.93 4.09 5.45 29.281 2.187 44.640 The moments of inertia in roll and yaw axes were much bigger than those of 40% scaled airplane" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000887_1.1897746-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000887_1.1897746-Figure2-1.png", + "caption": "Fig. 2 The linkage\u2019s line of symmetry in relation to opposing joint screws.", + "texts": [ + " 2 The Kinematic Tools The Bennett linkage is depicted skeletally in Fig. 1 and, just as in Refs. 2,4 , we choose its frame of reference to comprise link DA and the pin at A. Notation relating to the loop is sufficiently explicated by the diagram and we make use of Bennett\u2019s index p = a/s = b/s . We also employ the abbreviations s for sine and c for cosine. The axis of symmetry of the joint-screw velocities 4 is of great relevance here; its disposition relative to a pair of alternate axes is displayed in Fig. 2. We rely upon screw vector algebra in a dual format and represent a generic screw by Contributed by the Mechanisms and Design Committee for publication in the JOURNAL OF MECHANICAL DESIGN. Manuscript received May 22, 2004; revised September 8, 2004. Associate Editor: G. R. Pennock. Journal of Mechanical Design Copyright \u00a9 20 rom: http://mechanicaldesign.asmedigitalcollection.asme.org/pdfaccess.as S = S + S* = S n + M* , where 2=0. In this set of denotations, n is a unit vector with the direction of the ISA" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001585_iros.2006.282587-Figure3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001585_iros.2006.282587-Figure3-1.png", + "caption": "Fig. 3. Three damped springs control three DOF arm that models a humanoid limb.", + "texts": [ + " The torques applied to the joints will drive the arm along a trajectory. The expression is given in Equ. (4). \u03c4 = I(\u03b8)\u03b8\u0308 + C(\u03b8, \u03b8\u0307)\u03b8\u0307 + G(\u03b8) (4) where \u03c4 is a 3\u00d7 1 vector of joint torques applied to shoulder, elbow and wrist, I(\u03b8) is a 3\u00d73 matrix representing the kinetic energy, C(\u03b8, \u03b8\u0307) is a 3 \u00d7 3 matrix of centrifugal and Coriolis effects, G(\u03b8) is a 3 \u00d7 1 vector of gravitation. \u03b8, \u03b8\u0307, and \u03b8\u0308 are vectors of joint values, velocities and accelerations. The required torques in Equ. (4) are generated by three spring muscles in Fig. 3, where each spring controls one DOF. Although actual muscles that control human limbs are much more complicated than the springs here, our simplified muscle model captures the main scope of its movements. To model the muscle forces, a well established method is using damped springs to simulate the dominant mechanical behavior of muscles. The forces are dependent on the muscle length and its rate of change as represented in Equ. (5): F = K \u00d7 (l \u2212 lnatural) + B \u00d7 l\u0307 (5) where K and B are the spring stiffness and viscosity parameters respectively, l is the current spring length, lnatural is the spring natural length, and l\u0307 is the velocity of spring length" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000210_iros.1996.571043-Figure5-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000210_iros.1996.571043-Figure5-1.png", + "caption": "Figure 5 : Position error of pseudo center with respect to Col", + "texts": [ + " This control rule determines the linear and rotational velocities of the mobile robot according to the position and orientation errors with respect to the desired trajectory expressed in the coordinate frame fixed on the robot, and guarantees that the actual trajectory of the robot converges uniformly asymptotically to the desired trajectory. Let CO, denote the coordinate frame with its origin at the pseudo center and its X axis parallel to the line which joins the contact point and the center of friction (see Figure 5). Then, we consider the position error of the pseudo center expressed in CO, with respect to the desired trajectory. The error is denoted as O b e = [O'xe, \"'ye],. Denoting the actual position of pseudo center expressed in C u as \"p,,, the desired position as and the rotational matrix from CO) to CTJ as \"Ro,, the error O p e is given by (9) Denoting the orientation error as 8, = 8 d - 0, and adopting a differential feedback rule described by ] > (10) [ a ; ] = [ Bo! d COS Oe f Kz O'Xe 0 d + @,/d(KYo'ye + KO sin 0,) it is guaranteed that e," + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001627_iros.1989.637927-Figure5-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001627_iros.1989.637927-Figure5-1.png", + "caption": "Figure 5: Multiply-Connected Backprojection", + "texts": [ + " Mark every goal vertex such that it is possible to slitlc away from the vrrtex on the non-goal edge. 2. At every marked vertex erect two rays parallel to the edges of the inverted control uncertainty cone. Compute the intcrscction of thcsc rays among themselves and with CContact. Interrupt each ray beyond the first intersection. 3. Beginning a t the goal edge trace out the backprojection region. Thr operations of this algorithm are illustrated in Figure 4. The above algorithm can only construct simply connected backprojection region, which may be a limitation. For example, consider Figure 5. The result of applying the algorithm to the edge 7\u2019 is shown in 5.a. T h e maximal backprojection is shown in 5.b. Indeed, if the robot configuration reaches the vertex denoted by X, it non-deterministically slides on one of the two edges abutting at X and, in both cases, ultimately reaches 7d. An extension of Erdmann\u2019s algorithm to handle this kind of situation is presented in [12]. Extending the algorithm to a collection of edges is not simple, since the union of the backprojections of different edges may not be a maximal backprojection" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003815_s147355041000025x-Figure12-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003815_s147355041000025x-Figure12-1.png", + "caption": "Fig. 12. Sacrum", + "texts": [ + " Hydrostatic collapse causes the ventral half of the outer layer to fold over upon the dorsal half, forming the zygomatic arch and prefrontal ridges. The remaining ventral segments of the apical cap fold underneath the presumptive mandible (Fig. 19). The sacrum is formed from the caudal cap comprising five concentric zones divided by four meridians, with a hole in the centre of each. The coccyx of the spinal column comes to rest in the centre of the caudal cap, covering the dorsal half. The ventral half folds back onto the dorsal half, completing the form of the sacrum (Fig. 12). The presumptive alimentary canal results from the segments of the surface entering the internal canal at the poles that meet interiorly to form the gut tube. An aneurismic inflation forms the stomach. The segmented large intestine is formed from the incursion of the segments of the caudal cap (Fig. 13). The subject model is of the human species ; it is also applicable to the other mammalian species. The fact that reptiles and other predecessor forms, except for birds, do not possess a sternum does not exclude them from the paradigm" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001826_6.2006-6703-Figure4-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001826_6.2006-6703-Figure4-1.png", + "caption": "Figure 4. Control system, method of realizing.", + "texts": [ + " At spinning object one channel is used to control both angles: of attack and side-slip. These assumptions can be realized by gasdynamic impulse interaction on the object gravity center. This solution gives us not only quicker object response on seeker information and follows it makes simpler servo control system but also more precision guidance. The complicated mechanics of aerodynamic servo are not needed. In that solution control is realized by correction engines located around the flying object center of gravity (figure 4). When the target is selected, it is tracked during the rest of the flight of the projectile. The error \u03ba between the center of the target and foreseen impact point of missile is continuously monitored. As soon as this error or it\u2019s time derivative exceeds a reference value, one or several rocket correction engines are fired in a direction to minimize the value of the error. The impulse of the rocket correction engines passes through the center of gravity of projectile, which gives instantaneous course correction", + " Task of the rocket engines is to correct the course of the projectile in the last phase of the trajectory, homing it to the target, achieving a direct hit. Correcting rocket engines are located in a cylindrical unit, arranged radially around the periphery. Each one of the correction rocket engines can be fired individually in a calculated sequence. The correction engines set is placed close to the center of gravity of the projectile. When the rocket engine is fired, the course of the missile is changed instantaneously. Way to course change shows figure 4. By successive firing of several rocket engines, the projectile is steered with high precision on to the target. The chosen steering system gives a very fast response to the guidance signals. The decision when the correcting rocket engine is fired depends on the value of the reference control error. The frequency of the correcting engines firing N, which is defined as the missile rotation number between correcting engines firings, increases according to the control signal value K. The direction of control forces depends on time of firing the control engine", + " Results are especially good when flight trajectory is steep (pitch angle \u0398 is higher than 60o). But is a problem with \u201cflat\u201d trajectory at maximal range of fire. When range is higher than five kilometers quality of guidance process give the poorer results. Flights with flat trajectories often give the result in missiles dropping before target. In extreme cases reaching the target is impossible. Missile is controlled with PD control law. Steering signals from regulator depends on angle between main symmetry axis of missile and line setting from missile seeker to target center \u03ba (figure 4) and derivative d\u03ba/dt. This is no difference for control system if target is detecting above or below of projectile main symmetry axis. In simulation we recognized that impulse guiding system, like one described earlier, is ineffective when pitch angle \u0398 is lower than 60o. Our impulse control system consist of 12 impulse correction engines and guidance process has good quality only with steep trajectory for \u0398>60o. During attack faze when target is detected below missile main symmetry axis too early decision about using correcting rocket engines system can get worse trajectory of flight (makes it more flat)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003266_detc2009-87020-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003266_detc2009-87020-Figure2-1.png", + "caption": "Figure 2. Schematics of a torsional system with lumped inertias and torsional elastic springs.", + "texts": [ + " The values of the loss factor associated with the coupling and the engine are supplied by the manufacturers as 0.124 and 0.08, respectively. A relatively small value of 0.0015 is assumed for rest of the driveline components. The equations of motion (EOMs) of the test cell system are derived by representing it as multiple lumped inertias, 1 2, , , ,\u22ef nJ J J connected with torsional elastic springs with stiffnesses, 1 2 1, , , .\u2212\u22ef nK K K The schematics of such a pure torsional system with n degree-of-freedom (DOFs) or inertias and n-1 torsional springs are shown in Fig. 2. The electrical drive rotates the system by providing a torque ( )t\u03c4 at the first disk. Euler\u2019s law is applied to each of the disks to obtain n coupled EOMs. When the drive torque ( )t\u03c4 is applied, the reaction torques ( )1 1 2K \u03b8 \u03b8\u2212 \u2212 and ( )1 1n n nK \u03b8 \u03b8\u2212 \u2212 \u2212 result at mass 1 and mass n, respectively. However, for mass 2 to mass n-1, a reaction torque of ( ) ( )1 1 1 2 ,+ + + +\u2212 \u2212 \u2212i i i i i iK K\u03b8 \u03b8 \u03b8 \u03b8 is resulted, where 2 1.\u2264 \u2264 \u2212i n Equating the equal and opposite torques at each mass, n distinct EOMs are formed as, ( ) ( ) ( ) ( ) 1 1 1 1 2 2 2 1 1 2 2 2 3 1 1 ( ) " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002200_09544054jem699-Figure4-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002200_09544054jem699-Figure4-1.png", + "caption": "Fig. 4 Sketch of an SSM system", + "texts": [], + "surrounding_texts": [ + "The constraint condition of a TP, namely equation (8), specifies the lower limit of the information content of a TP, whereas the upper limit of the information content of a TP is constrained by the control condition, controller choice, and price cost, as well as the control efficiency to be discussed hereafter. To measure the control efficiency, a measure of the system function is needed. In this paper, the summation of the information content of the FR is used as the measure of system function according to IR \u00bc X j IRj \u00f013\u00de In the above equation, j indicates the number of FRs and IR is the measure of the system function, which is comparable between different design schemes. Based on IR, the control efficiency between different designs can be evaluated by defining the control efficiency h as h \u00bc IR M \u00f014\u00de h represents the control efficiency of the product control system. When M is decreased by reducing the number of TPs or TP information content, or when IR is increased by integrating or reconstructing structural components, h will increase. The control efficiency h is an important parameter to measure the design quality of a mechatronic system. By this standard, the second principle of the mechatronic design can be proposed: the maximumcontrol-efficiency principle states that an appropriate design for mechatronic system, which satisfies the independence axiom, the information axiom, and the constraint condition of the information content of a TP, is the design with maximum control efficiency. Since the comparison of the information contents of functions is constrained to be that of congeneric functions, the comparison of the control efficiencies of TPs is constrained to be that of congeneric systems, especially between different design schemes of the same system. The above restriction does not hinder the application of the maximum-controlefficiency principle, which is needed in system design to make a choice between many schemes. Whether a design scheme is good or bad can be confirmed only through comparison with other schemes of the congeneric systems. There is not an absolute criterion which can cover all system functions." + ] + }, + { + "image_filename": "designv11_69_0003303_j.ejc.2009.09.007-Figure8-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003303_j.ejc.2009.09.007-Figure8-1.png", + "caption": "Fig. 8. Fold-in and pull-out operations.", + "texts": [ + " Among the two regions of S divided by \u03b3 , let \u0393+ be the region corresponding to the red-face-side ofM , and \u0393\u2212 be the region corresponding to the blue-face-side ofM . Suppose area(\u0393+) > area(\u0393\u2212) inM \u2032. Then, by reversingM \u2032, we have area(\u0393+) < area(\u0393\u2212). Hence, in the midway of the deformation, it happens that area(\u0393+) = area(\u0393\u2212). However, since \u03b3 is shorter than the great circle, this is impossible by the above lemma. Let us introduce here a few special origami-deformations related to a rectangular tube. (1) Fold-in- and pull-out-operations. By subdividing a rectangular tube suitably, we can \u2018fold in\u2019 a part of the tube as in Fig. 8. Let us explain a little more. In Fig. 8 left, put x = OC, y = AC = BC , and let a\u00d7 b be the size of the base rectangle. Then y < min{a/2, b/2}. In order to fold in as shown in Fig. 8 right, the three verticesA, C, Bneed to become collinear in themidway of deformation. Hence, if y/x > \u221a 2, one of A, B goes outside the a\u00d7b rectangle in themidway of deformation. But, if y/x < \u221a 2, then A, B can remainwithin a\u00d7b rectangle. (This is important to introduce fold-out-operation.) If y/x < \u221a 2\u22121, then A, B cannot go down to the level of O. Hence we also assume y/x > \u221a 2 \u2212 1. If we take x, y to satisfy y < min{a/2, b/2} and \u221a 2\u2212 1 < y/x < \u221a 2 then we can fold in (and pull out) the tube by length x, with keeping A, Bwithin the a\u00d7 b rectangle" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001937_epepemc.2006.283096-Figure5-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001937_epepemc.2006.283096-Figure5-1.png", + "caption": "Fig. 5. Flow of the generation of A1", + "texts": [ + " To realize an open-loop amplifier with high linearity and low output impedance, the amplifier in Fig. 4(a) is considered to consist of two voltage amplifiers, shown in Fig. 4(b) and (c). If the nonlinearities of amplifier Amp1 and Amp2 are reversed, their product can achieve high linearity, as discussed above. Amplifier Amp2 is a conventional common-source amplifier. Its gain A2 is decreasing as the input amplitude becomes larger, which means \u03b22 is negative. Thus, to compensate the nonlinearity of A2, \u03b21 should be positive with a proper value. Fig. 5 shows the generation of A1 intuitively. Assuming M7 and M8 are ideal current sources, the gain between VS1 and V + I N can be described as (6), which is derived from (4) and (5). gm1 ( V + I N \u2212 VS1 ) + (V \u2212 X \u2212 VS1) / rO1 = 0 (4) VS1 / rO3 + gm3V \u2212 X + (VS1 \u2212 VS2) / RS = 0 (5) VS1 / V + I N = 1 1/ rO3+2/ RS gm1gm3rO1 + 1 gm1rO1 , (6) where VS1 = \u2212VS2. As in the conventional amplifier, gm1rO1 decreases when input amplitude becomes larger. Therefore, VS1/V + I N is decreasing, as shown in Fig. 5(a). Because V + I N is also VG1, drawn as a dashed line, the difference between the dashed line and VS1 is the gray area, VGS1. The gray area indicates that, the gain between signal VGS1 and V + I N is increasing when input amplitude becomes larger, as shown in Fig. 5(b). Because the current in M1 is almost constant, VDS1 is proportional to 1/(VGS1,2 \u2212 VT H )2. Therefore, the area drawn in gray is amplified and the gain between signal VDS1 and V + I N is increasing when input amplitude becomes larger, as shown in Fig. 5 (c). As a result, the gain of amplifier Amp1 is also increasing when input amplitude becomes larger. Therefore, a positive \u03b21 is realized. Until now, a positive \u03b21 and a negative \u03b22 are obtained, so the linearity of the amplifier is already enhanced. To further suppress the nonlinearity, the values of \u03b21 and \u03b22 need to be matched, described as follows. The signal VD1 is feedback to VS1 through transistor M3. The feedback factor is described by (7): Loop gain \u2248 gm3(rO3||RS) (7) As shown in Fig. 5, curves VS1, VGS1, and VDS1 are in the feedback loop, therefore the shapes of these curves are affected by the feedback loop gain. In (7), the value of gm3, rO3, and RS can affect the nonlinearity by adjusting the local loop gain. In the design, gm3 is decided by the current and VGS3 \u2212 VT H in Amp1, which are designed under the consideration of power efficiency and transistor\u2019s operation region. The value of RS is decided by the amplifier\u2019s gain and RL , where small resistances are preferred. Thus, the parameter rO3 is selected to adjust the loop gain in the design" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002165_bf02153076-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002165_bf02153076-Figure1-1.png", + "caption": "Fig. 1. The shaft in the undisturbed state, showing (grossly exaggerated) initial bend.", + "texts": [], + "surrounding_texts": [ + "SHAFTS TO EXCITATION Gianfranco Capriz * S O M M A R I O : Per prevedere con accuratezza /'ampiezz a di vibrazione di un asse lievemente sbilanciato che ruoti sd supporti lubriflcati ~ necessario rappresentare adeguatamente il comportamenlo dei supporti stessi. Un modello /ineare ~ sufflciente per molti scopi pratici purch~ si tenga conto chela risposta non ha generahnente simmetria assiale. D i conseguenza ~ necessario poter considerate casi nei quail la traiettoria di ciascun punto dell'asse ~ ellittica piutlosto che circolare. Percib si richiede una generalizzazione de/metodo di K#nig; generalizzazione che ~ indicata nel presente lavoro. II metodo proposto fa perb ancora uso di semplici operazioni matriciati. S U M M A R Y : To obtain an accurate forecast of the bdmviour of a shaft rotating on lubricated bearings under excitation, an adequate model must be introduced to represent the response of the bearings; a linear model of the response is suffiident for many practical purposes, but allowance must be made for the variation of film stiffness, cross-stiffness, etc., with direction. A s a consequence the shaft must be assumed to move in an elliptical, rather than circular, whirl. The phenomenon which must be attaljzed is thus more complex than that envisaged in K#nig's study, for instance, see Ref. [2]. IVe show here nevertheless that it is possible to devise an appropriate generalization of the Myklestad-Holzer method, so that )be problem can be solved bj, matrix manipulations. 1. Introduction. The investigation was prompted by the following cortsiderations: I) To base the design of a shaft largely on the calculation of its critical speeds is at times misleading since damping may allow the running of a well-balartced shaft close to or even at a critical speed. II) It is of interest to have means of predicting the behaviour of a shaft under accidental unbalance, such as may be caused by the stripping of some blades in a turbine. III) The resportse of a shaft to disturbance is greatly influenced by the characteristics of its bearings: hence a careful representation of these is important. The peculiar properties of fluid films usually lead to elliptic motion. The paper is divided into sections. Sects. 2, 3, 4 are devoted to establishing the basic equations; a method (of the Myklestad-Holzer type; see, for instance, Re\u00a3 [1]) for the numerical solution of the problem is then described, having in mind, of course, the use of a computer. In * Professore di Meccanica Razionale, Centro Studi Calcolatrici Elettroniche, Universit8 di Pisa. Sect. 6 the problems arising from bearings effects are tackled. To achieve compactness many details (which are called for to plan a flexible computer programme) are left out of this paper: for instance distributed damping and magnetic forces are not mentioned here. We refer to the internal report of Re\u00a3 [9] for these and other questions. 2. Dynamic equations. The analysis is based on the simple theory of beams (Refs. [3], Ch. XVIII ; [1], Ch. IV; [4], Chs. V, VI) with the additional consideration of shear and gyroscopic effects (Ref. [41, nos. 48, 55; Ref. [51). However, a brief restatement of the principles involved is in order, as it gives the opportunity for precise definitions. A fixed system of reference s* (origin I2; unit axial vectors cl, c~, cs) is chosen so that the third axis (axis of ~ goes through the line of centres of two bearings, but either the line of the elastic centres C or the line of the centres of mass G (of elementary slices) may be different from this axis even in the undisturbed state (in the first case due to a permanent bend, in the second due to lack of balance). Because of this we introduce the notatiort C. , G, for the positions of C and G in the undisturbed state. To calculate correctly the elastic forces and moments arisirtg during whirl the displacement w(z, t) must be interpreted as the displacement at the instant t of the point C (relating to the cross-section of coordinate Z). At the same time a local measure of unbalance is given by the vector e(z) which joins O ( ~ (the point of the crosssection of coordinate Z, on the axis) to G.(Z). I f w0(~ is the displacement due to permanent bending, then: e(O = w0(~) + C,G,. DECEMBER 19~ 213 A second system of reference s(g2, in, i2, is = ca) is also introduced, which rotates with the shaft at the same steady speed ~o (radians per unit time) around the third axis. Calling the mass ttAc]g, the resultant external force R(eI and i ( \" ) the rate of change of moment of momentum (per unit length of the shaft), the equations of motion of the shaft are: c)S /,M~ O~-w OZ g Or2 c)M ..x(m) = S X c a + ~ v x , o.~ (2.1) where S is the shear force and M the bending moment. The external forces contributing to R(e) are unbalance, damping and, in rotors of electrical machines, magnetic forces. For simplicity, only the first source of excitation is considered here. The vector e is usually known through its components on the moving frame: e = elil + e~i2, he.rice M 0'\u00b0 = j3~rcs X 0 + v2j[(0 \u2022 i 2 ) i l - (0 - i,)i2]; in our case, of course, co = v. (2.6) 3. Dependence of d i s p l a c e m e n t s on forces a n d moments . Boundary c o n d i t i o n s . When it is deemed convenient to account for the direct effect of shear force on the displacement in a bar, the usual relationship between bending moment and curvature of the displaced centre line is complemented by the following assumption: the displacement w can be split into the sum of two components w , , w**, of which the first obeys the usual relationship : (~2W. M = Efc3 \u00d7 OZ ~ (3.1) (/\u00a3, modulus of elasticity; I, second moment of area o f cross-section), whereas the derivative of the second is proportional to the shear force and we will express forces due to lack of balance in terms of el~ e2: MoB c)w** S = - - (3.2) k 0Z R(e) -----/,.d___._% co~_(eii 1 + eziz). (2.2) g An explicit expression for M(,\"~ can be based on the observation that the motion of an3, elementary slice of the shaft is approximately rigid, with angular velocity p t i l +p2i2 +paiz , if the vector p = p l i l +p.,i.~ is defined through the formula: 0 (O(wo + w ) ) (2.3) OZ Ot = p \u00d7 ca, and pa = o,. Using the notation j ldz , j~dz, jadg(j j = j~_ = j ) for the principal moments of inertia of art elementary slice of the shaft of thickness dz, (calculated as though there was rto permanent bending) we have then for MO,,l: M ( \"~= ~ , - - (j.,+,--j,+e)p,+,p,~.2 i , , (2.4) where the indices must be interpreted cyclically. Naturally throughout these formulae second order terms are dropped, which would be required to account for lack of exact parallelism betaveen the principal axes of inertia of the slice during motion and the unit vectors il, i2, in. From (2.4), calling 0 the slope, we get: M ( , , , . ao 0 2 0 = j3o) -~ + y e a \u00d7 01-----7- (2.5) But during a circular whirl of steady angular speed re3 00 0z0 Ot = yea X O, 0t~ - - 7'2e' (B, shear modulus; _/1~, area of cross-section; k, Timoshenko's constant). Eqs. (3.1), (3.2) imply that M and w can be put into a direct relationship: o.,w M = E/ca X - OZ,., 0 k s)l and there is no need to make further reference to w. , w**; we can use Eq. (3.3) and w only. Eqs. (3.t), (3.2) justify, however, a special notation, 0, for the slope due to bending 0 w k 0 = - - - S . (3.4) 0Z AcB No essential change would actually occur in the following analysis, if O were taken to be equal simply to 0w/0z; but with the notation (3.4) some simplification is achieved. Note also that rigour requires that w . rather thart w enters into Eq. (2.3), and although gyroscopic and shear actions are normally small so that there )s no strict rteed to account for compound effects, we will now use the formula: o(e + %) dwo c)t = p X ca, O0 = dz (3.5) rather than (2.3), to define pl and p2. The indefinite equations (2.1), (2.2), (2.4), (3.4), (3.5), must be associated with appropriate boundary conditions (at each end of the shaft) and with transitiort conditions (at the bearings). I f the shaft runs oa plairl lubricated bearings, fluid film effects are of great importance. Within the limits of our analysis they must be thought of as linear effects 214 MECCANICA (flexibility, cross-flexibility, damping, cross-damping), implying a jump in the shear force and bending moment at the bearing centre, which is a function of the components of displacement and speed. The differential system (4.3) must be accompanied by the bourtdary and transition conditions. 5. N u m e r i c a l m e t h o d . 4. E l l i p t i c synchronous whir ls . Wc assume here that each point of the centrc line of the shaft moves in an elliptic path, the axes arid the phase angle varying with the axial coordinate. The angular speed of the whirl is assumed to be equal to the speed of rotation w, so that the displacement can be written as follows: w(z, t) = [O,(x)(Z) coso)l +y(-\u00b0)('Z) sincot] c] + + [y(a)(Z ) cos,ol +.yI~)(Z ) siR:o/] c.o, (4.1) and parallel expressions hold true for 0, M and S: o(z, t) = [0( , (~) cos,or + 0(2)(~) sin:,,t] c~ + + [0(a)(z ) cos~0t + 0(.u(z ) simot] c2, M(Z, t) = - - [M(a)(Z) c o s , o / + M('t)(Z) sin~o/] c , + + [M(*)(Z ) cos(ot + M(2)(z) simot] c~, (4.2) S(Z, t) = - - [Sm(z ) cos~ol + S(\u00b0-)(Z) sin~ot] c t - - - - [S(a)(Z ) coswt + S(a)(Z) sin:,,/] ce. Here notation and signs are chosen so as to simplify, as far as possible, some ensuing formulae. The question naturally arises of the compatibility of expressions (4.1), (4.2) with the indefinite equations of Sects. 2, 3. It is not immediately obvious that gyroscopic forces for instance do not excite higher harmonics: simple calculations assure us of compatibility, however, and lead to a system of first order differential equations in the functions .),(~), 0(~), Mm, S(~): 4)'(*) = 0 ( * ) - - k S(\") (~ = 1, 2, 3, 4); dz A ~ B dO(s) II\u00a2I1~) & d114(1) & d21,B~) a< dMla) dM(~) & dS(U /5I (s = 1, 2, 3, 4); (4.3) s , , , - ,~{/10., + q , ~ ] - / d o . ' + ~,11. s, '~,- ,.:-{j[o(\"-,- *\"-I + j..[o(., + ~d}. S(a,-- ,,;'{jlo(a, + \u00a2-~l + ja[0\"- ')-- ~r._,]}, 3\"(a) - - o)~{j[0~\") + 9~,1--jalO(\" + e-,]} ; tt.Ac _ _ (u'-'(y(U + e 0 , etc. g The first group of these equations derives from the vector equation (3.4) which defines 0; the second group from the relation (3.1) between bending moment and curvature; the third group from the second Eq. (2.1), with the specification (2.4), (3.5) of the moment of momentum and calling ~a, ~. the components on s of 0o. Finally the last group derives from the first Eq. (2.1) with the specification (2.2) of the external force. The differential system (4.3) could be integrated numerically using any suitable general method. However, to exploit certain pecularities of the system (and also to streamline data requirements) the special procedure of Myklestad-Holzer is preferred. The application of that procedure to our system does not present difficulties; we quote here only the difference equations on which the numerical calculations must be based. The shaft is divided into sections corresponding to intervals on the z-axis: (Z0 = 0, Zl), (Zl, Z2) . . . . . (Zn, Z,,+I) . . . . ,(Z,v-1, ZN). The choice of Z, will be partly suggested by the details of rotor design (for instance it is convenient to have a section for each wheel of a steam turbine), but will also depend on the accuracy desired. The size 1, = Zn+l - -Z\" of the intervals may vary; it may even be convenient to introduce sections of zero length to deal more easily with discontinuities, such as arise through the scheme suggested at the end of Sect. 3 to deal with fluid film action at the bearings. The components of the shear force are taken to have the constant values S#(*) over the n-th section; 2l//n(s), 0,(*), y , m are approximate values of M(~), 0(*), .y(*) at Z,. Some constants, to be considered as given, are also required: )\"+'(Ez)-~ az. t~,, . (Ez)-, (z-- ~,,)~z, \" g n \" g n = ~ t l+ l . , j \"gtl+ 1 :'\" i (EI)-~ (~.-- ~,,)\"&., o,,= (.,,l~B)-* kd~Z, ~ n \" 2 n \"Zn+l ~'2n+l = 2 ~ 1 ~ 1 [ ' z n + t e,,.i ,~, q,, g ,,,A~eMz, (i = 1, 2). Then the difference equations are: (5.1) ),(,') = ),(,') + I o(~) - - t~,,M(\"),, - - 6 ' . + a )S~*) \u2022 1 ' + I \" ~1 II 11+~. ) (5.2) O(,') = O~ ~) + ,~ M(') + / / S (~) (s = 1, 2, 3, 4)\" n + 1 n n II - - l l )) ' ~1\u00a2.) = M~)) \u00b1 ] s ! ? ) - - ~ [0(- + %+lal + i1 + l T It ,, ~ tl t ) l + l + \u00a2 [0\") + %+~a] n L n + 1 [ ra) + e,.1] , etc. S(~) = 3'(~) + u, . . . . ,,+l etc.; Note that ~025,,, co2\u00a2n are approximately the central moments of inertia of the n-th section on the axes parallel to i l (or in.) and ia respectively if the slices are sufficiently thin; co-~q,, is the mass o f the n-th section; e,,a, ena are the components in the directions of il, i2 of the displacement of the centre of gravity of the n-th section. DECEMBER 19~ 2 1 5 6 . B o u n d a r y e f f e c t s . To achieve the numerical integrat ion of our differential problems, the difference equations (5.2) must be associated wi th condit ions which express boundary effects. I t is particularly simple to write these condit ions in the case of a shaft whose ends are free and which is supported on a number of perfectly smooth, self-aligning, rigid bearings. In that case the numerical integrat ion is started at a free end with a vector (,\";'I \u00b0'o\"1 g')I s,:)), whose last 8 components are null, whereas the first 8 are indetermined; these 8 quantities remain indeterminate throughout the step-by-step procedure, being finally determined by the condi t ion that also in the last vector ,(s) (-) x I 0g)IM591 a'~,) the last 8 components vanish. The ease when only rigid or elastic supports are envisaged is trivial, from the point of view of the present investigation; in fact, ir~ that case the response of the shaft to e.xcitation due to lack of balance is a circular whirl. Instead, elliptic whirls necessarily occur when the response at the bearings is of the general type: Ou(~) Oum) q ~ = q bnu(') + bm_u \u00a2-\u00b0) + b~a-7~-\" + b.) ,oOt 2 ' (6.1) - -aU + e---U + c,s + l , here Q;, N t are the components along cl, cz of the jumps at the bearings centre of the shear force and beading moment respectively and uCtl, u(~l are the components of w along the same unit vectors also at the bearing centre: u{l) =ye(1) cos w/ +yA-(2) sin(or, am) _--yx(a) cos(o/+.yA:(4) sinco/, etc. Such complex response must be envisaged if a reasonably accurate model is required for the effects of the lubricant film at the bearings. Then q and n, b~e and c;A- are constants that cart be specified analytically for the case of short plain cylindrical bearings (see Refs. [6] and [7]) and are given for many other complex cases in tables and graphs (Ref. [8], for instance). Subst i tut ion in (6.1) leads to the fol lowing condit ions for S~+I ( s ) - SA-(*~: b },(4) b j,(1) b ,(3) , S ( m - - S ~ = - - b ,\u00a22)+ . _ _ __ .) ] e+ l ~. q [ l i J k 1. o /\u00a2 18 A- 14 k S(3) - - S ( 3 ) ~ - q [ b ~A! 2~ I: 23 k 2/1 /\u00a2 - ,+1 k i r ~ + b ),ca) + b .),'~ + b rc4q, S\u00a2'~ - - S\u00a2 't~ = - - q [ b 3 '(2) + b y(,a) _ _ b r m - - b ),\u00a2z~] : ~+1 /~ 21 k 2~ k 23 A\" 2~ k - and parallel expressiorts for M,+ lm- -Mk\u00a2*l . These formulae, as the previous set (6.1), are writ ten explicitly only for the case when there is no initial bend in the shaft. Obvious changes occur in the more general case. The important remark is that Eqs. (6.1) cannot be satisfied for a general choice of but, c~t w h e n j es) ----- __ y(2), 3,(,t) =3a l l (circular whirl), hence the need for the developments of the present paper. Actually, wi th in the framework of our developments even more complex circumstances cart be simulated (such as, for instance, the case of a plain lubricated hearing on art elastic support); we refer again to the report Ref. [9] for details. Received 4 September 1967. A t Cl, C2, C3 C. , C = e el, e2 E, B g = G, , G = i l , i2, i3 = j~, j~, j3 = / . = L M = Zff\u00a2o = Af,,\u00a2') = M(m) = N = N , Q = O = p = pl, p 2 , ~ = R(O = LIST OF SYMBOLS area of cross-section of shaft unit vectors of fixed frame of reference positions of elastic centre of cross-section in the undisturbed and disturbed states vector which characterizes the lack of balance of an elementary slice: e = O G. components of e on the first two axes of moving frame modulus of elasticity, shear modulus acceleration due to gravity position of centre of mass of elementary slice of shaft in the undisturbed and disturbed state unit vectors of moving frame moments of inertia (per unit length of shaft) around axes parallel to il, i2, ia length of n-th slice in difference formulae: 1. = Z.+I - - Z, length of shaft bending moment components of M, see formulae (4.2) values of Aim at Z. rate of change of moment of momentum (per unit length of shaft) running index in numerical integration maximum value of n moment of the couple and force due to the lubricant and acting on the journal point the cross-section of shaft on the z-axis angular velocity of elementary slice of shaft components of p on the moving frame resultant external force (per un i t length of shaft) fixed frame of reference 216 MEGGAN I CA S S(\") = S,,(o = t = U = W = W , , W * * = w0 frame with third axis common with s., rotating at speed o0 shear force components of S, see formulae (4.2) approximate values of SO) at Z, time displacement of journal centre from steady state position displacement of C during vibration component vectors of displacement caused directly by bending and shear, respectively vector characterizing permanent bending: w0 = OC y\u00a2') = components of w, see formula (4.1) y.(') = approximate value of yco at Z. Z = coordinate along the shaft axis" + ] + }, + { + "image_filename": "designv11_69_0000965_2004-01-1723-Figure8-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000965_2004-01-1723-Figure8-1.png", + "caption": "Figure 8: Synchronous Belt Drive System", + "texts": [ + " In summary, periodicity in the analysis intervals for order analysis may be ensured by setting the interval length to a finite number of revolutions of the source of the vibration. However, frequency analysis on arbitrary time intervals requires the use of FFT windowing functions. Two applications requiring multichannel measurement and analysis will now be presented in order to illustrate the capabilities of RAS equipment. The primary purpose of synchronous belt drive systems is to synchronise camshafts and crankshafts. The belt system may also be used to drive integrated auxiliaries. In the 4-cylinder diesel engine shown in Figure 8 the crankshaft pulley drives the camshafts via a toothed belt. The idler on the tight side ensures that the belt wraps properly around the crankshaft's toothed wheel. The tensioner on the slack side provides tension. The fuel pump is also integrated into this drive system. The influence of the effective tension (dynamic tight minus slack belt tension) on the life and durability of the The rotational speed of the crankshaft, fuel pump and camshafts was measured by scanning their toothed wheels with magnetic proximity sensors" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001222_isie.2006.296103-Figure8-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001222_isie.2006.296103-Figure8-1.png", + "caption": "Fig. 8. Claw coupling, exploded view", + "texts": [ + " As the measurement results will show, the measurement of the position has to be carried out with a high resolution. For that reason the sinusoidal analogue signals of a 2048 pulse incremental encoder are used. The resolution reached is approximately 0,5 Mio. PPR. Fig. 7 depicts the mechanical setup. A permanent magnet synchronous machine with a rated torque of 15 Nm drives the mechanics. The machine is fed by an inverter with field oriented control. In order to enable thorough investigations in a wide range of backlash values a special mechanical backlash element has been designed and realized (see Fig. 8). The realized module is a claw-type construction [6]. It provides the possibility to change the backlash angle manually to certain values within a wide range. Due to wear, the contact elements of the backlash element, which are depicted in Fig. 8/4 are replaceable. This makes sure that the initial precision can be reinstalled after longer operation. The contact elements themselves are milled very precisely. By twisting the positioner (Fig. 8/2) between both halfcouplings (Fig. 8/1, Fig. 8/3) the backlash can be continuously adjusted in a range up to some mechanical degrees. The backlash is fixed by cap screws (Fig. 8/5) that exert a compressive load on the parts of the coupling. Taper pins (Fig. 8/7) can be inserted in plain holes (Fig. 8/6) in order to get a precise and reliable setting of the backlash in discrete steps 2= 0\u00b0; 0,05\u00b0; 0,1\u00b0, 0,2\u00b0; 0,5\u00b0;1\u00b0; 2\u00b0; 50;10';15'. Additional plain holes (Fig. 8/8) were designed for a precise fitting of the coupling parts with no backlash at all. In the following figures the measurement results for 2g= 2, 2 =1\u00b0 and for 2 =0,50 are displayed. On the left hand side the signals nM(t) and iq(t) are depicted and on the right hand side there is a zoom, pointing out the interesting area of nM(t). The presented measurement results point out that t2 can be detected clearly in any case. As mentioned above, uncertainties in finding t1 exactly do not cause serious problems, because the area between the signals QM(t) and QL(t) for t t1 is negligible" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002065_ecc.2007.7068339-Figure7-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002065_ecc.2007.7068339-Figure7-1.png", + "caption": "Fig. 7. A conceptual illustration of the used ship simulator with its rotational axis.", + "texts": [ + " 6 shows the phase portrait of the antenna tracking output error angles \u03d5e and \u03b8e with employed NIMC subject to sea motions. The objective of this simulation is to show that both errors go to zero from randomly chosen initial conditions. Sinusoidal waves are used to simulate the pitch and roll disturbances due to the wave effect standards according to [7]. The idea of having a real test of STA resulted in taking a practical simulation of the antenna operation on a virtual ship. Ship simulator, Fig. 7, can simulate the movement of the ship in different conditions. This system provides a reliable simulation environment when heave, surge, and sway forces are negligible specially because they result in translative motions which are not taking into account by considering the long distance between the ship and the satellite. In the verification tests, the maximum amplitude and frequency of the pitch, roll, and yaw disturbances, according to [7], has been set to the ship simulator. Fig. 8 shows the low-pass filtered output of the roll and pitch gyros" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002608_iembs.2008.4649685-Figure9-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002608_iembs.2008.4649685-Figure9-1.png", + "caption": "Fig. 9. A high-speed camera system.", + "texts": [ + " This tendency may come from the accumulation of the blood mass at the fingertip. Fig.8 shows the change of the accumulated blood mass with respect to time where the horizontal and the vertical axes denote the time and the blood mass, respectively. As expected, the accumulated blood mass increases at the fingertip under the pressed condition. In order to observe the skin surface deformation under the pressed condition compared with that under the nonpressed condition, we set up the high-speed camera system [11] as shown in Fig.9. Fig.10 shows the result of the skin surface deformation where Fig.10 (a) and (b) are under the non-pressed and the pressed conditions, respectively. The \u201cBefore\u201d and \u201cAfter\u201d in Fig.10 denote the surface profiles before and after the force impartment, respectively. The finger deformation during the force impartment is obtained by chasing the slit laser by the high-speed camera. The result provides us with the rich information on the deformation under both conditions, compared with the point-typed stiffness sensor" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001788_j.compstruct.2006.02.029-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001788_j.compstruct.2006.02.029-Figure2-1.png", + "caption": "Fig. 2. Levels of relative energy density. Vertical arrows for external loads and horizontal arrows for reactions at symmetry boundary. Non-displaced boundaries as well as the displaced boundaries (parallel), i.e., separated by the pure translation e0 = 52.78 lm.", + "texts": [ + " It may be criticized that the problem is solved inversely, assuming the contact size a and determining the total contact force F. However computer times are a few seconds and parameter studies are therefore easily performed. The matrix [K] = ([Sd] 1 + [Sp] 1) 1 is a stiffness matrix, which is positive definite and strictly diagonal dominant. From this follows for each row (or column) in [K] thatPn j\u00bc1Kij > 0 for all i, and therefore {Au} > {0}, and we can always find an indentation e0. Thus the contact problem has a solution with only positive contact pressure. Fig. 2 shows for illustration a finite element result from the case of DR = 100 lm and a = 1.0181 mm. As the von Mises stress has no meaning in relation to non-isotropic materials, the levels of relative energy density are presented. The highest value is (for all the cases of Table 1) found in the interior of the orthotropic disc. Illustrations for the variation along the boundaries of the models are presented in Section 5. Fig. 2 shows in the middle, the initial positions of the boundaries with only point contact at x,y = 0,0, i.e., before contact deformations. Returning to the results in Table 1 we found the total flexibilities md + mp to vary in the range 3.4\u20134.0 (10 11 Pa 1). The mean value \u00f0md \u00fe mp\u00deestimated \u00bc 3:7 10 11 Pa 1 \u00f013\u00de we use for comparisons, and to validate this value we may also argue from a more direct point of view. In the model behind Table 1 the pin is assumed to be made of aluminum with modulus of elasticity E = 7 \u00b7 1010 Pa and Poisson\u2019s ratio m = 0", + " The cylindrical length (thickness) t = 3.5 mm, the pin radius R = 2 mm, the constitutive matrix components for the pin as specified in (2), and the constitutive matrix components for the orthotropic disc as specified in (3) were unchanged. We now rotate the orthotropic disc 90 getting the flexible orthotropic disc direction in the direction of the external force F, and thus change (3) to Cxxxx \u00bc 7:0 1010 Pa Cyyyy \u00bc 0:27 Cxxxx Cxxyy \u00bc Cyyxx \u00bc 0:38 Cxxxx Cxyxy \u00bc 0:40 Cxxxx \u00f014\u00de The specific result shown in Fig. 2 is for the case of clearance DR = 0.1 mm and contact size a = 1.0181 mm. We choose the same values with the rotated orthotropic disc, and get the result shown in Fig. 8, with the resulting force being F = 1256.3 N and the resulting indentation being e0 = 82.19 lm. Comparing the results in Fig. 8 with the results in Fig. 2 we note \u2022 The contact forces has decreased (totally from 1916 N to 1256 N). \u2022 The displacements at the orthotropic disc contact surface has increased (the indentation e0 from 53 lm to 82 lm). \u2022 The \u2018\u2018reactions\u2019\u2019, i.e., the stresses at the model symmetry line has increased in the orthotropic disc and decreased in the pin (see lengths of arrows). \u2022 The energy distribution is accordingly also changed in magnitude as well as in distribution, resulting in an increase in maximum with a factor of 1.6 in the orthotropic disc" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002597_acc.2008.4586601-Figure12-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002597_acc.2008.4586601-Figure12-1.png", + "caption": "Fig. 12. Two-piece snake.", + "texts": [ + " 10 the creature goes backwards initially when opening its flappers and moves forward when they are closed. The net effect is a forward motion due to the differential friction. Previous discussions can be extended to other robotic devices. One could offset two flappers to obtain a tortoise Fig. 11. In this case the flapper angles \u03b81 and \u03b82 have opposite phases to accommodate locomotion. In the next section we will extend the flapper to a simplified snake model. We extend the previous analysis to a simplified model of a snake, Fig. 12. We will derive the friction forces exerted by the environment on the snake. First, consider a small piece with length dr of the snake from either the upper or lower bar, which is located at r(t) with respect to a inertial reference frame with basis {ex, ey}, Fig. 13. According to the differential friction model (1) the friction experienced by this differential slab is FBx (r, \u03b2) = \u2212\u00b5a(\u3008r\u0307, Bx\u3009)\u3008r\u0307, Bx\u3009Bx FBy (r, \u03b2) = \u2212\u00b5T \u3008r\u0307, By\u3009By, (11) where \u3008\u00b7, \u00b7\u3009 denotes the inner product, Bx = [ cos(\u03b2) sin(\u03b2) ]\u2032 is the axial direction and By = [ \u2212 sin(\u03b2) cos(\u03b2) ]\u2032 is the transversal direction, both with respect to the inertial reference frame. The total x and y components of the friction force are Fx(r, \u03b2) = \u3008FBx , ex\u3009 + \u3008FBy , ex\u3009, and Fy(r, \u03b2) = \u3008FBx , ey\u3009 + \u3008FBy , ey\u3009, respectively. The virtual work due to virtual displacements in the x and y direction of the differential piece at r is \u03b4Wr = Fx(r, \u03b2)\u03b4x + Fy(r, \u03b2)\u03b4y (12) We now proceed to derive the friction forces acting on the entire snake. Let rh = (x, y) denote the position of the hinge of the snake, Fig. 12. We further assume that both links have unit length. A point r \u2208 [0, 1] units away from the hinge on the upper and lower bar is ru = rh + rBu x , (13) rl = rh \u2212 rBl x, (14) where with respect to the inertial reference frame, Bu x = [ cos(\u03b8 + \u03b3), sin(\u03b8 + \u03b3) ]\u2032 and Bl x = [ cos(\u03b3 \u2212 \u03b8), sin(\u03b3 \u2212 \u03b8) ]\u2032 , respectively. We will use the Lagrangian dynamics approach to find the generalized friction forces. Recall the Euler-Lagrange equations d dt \u2202L \u2202q\u0307 \u2212 \u2202L \u2202q = Q, (15) where L is the Lagrangian and Q contains the external and control forces. Let the generalized coordinates be q = [x, y, \u03b8, \u03b3], Fig. 12. From (13) the virtual displacements at a distance r from the hinge on the upper bar, \u03b4ru = (\u03b4xu r , \u03b4yu r ), due to the virtual displacements in the general- ized coordinates, are \u03b4xu r = \u03b4x \u2212 r sin(\u03b8 + \u03b3)(\u03b4\u03b8 + \u03b4\u03b3), (16) \u03b4yu r = \u03b4y + r cos(\u03b8 + \u03b3)(\u03b4\u03b8 + \u03b4\u03b3) Similarly from (14), the virtual displacements on the lower bar are \u03b4xl r = \u03b4x + r sin(\u03b3 \u2212 \u03b8)(\u03b4\u03b3 \u2212 \u03b4\u03b8), (17) \u03b4yl r = \u03b4y \u2212 r cos(\u03b3 \u2212 \u03b8)(\u03b4\u03b3 \u2212 \u03b4\u03b8) Substituting (13), (14), (16), (17) into (12), the friction forces on both the upper and lower bar r units away from the hinge rh may be expressed in the form \u03b4Wr = Tx(r, x\u0307, y\u0307, q)\u03b4x + Ty(r, x\u0307, y\u0307, q)\u03b4y + +T\u03b8(r, x\u0307, y\u0307, q)\u03b4\u03b8 + T\u03b3(r, x\u0307, y\u0307, q)\u03b4\u03b3" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003293_s0263574710000421-Figure4-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003293_s0263574710000421-Figure4-1.png", + "caption": "Fig. 4. The 5R manipulator escaping from a singularity.", + "texts": [ + " (23) is rewritten in a matrix\u2013vector form as Js \u23a1 \u23a3 1\u03c92 2\u03c93 3\u03c94 \u23a4 \u23a6 = \u2212 0\u03c91 0$1 \u2212 4\u03c95 4$5, (24) where Js = [1$2, 2$3, 3$4]. In order to solve (24) it is necessary that det(Js) = 0. In other words, the closed chain is at a singular configuration if the screws 1$2, 2$3, and 3$4 are linearly independent, which implies that dim(Js) < 3, which occurs mainly when the revolute joints of such screws are aligned; see Fig. 3. In what follows, it is shown how the closed chain can escape from a singularity by means of a simple case. To this end, consider the singular configuration depicted in Fig. 4. In order to escape from the singularity, the following steps are suggested: (1) Detect the singularity. Since det(Js) = 0, the closed chain is at a singular configuration. (2) Lock the revolute joint q1, e.g. q\u03071 = 0, and unlock q2 such that this revolute joint becomes a passive element. (3) Consider the third limb, containing q3, as an active leg which implies that point C \u2032 can move along a circular trajectory. (4) Since none of the closed chains are at a singular configuration (dealing with the inverse velocity analysis), the motor q3 can be actuated producing a circular trajectory over point C \u2032 eliminating the undesirable alignment of the revolute joints responsible for causing the singularity of the original closed chain" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002034_2007-01-2860-Figure5-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002034_2007-01-2860-Figure5-1.png", + "caption": "Figure 5: Asymmetrical vertical movement", + "texts": [ + " These norms establish vehicle reference axes, as well as the degrees-of- freedom, which are listed as follows, and presented in Figure 3: - yaw angle: Z-axis rotation angle; - pitch angle: Y-axis rotation angle; - roll angle: X-axis rotation angle; For kinematic validation of the suspension, two basic movements are studied: first the symmetrical vertical movement of the wheels, caused by pitch for instance, as illustrated in Figure 4. The kinematic behavior influenced by the body roll, as in a curve maneuver, is also studied, generating an asymmetrical vertical movement of the wheels, as illustrated in Figure 5. 3 The total dimension of both symmetrical and asymmetrical movements depends on project parameters such as unsprung mass, in-roll mass transference, spring stiffness and end-restraints characteristics. Detailed investigation of all these basic characteristics during vehicle conception phase is essential. Camber and toe parameters directly interfere with vehicle behavior in maneuvers. A brief description is made about this influence in vehicle handling behavior. Camber (\u03b5w) is the angle formed between the vertical line of each wheel and its respective center axis, in the vehicle front view" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003545_00032710802677175-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003545_00032710802677175-Figure2-1.png", + "caption": "Figure 2. Schematic of the alginate microsphere preparation.", + "texts": [ + " The enzyme-mediated hydrolysis of urea leads to a shift in the pH of the assay medium, which was then monitored using absorbance values at the two wavelengths described earlier. A calibration curve was then constructed as shown in Fig. 1. Alginate Microsphere Preparation Alginate microspheres were prepared according to a technique described earlier (Jayant and Srivastava 2007). Briefly, 10ml of solution, comprising 2% alginate, was mixed with 2mg=ml urease enzyme solution. The mixture was then extruded using the VAR J 30 droplet-generator equipment (Fig. 2) into a vessel containing 250mM calcium chloride solution for external gelation under continuous stirring. The hardened alginate microspheres were then separated by centrifugation (1000 rpm for 1 min). Scanning electron microscopy (SEM) was then performed on 794 M. Swati and R. Srivastava the prepared alginate microspheres using a Hitachi S3400N instrument as shown in Fig. 3. Particle-size analysis was then performed using a Cuvette Helos (CUV-50ML=US) instrument as shown in Fig. 4. Cresol Red Incorporation and Polyelectrolyte Coating of Alginate Microspheres Following urease immobilization, alternating nanofilm coatings of polyelectrolytes were deposited on alginate microspheres", + " An approximate linear response was observed for urea concentrations in the range of 0.1 to 6.7mM, whereas a higher sensitivity was observed for urea in the concentration of 0.01 to 0.1mM (shown in inset). This is an important observation because urea concentrations in the dialysate are expected to be at least 10 times lower than in the physiological fluid, thereby warranting increased sensitivity at lower ranges. 798 M. Swati and R. Srivastava Alginate microspheres were then prepared according to the protocol described earlier as shown in Fig. 2. Urease enzyme was encapsulated in the alginate precursor solution before being extruded in the gelling medium. The encapsulation efficiency was determined by estimating total protein concentration using Bradford\u2019s reagent and was calculated to be approximately 93%. SEM images of urease-loaded alginate microspheres are shown in Fig. 3. The microspheres display a round morphology with sizes in the range of 50\u201380mm. Particle-size analysis on the prepared alginate microspheres yielded similar results with more than 90% of the microspheres, with sizes between 50 and 80mm, as shown in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003740_10402000903491283-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003740_10402000903491283-Figure1-1.png", + "caption": "Fig. 1\u2014Tripod sliding universal joint assembly: 1, input shaft; 2, sleeve; 3, slide rod; 4, joint bearing; 5, tripod arm; 6, tripod.", + "texts": [ + " Serveto, et al. (12) recently presented the theoretical and experimental studies on the axial force generated by the tripod joint. With the development of the automobile industry, however, the tripod sliding universal joint (TSUJ) has been advanced recently for transmitting torque larger than previously (Wang, et al. (13) and Wang and Chang (14)). It generally consists of the input shaft with three metal sleeves, three slide rods, three joint bearings, and the tripod with the output shaft, shown in Fig. 1. Compared with the traditional tripod joint, its input shaft and tripod link up through three slide rods instead of three spherical or cylindrical rollers. However, the premature failures induced by the lubricating insufficiency of the main mating surfaces restrain the popularization and utilization thereof. Especially, severe failures of the mating surfaces between the sleeves and slide rods were frequently found in that they are the main elements transmitting torque. In order to improve the lubricating performance between them, therefore, the slide rod was redesigned into a novel one with several annular bumps around a column (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001188_03321640610649023-Figure5-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001188_03321640610649023-Figure5-1.png", + "caption": "Figure 5. Magnetic-bearing model: (a) FE mesh; (b) magnetic flux lines (unbiased excitation) and (c) magnetic flux lines (biased excitation)", + "texts": [ + " The magnetic flux density (Br, Bu) with respect to the standstill polar coordinate system can be gathered into a complex-valued field: B \u00bc Br \u00fej Bu \u00bc l[L X2jl r 2al r r rt l e2jlu: \u00f031\u00de Similarly, the force components Fx and Fy, expressed by the Maxwell stress tensor, can be brought together: F \u00bc Fx \u00fej Fy \u00bc lz Z 2p 0 n0 2 B2e jur du: \u00f032\u00de Introducing equation (31) into equation (32) and working out the integral leads to: F \u00bc 2 4plzn0 rrt l[L X l\u00f01 2 l\u00deala12l: \u00f033\u00de The 2D FE model of a magnetic bearing (Figure 5(a)) is equipped with the eccentric air-gap element. Transient simulations are carried out to test the numerical behavior of the eccentric air-gap element and the embedded force computation. The rotor is submitted to a prescribed movement which is a combination of a translation from the position (0.5 mm, 9p/8) to the position (0.5 mm, p/8) and a rotation around its axis. The stator coils are excited such that a force under an angle of 22.58 is generated. Both unbiased and biased current excitations are considered (S\u030ctumberger et al., 2000). The magnetic fluxes when the rotor is at the center position are shown for both unbiased and biased excitation in Figure 5. The force experienced by the rotor depends on the position of the rotor (Figure 6). Eccentric air-gap element 353 D ow nl oa de d by A th ab as ca U ni ve rs ity A t 0 8: 56 2 3 Ju ne 2 01 6 (P T ) During the transient simulation, the stator and rotor FE meshes do not have to be re-constructed. The movement of the rotor between two successive time steps only affects the operators T 1; G 1 and Ra, which are embedded in the eccentric air-gap stiffness operator Kag. In practice, only the parameters specified in the routine (27) have to be adapted" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003146_978-3-642-16584-9_14-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003146_978-3-642-16584-9_14-Figure1-1.png", + "caption": "Fig. 1. Giant Magnetostrictive Actuator. (a) Section. (b) Prototype.", + "texts": [ + " Compared to conventional magnetostrictive and piezoelectric materials, it has higher strain value, more rapid response speed, better frequency characteristics and higher reliability. Furthermore, it is not subject to fatigue, overheating and failure. GMM is widely used in the areas of aviation, aerospace, and precise manufacturing, etc. One of the most important applications is to construct actuators which are often called Giant Magnetostrictive Actuators (GMA). A type of giant magnetostrictive actuator produced by Department of Materials Science and Engineering of Beihang University is depicted in Fig.1. However, hysteresis nonlinearity is the inherent property of GMA, it can reduce the control accuracy and even lead to oscillation. There are mainly three classes of approaches for the modeling of hysteresis. One is the physical modeling theory, which is represented by Jiles-Atherton model [1] and Duhem model [2]. The other is the operator modeling theory, which is represented by Preisach model [3] Krasnoselskii-Pokrovskii (KP) model[4], Prandtl-Ishlinskii (PI) model [5], etc. Another is the intelligent modeling theory based on computational intelligence [6]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001538_20060906-3-it-2910.00121-Figure3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001538_20060906-3-it-2910.00121-Figure3-1.png", + "caption": "Fig. 3 Definition of contact angle", + "texts": [ + " The two hips are connected to the body via a differential which has an angle \u03c1 on the left side and \u03c1\u2212 on the right side. On a flat surface \u03c1 is zero but becomes non-zero when one side moves up or down with respect to the other side. The differential joint \u03c1 is passive (unactuated) and provides for the compliance with the terrain. The wheels are steerable with steering angles denoted by i\u03c8 . The wheel terrain contact angle i\u03b4 is the angle between the z-axes of the i-th wheel axle frame Ai and contact coordinate frame ic as shown in Fig. 3. In order to derive the kinematics equations, we must assign coordinates frames. Fig. 4 illustrates our choice of coordinate frames for the left side of the rover. The right side is assigned similar frames. In Fig. 4, R is the rover reference frame whose origin is located on the center of gravity of the rover, its x-axis along the rover straight line forward motion, its y-axis across the rover body and its zaxis represents the up and down motion. The differential frame D has a vertical (along z-axis) offset denoted by 1k and a horizontal distance of 2k from R. The distance from the differential to the hip, denoted by 3k , is half the width of the rover. We now introduce three more frames, all of which have origin at the wheel axle. The length of the legs from the hip to the wheel axle is 4k . The hip frames 41 H,,H L for the four wheels are obtained from the differential frame by rotation and translation as shown with the Denavit-Hartenberg (D-H) parameters dhdhdh ad ,,\u03b3 and dh\u03b1 in Table 1 and in Fig 3. Similarly the steering frames 41 ,, SS L and axle frames 41 A,,A L are defined in Table 1 and Fig 3. We must now use the basic frame to frame equations (1)-(2) and go through the frames sequentially from wheel i terrain contact ic , wheel axle iA , steering iS , hip iH , differential D, and finally to the rover reference R. Equation (1)-(2) for the contact to the axle becomes ( )( ) ( )Ticici,AiAi T cicici,AiAi 00R r00uRu \u03b4\u03d5\u03d5 \u03d5 &&& &&& \u2212+= \u00d7+= (5) where the rotation matrix is \u239f \u239f \u239f \u23a0 \u239e \u239c \u239c \u239c \u239d \u239b \u2212 = cici cici ci,Ai c0s 010 s0c R \u03b4\u03b4 \u03b4\u03b4 , as evident from Fig. 3. Next we form wheel i axle to steering velocity propagation as ( )( ) ( )TiAiAiSiSi T AiAiAiSiSi R uRu \u03c8\u03d5\u03d5 \u03d5 &&& &&& \u2212+= \u00d7+= 00 000 , , (6) where \u239f \u239f \u239f \u23a0 \u239e \u239c \u239c \u239c \u239d \u239b \u2212 = 100 0 0 , ii ii AiSi cs sc R \u03c8\u03c8 \u03c8\u03c8 . The next in the chain is the hip frame, and we can write ( )( ) ( )TiiSiSiHiHi T SiSiSiHiHi hR uRu \u03c3\u03d5\u03d5 \u03d5 &&& &&& \u2212+= \u00d7+= 00 000 , , (7) with \u239f \u239f \u239f \u23a0 \u239e \u239c \u239c \u239c \u239d \u239b \u2212 \u2212 = 010 )(0)( )(0)( , iiii iiii SiHi hshc hchs R \u03c3\u03c3 \u03c3\u03c3 , 14 \u03c3\u03c3 = , 23 \u03c3=\u03c3 , and 4,3 2,1 1 1 = = \u23a9 \u23a8 \u23a7 \u2212 = i i hi . The differential frame velocities are obtained from (1)-(2) and Table 1 as ( )( ) ( )TiiiiSiHi,DiDi T 3i4HiHiHi,DiDi bh00R kb0kuRu \u03c1\u03c3\u03d5\u03d5 \u03d5 &&&& &&& ++= \u2212\u00d7+= (8) where \u239f \u239f \u239f \u23a0 \u239e \u239c \u239c \u239c \u239d \u239b ++ +\u2212+ = 100 0)()( 0)()( , iiiiiiii iiiiiiii HiDi bhsbhc bhcbhs R \u03c1\u03c3\u03c1\u03c3 \u03c1\u03c3\u03c1\u03c3 , \u03c1\u03c1\u03c1 == 14 , \u03c1\u03c1\u03c1 \u2212== 32 , and 3,2 4,1 1 1 = = \u23a9 \u23a8 \u23a7\u2212 = i i bi " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000176_amc.2000.862847-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000176_amc.2000.862847-Figure1-1.png", + "caption": "Fig. 1: The image of power-assistant platform", + "texts": [ + "his paper considers a power-assistant platform that is operated comfortably independently of environmental conditions. Power-assistant platform is a mobile robot with force sensors, and a schematic view of the robot is shown in Fig.1. Here an operator moves the mobile robot by pushing and pulling a handle bar. The handle bar has two force sensors to detect the applied forces by the operator. Then the translational and rotational forces are also calculated, and are utilized to determine a motion controller of the driving wheels assisting the movement of the mobile robot. In general, to carry a baggage by a normal platform, the operator is required to consider the following conditions: 1. control of moving speed, 2. configuration of the platform, 3", + " 8 = Jac0-'(Z -i- J;,,b) (3) Where 8 = [Or B 1 I T , i = [v wIT. 2.2 Dynamics In case the center of gravity of the robot is at P, as shown in Fig.2, the motion equation of the robot is described as follows. (4) Where M : mass of the platform J , : inertia of each wheel J : inertia of the platform around z, Substituting eq.(3) into eq.(4), the equivalent mass matrix in the robot coordinates is written as follows. = [F ik] As described before, the power-assistant platform is a mobile robot with force sensors as shown in Fig.1. A robot operator moves the mobile robot by pushing and pulling a handle bar. The handle bar has two force sensors, and the applied force is detected. Here the detected force of the left and right force sensor is resolved into the translational and the rotational force by using eq.(7), and the assist motion is determined in the robot coordinates. (7) F h = F,h+F; N h = r , x F,h + r l x F! Y r Fig.2: The top view of the power-assistant platform Fig.3: Generation of the force to the robot In the proposed approach, a compliance controller based on F h and N h is introduced" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003878_gt2010-22105-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003878_gt2010-22105-Figure1-1.png", + "caption": "Fig. 1 FE model of the fan bladed disk", + "texts": [ + " In the identification process, the FE model which makes the of every blade of the real mistuned bladed disk equal to zero is defined as the \u201cideal\u201d tuned FE model avgp avgp avgp [12]. Finally, according to different situations, we introduce appropriate constraint conditions to Eq. (13), adjust Eq. (18)-(20), and assemble and solve the linear equations using the linear constrained least square method to obtain the identification vector . \u0398 To validate the effectiveness of the presented mistuning identification method, an industrial fan bladed disk is presented as the numerical example. The bladed disk has 13 blades and its FE model is shown in Fig. 1. There are 2,167 elements and 6,400 nodes in a single sector. According to the identification theory introduced in Section 2, the input parameters include the natural frequency i\u03bb and the corresponding mode shape i\u03c6 of the tuned model, and the damping parameters (\u03b1 and \u03b2 ), the steady-state response amplitude and excitation frequency ju j\u03c9 of the mistuned model. All these parameters can be obtained by definition or computation of two benchmark models, one for the tuned bladed disk and the other for the mistuned (represents the real mistuned bladed disk)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003308_s1064230710040143-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003308_s1064230710040143-Figure2-1.png", + "caption": "Fig. 2. Schematic diagram of haptic interaction mechanism.", + "texts": [ + " On the other hand, the obtained results are inter esting because of the fact that they give the idea of the limiting possible improvement of operation of the control system due to reducing the time interval of cal culations and the justified character of efforts for reducing this interval. It was already mentioned in Introduction that the multidimensional multilink mechanism of manipula tor type is used as the interaction mechanism; remov able models held or moved by human hands can be fixed at the free end of this mechanism (Fig. 2). The manipulator is equipped by the controllable drives, usually DC motors. The current state of output shafts of the motors is characterized by the coordinates d1, d2, \u2026, dn, which can be measured by special sensors together with their velocities and accelerations . The drives provide measurement of joint coordinates of the manipulator , , \u2026, using reduction gears. The connection between joint coordinates and drive coordinates is represented as (1.1) where = ( , , \u2026, ) and d = (d1, d2, \u2026, dn), and g = (g1, g2, \u2026, gn) = P(d) are the n dimensional vectors of joint coordinates after reduction gears and coordi nates of output shafts of the drives and these coordi nates reduced to joints; P(d) is the vector function of reduction transformation; and \u03c4 = (\u03c41, \u03c42, \u2026, \u03c4n) is the vector of elastic deformation of reduction gears" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002437_0954406jmes197-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002437_0954406jmes197-Figure1-1.png", + "caption": "Fig. 1 A rotor\u2013bearing\u2013seal system", + "texts": [ + " The key point, the influence of non-linear oil-film forces and non-linear seal forces is put in the analysis of this paper. As a result, torsional vibration of rotor and gyroscopic effects may be neglected and only transverse vibration of rotor needs considering. Thus rotor\u2013bearing\u2013seal system can bemodelled as a Jeffcot rotor system, inwhich the rotor is simplified to one disc with two transverse stiffness and the masses of the shaft is equivalent to the disc and two bearings. The model of rotor\u2013bearing\u2013seal system is shown in Fig. 1. The mathematical model of the rotor\u2013bearing\u2013seal system takes into account four degrees of freedom\u2013 horizontal and vertical displacements of the rotor at the disc location (X2,Y2) and at the journal (X1,Y1), correspondingly. Then the dynamic equations of system are established as follows\u23a7\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23a8 \u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23a9 m1X\u03081 + c1X\u03071 + K 2 (X1 \u2212 X2) = fX1 m1Y\u03081 + c1Y\u03071 + K 2 (Y1 \u2212 Y2) = fY1 \u2212 m1g m2X\u03082 + c2X\u03072 + K (X2 \u2212 X1) = FX2 +m2r\u03c92 cos \u03c9t m2Y\u03082 + c2Y\u03072 + K (Y2 \u2212 Y1) = FY2 +m2r\u03c92 sin\u03c9t \u2212 m2g (1) Considering the non-linear oil-film force model under assumption of short bearing [7], a dynamic model of the non-linear oil-film force is established, for the model has better accuracy and convergence" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002510_dscc2008-2176-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002510_dscc2008-2176-Figure1-1.png", + "caption": "Figure 1. Standard notation and sign convention for ship motion description (SNAME, 1950) .", + "texts": [], + "surrounding_texts": [ + "Consider the following generic mathematical model that captures the ship motion characteristics in the horizontal plane: m(u\u0307\u2212 vr) = Xh + \u03c4x , m(v\u0307 +ur) = Yh +Y\u03b4\u03b4 , Izr\u0307 = Nh +N\u03b4\u03b4 (1) where (\u00b7)h are the hydrodynamic forces and \u03c4x and \u03b4 are the propeller thrust and the rudder deflection. m and Iz are the vehicle mass and mass moment of inertia. u,v are the body-fixed linear velocities (surge and sway), and r is the yaw rate. The hydrodynamic forces are often modeled as a nonlinear function of surge velocity, sway velocity, and yaw rate, for example, as shown in [8]- [11]. A Serret-Frenet formulation can be used to represent the vessel kinematics in terms of path parameters, which allows for convenient definition of cross track and course keeping error. The following equations, first introduced in [12], cast the path following error dynamics in the Serret-Frenet framework: e\u0307 = usin(\u03c8\u0303)+ vcos(\u03c8\u0303), \u02d9\u0303\u03c8 = \u2212\u03ba 1\u2212e\u03ba (ucos(\u03c8\u0303)\u2212 vsin(\u03c8\u0303))+ r (2) where e (defined as the distance between the center of gravity and the path, see Fig. 2 for the geometric interpretation), \u03c8\u0303 = \u03c8\u2212\u03c8s are referred to as the cross-track error and relative heading error, respectively. \u03c8S is the path direction as shown in Fig. 2. \u03ba is the curvature of the given path. The control objective of the path following problem is to drive e and \u03c8\u0303 to zero. For surface vessels operating in the open sea, the path is often a straight line or way-point path, which consists of piecewise straight lines. In these cases, the curvature \u03ba is zero, therefore the heading error dynamics could be simplified into: \u02d9\u0303\u03c8 = r. (3) Furthermore, in this study, we maintain a constant propeller speed rather than a constant surge speed, which is more realistic for most ship maneuvering operating conditions. Hence, the surge velocity is assumed to be bounded, but time varying throughout the paper." + ] + }, + { + "image_filename": "designv11_69_0000801_j.jmatprotec.2004.07.013-Figure3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000801_j.jmatprotec.2004.07.013-Figure3-1.png", + "caption": "Fig. 3. Specific power loss and flux density distributions under sine wave excitation at core back flux density of 1.3 T.", + "texts": [ + " The coefficient of resisance of each thermistor was measured over the temperature ange 15\u201320 \u25e6C before being placed on the test lamination. he error for localised flux and loss measurements were esimated to be within \u00b15 and \u00b18% of measured value. . Results and discussion Localised flux and loss distribution under sine and PWM oltage excitation in locations behind an arbitrary tooth and lot were found using the measurement system described bove. The tangential and radial flux density variation with ime at locations behind a slot and a tooth under sine and WM voltage excitation were measured at peak core back ux density of 1.3 T. Fig. 3a, b and Fig. 4a, b illustrate flux ensity and power loss at positions S1\u2013S4 under sine wave nd PWM excitations, respectively. The modulation index nd the switching frequency remained constant at 0.7 and .5 kHz, respectively. The flux density distribution under PWM voltage excitaion is similar to that under sinusoidal excitation behind both lots and teeth as shown in Figs. 3a and 4a. The peak value f resultant flux density in both S and T locations decreased towards the outer region of the stator core" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001445_20050703-6-cz-1902.01954-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001445_20050703-6-cz-1902.01954-Figure1-1.png", + "caption": "Fig 1. Tanker Reference Frames", + "texts": [ + " This model has been used in previous studies by a number of other investigators (Kallstrom et al., 1979; Fossen, 1994; McGookin et al., 2000). This model can be represented in the following standard state space form: (1) )u,x(fx = Assuming all the states are available, here x represents the system state vector and u is the input vector to the tanker. The dynamics (defined about the body-fixed inertial reference frame) and the kinematic states (defined about the earth-fixed inertial reference frame) for this model are defined in Fig 1 and Table 1 (Fossen, 1994) shown below: From the table above, it should be noted that water depth, h, influences the model dynamics. In this investigation, it is assumed that the vessel is travelling in deep water (with water depth 200m) so that water depth effects have negligible influence on the yawing motion and surge velocity of the vessel. It can be noted that there is a distinction between commanded inputs and the actual state vector of rudder deflection, \u03b4r, and propeller speed, n. This imposes a maximum actuator rate of = 2" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000910_detc2004-57046-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000910_detc2004-57046-Figure1-1.png", + "caption": "Figure 1. SYSTEM SET-UP", + "texts": [ + " Our aim is to provide the surgeon with an operation environment very similar to manual instrumental surgery (i.e. the surgeon can always feel forces exerted on the instruments). According to [15], the influence of force feedback on operation time seems to be even more important than it is for visual feedback. Following this analysis of deficiencies, we developed an open evaluation platform for robotic surgery that was tailored to the needs of sensitive force feedback for delicate operations like bypass operations in cardiac surgery (Fig. 1). Our workstation is not a telemanipulator that is controlled by visual servoing of the surgeon. Instead, it can be directly controlled by transmitting 6 DOF coordinates to its control unit. This is an important feature for closing control loops in machine learning applications, which can be applied in order to autonomously perform certain recurrent tasks, e.g. automated cutting or knot-tying. Similar to other systems, our setup comprises an operatorside master console for in- and output and a patient-side robotic manipulator that directly interacts with the operating environment. As shown in Fig. 1, our system has two manipulators, which are controlled by two input devices. Each of the two arms of our surgical robot is composed of the following subsystems. A low-payload robot bears a surgical instrument that is deployed with the surgical workstation daVinci (TM). We have developed a special adapter that interconnects the robot\u2019s flange with the instrument. The surgical instruments have three degrees of freedom. A micro-gripper at the distal end of the shaft can be rotated and adaptation of pitch and yaw angles is possible", + " Since the rotation of the robot\u2019s flange and the rotation of the instrument share one axis, the combination of robot and instrument results in a manipulator with eight degrees of freedom. That means our system is a redundant manipulator. This can be exploited to evaluate different kinematical behaviors. The most important one is trocar kinematics. This allows 6 dof control of the end effector, while the shaft of the instrument has to be moved about a fixed fulcrum (keyhole surgery). Position and orientation of the manipulators are controlled by two PHANToM devices (Fig. 1). This device is available in different versions with different capabilities. Our version provides a full 6 dof input, while force feedback is restricted to three translational directions. The user controls a stylus pen that is equipped with a switch that can be used to open and close the micro-grippers. The most interesting feature of the PHANToM devices we used, is their capability of providing the user with haptic feedback. Forces are feeded back by small servo motors incorporated in the device" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001480_sme-200067068-Figure6-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001480_sme-200067068-Figure6-1.png", + "caption": "Figure 6. The experiment set-up of vibration of the spindle/disks assembly (a) For measurement and postprocessing (b) Closer view of the shaker with head expander (c) For rotating disks at different rotational speed.", + "texts": [ + " The parameters of the disks and the stationary part are listed in Table 1. Some natural frequencies of the disks and the stationary part are calculated using finite element analysis software ANSYS and listed in Table 2. In order to test the program, we first studied the natural modes of the disks/spindle system without the top cover. The natural modes of the case can be calculated using Eq. (43) after setting the stationary part as a fixed body. The natural modes of the system are measured and performed by using a shaker. The experiment setup is shown in Fig. 6. The HDD with the top cover removed was fixed on the shaker through a head expander. A sweeping sinusoidal signal was produced by a Nicolet dynamic signal generator and analyzer and applied to the shaker to excite the base of the HDD. The velocity of the top disk and the velocity of the base were measured using a Polytec LDV and B&K accelerometer, respectively. The transmissibility ratio of disk velocity to base velocity was obtained. A HC5802 Motor Driver together with a power supply was used to spin the disks at different speeds" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002781_wcica.2008.4592987-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002781_wcica.2008.4592987-Figure1-1.png", + "caption": "Fig. 1 A six-wheeled mobile robot with a rocker-bogie configuration.", + "texts": [ + " This paper is organized as follows: Section 2 presents the definition of coordinate frames, wheel-terrain contact model and kinematics model of an all-terrain mobile robot with passively compliant mechanism. Section 3 describes odometry and slip calculation process. In Section 4, several physical experiments and their results calculated by field data are shown and discussed. We present our conclusions in Section5. II. KINEMATICS MODELING 978-1-4244-2114-5/08/$25.00 \u00a9 2008 IEEE. 581 Consider a six-wheeled mobile robot with a rocker-bogie configuration, six wheels dependently driven, and four wheels dependently steered, showed as in Figure 1. In the following context the kinematical equations of the mobile robot will be established on two assumptions: (1) Both the wheels and the terrain on which the robot traverses are rigid; (2) Each driven wheel is modeled as a disk, so the wheel-terrain contact will be considered as single point contact. The coordinate frames are illustrated as in Figure 2. The front, middle, and rear wheel on the right are numbered as wheel 1, wheel 2, and wheel 3 respectively, while the front, middle, and rear wheel on the left are correspondingly numbered as wheel 4, wheel 5, and wheel 6" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002076_2008-01-1044-Figure10-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002076_2008-01-1044-Figure10-1.png", + "caption": "Figure 10: Cylinder bore deformation", + "texts": [ + "25 mm2 for the minor thrust side. Figure 7 shows the meshed piston geometry and Figure 8 shows the skirt profile of the piston. The engine was modeled in Ricardo Wave to obtain combustive pressure traces. Figure 9 shows the pressure trace at 1000 RPM, the engine speed used for the simulation results presented in this paper. The beginning of the intake stroke is at zero crank angle degrees. The cylinder bore deformation and temperature distributions are as of [15]. The cylinder bore deformation is shown in Figure 10. Table 1 summarizes the engine and piston geometrical and material properties. Figure 7: Piston mesh SAE Int. J. Engines | Volume 1 | Issue 1716 In the following sections a comparison is made between simulation results obtained via different piston dynamic modeling approaches. The model that considers piston transverse motion in the plane perpendicular to the crankshaft axis is referred to as 2D, and the one that considers motion in the planes perpendicular and parallel to the crankshaft axis is referred to as 3D", + " Also the noise in the predicted motion disappears (compared to the one predicted by the 3D model) since the second land deformation is assumed to be invariant to the hydrodynamic and contact pressures and the oil film between the land and the cylinder bore provides a lot of damping. Now, considering the wear on the second land the results are very different (Figure 21). Both models predict about the same order of magnitude of wear. However, the 2D/Land model predicts the wear to occur from about 0 deg. to 180 deg., whereas the 3D/Land model predicts the wear to occur around the land\u2019s circumference. The maximum wear occurs around 90 deg., and more wear is predicted above the minor thrust side than above the major thrust side. This agrees with cylinder bore deformation (Figure 10). COMPARING DRY AND FULLY FLOODED SECOND LAND CONDITIONS \u2013 The assumption whether the second land lubrication conditions are either dry or fully flooded affects piston motion and as a result wear prediction. SAE Int. J. Engines | Volume 1 | Issue 1 719 The motion in the Xp-Yp plane (Figure 22, Figure 23) is noticeably affected during the intake, compression and exhaust strokes. During the expansion stroke where the combustion pressure is very high, it dominates the motion; thus the effect of dry or fully flooded lubrication at the land is minimal" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001717_iecon.2004.1433448-Figure3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001717_iecon.2004.1433448-Figure3-1.png", + "caption": "Fig. 3. Flelation between current and flux with saturation", + "texts": [ + " Definition of coordinates and symbols Coordinates used in this paper are defined as shown in Fig1 and Fig.2. The ~y-0 coordinate is defined as the fixed coordinate. The d-Q coordinate is defined as the rotating coordinate. The 7-6 coordinate is defined as the estimated rotating coordinate. As shown in Fig.2, the largest value of rotor inductance is defined as L,,,, and the smallest value of rotor inductance is defined as Lmin. In this paper, the d-axis direction of the SynRM's model is selected as the 0~7603-873C-9~04/$2$20.00 02004 \\E\u20ac\u20ac kinds of inductance in Fig.3. One is a static inductance, Fig. 2. Coordinates of SynRMs. L , which represents a sIope from the origin to a drive same direction BS Lmgn because this d-axis direction is an appropriate one for position estimation [SI. Therefore, the relation between the d-axis inductance and the q-axis one is the same in a11 types of synchronous motors: & < to. The symbols used in this paper are as follows. (ud uslT voltages on the rotating coordinate, [ad iqlT currents on the rotating coordinate, [wa q j I T voltages on the fixed coordinate, [io ialT currents on the fixed coordinate, [ea eplr Extended EMF on the fixed coordinate, [U-, wslT voltages on the estimated rotating coordinate, [i7 islT currents on the estimated rotating coordinate, R stator resistance, KE back-EMF constant, Ld &axis static inductance, Lk d-axis dynamic inductance, 9-axis static inductance, i", + "(l) , with this model defined as a linear model. KE is zero in SynRMs, L d is equal to L, in SPMSMs. In case magnetic saturation phenomenon is generated, inductance variations are caused. We can still use the linear model in the case magnetic saturation is not so large, but we must use the mathematical model that considers magnetic saturation in the case magnetic saturation is generated to a large degree. This is because the difference between two kinds of inductance becomes clear [ lO][ l l ] . Fig.3 shows the relation between current and flux in the case of generating magnetic saturation. There are two point, and the other is a dynamic inductance, L', which represents the slope of a tangent line of a drive point. The difference between two inductances becomes large under large magnetic saturation; it is frequently generated in SynRMs [9] and permanent-magnet-assisted SynRMs [12]\\13]. In these motors, we must use the model that considers this relation. The mathematical model that considers magnetic saturation is derived in eq" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001862_indico.2004.1497760-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001862_indico.2004.1497760-Figure2-1.png", + "caption": "Fig. 2 Chientation of four actuators to obtain pitch motion.", + "texts": [ + " In MDUN INSTITUTE OF TECHNOLOGY, KHARAGPUR 721302, DECEMBER 20-22,2004 305 this stage the launch vehicle is above the atmospheric level and thrust vector control is used io control vehicle.motion along pitch,,yaw and roll channels. Digital Auto Pilot (DAF') takes the body error signals as inputs and generates the command for the actuators. There are four actuators, attached with two engines (nozzles), control the pitch, yaw and roll motions of the vehicle. Orientations of the four actuators have been shown in the Fig. 2. Each actuator is inclined with the pitch and yaw axes by 45'. The DAP along the pitch and yaw axes are nothing but controller with PID structure. But in case of the roll DAP takes the PD structure because we are not interested in tbe steady state error of roll motion of the vehicle and for a short period we can let the vehicle to roll about its axis. In this stage, attitude control is achieved by swiveling the movable nozzle. Considering rigid body dynamics, the basic dynamic equation of the vehicle along the pitch axis is given bY ' w Where T, thrust generated by the engines, is the equivalent distance of the engines from the cerrter of gravity, Sil is deflection of nozzle and is moment'of inertia" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003398_ijmr.2009.026577-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003398_ijmr.2009.026577-Figure1-1.png", + "caption": "Figure 1 Position of temperature measurement", + "texts": [ + " (2006) used a combination of numerical solutions, inverse approach algorithms and experimental tests using inserted thermocouples to determine the heat flux flowing into the tool through the rake face and the heat transfer coefficients between the tool and the environment. Boud (2007) used an inserted remote thermocouple to determine the influence of the bar diameter on the temperature while turning carbon steel bar using uncoated HSS inserts. For the work described in this paper the temperatures were measured at a remote location (Figure 1) close to the tool/workpiece interface and the chip. The inserted/remote thermocouple technique was used. This technique was preferred as it is inexpensive, easy to calibrate, has a quick response time and good repeatability during experiments. Mineral insulated, metal sheathed, type K thermocouples with a measurement range between \u2013200\u00b0C and 1200\u00b0C were used. The thermocouples are constructed from Nickel-Chromium and Nickel-Aluminum metals, welded at the tip to a 1 mm diameter stainless steel sheath (Sonnekus et al" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003303_j.ejc.2009.09.007-Figure6-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003303_j.ejc.2009.09.007-Figure6-1.png", + "caption": "Fig. 6. Not s-reversible.", + "texts": [ + " Orient the loops \u03b1 and \u03b2 in arbitrary way, and make (\u03b1, \u03b2) an oriented link. Let \u03b1\u2217 = f1(\u03b1), \u03b2\u2217 = f1(\u03b2). Then, (\u03b1\u2217, \u03b2\u2217) is a mirror image of the oriented link (\u03b1, \u03b2), and hence the linking number Lk(\u03b1\u2217, \u03b2\u2217) of (\u03b1\u2217, \u03b2\u2217) is equal to \u2212Lk(\u03b1, \u03b2). On the other hand, since no two faces go through each other in our origami-deformation, it follows that the oriented link (\u03b1\u2217, \u03b2\u2217) is isotopic to (\u03b1, \u03b2), and hence Lk(\u03b1, \u03b2) = Lk(f1(\u03b1), f1(\u03b2)) = Lk(\u03b1\u2217, \u03b2\u2217). Since Lk(\u03b1, \u03b2) 6= 0, this is a contradiction. Therefore,M is not s-reversible. Example 6. The polyhedral surface shown in Fig. 6 is not s-reversible, since its boundary forms a link with nonzero linking number. It is known (e.g., Conway and Gordon [4], Sachs [12]) that every spatial embedding of the complete graph K6 contains a pair of disjoint cycles (loops) that forms a link with odd linking number. Hence the next corollary follows. Corollary 1. If the complete graph K6 can be embedded on a polyhedral surface M, then M is not sreversible. The genus of a surfaceM with boundary is the genus of the closed surface obtained by capping off each of the boundary components ofM with a disk" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002979_978-3-540-70534-5_7-Figure8.7-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002979_978-3-540-70534-5_7-Figure8.7-1.png", + "caption": "Figure 8.7: EyeTrack robot and bottom view with sensors attached", + "texts": [ + " The sensor we selected for most of our projects has a resolution of 1\u00b0 and accuracy of 2\u00b0, and it can be used indoors: \u2022 Vector 2X [Precision Navigation 1998] This sensor provides control lines for reset, calibration, and mode selection, not all of which have to be used for all applications. The sensor sends data by using the same digital serial interface already described in Section 3.3. The sensor is available in a standard (see Figure 3.8) or gimbaled version that allows accurate measurements up to a banking angle of 15\u00b0. Gyroscope, Accelerometer, Inclinometer 59 3.8 Gyroscope, Accelerometer, Inclinometer Orientation sensors to determine a robot\u2019s orientation in 3D space are required for projects like tracked robots (Figure 8.7), balancing robots (Chapter 10), walking robots (Chapter 11), or autonomous planes (Chapter 12). A variety of sensors are available for this purpose (Figure 3.9), up to complex modules that can determine an object\u2019s orientation in all three axes. However, we will concentrate here on simpler sensors, most of them only capable of measuring a single dimension. Two or three sensors of the same model can be combined for measuring two or all three axes of orientation. Sensor categories are: \u2022 Accelerometer Measuring the acceleration along one axis \u2022 Analog Devices ADXL05 (single axis, analog output) \u2022 Analog Devices ADXL202 (dual axis, PWM output) \u2022 Gyroscope Measuring the rotational change of orientation about one axis \u2022 HiTec GY 130 Piezo Gyro (PWM input and output) \u2022 Inclinometer Measuring the absolute orientation angle about one axis \u2022 Seika N3 (analog output) \u2022 Seika N3d (PWM output) 3.8.1 Accelerometer All these simple sensors have a number of drawbacks and restrictions. Most of them cannot handle jitter very well, which frequently occurs in driving or especially walking robots. As a consequence, some software means have to be taken for signal filtering. A promising approach is to combine two different sensor types like a gyroscope and an inclinometer and perform sensor fusion in software (see Figure 8.7). Devices, measuring a single or two axes at once. Sensor output is either analog Sensors 60 3 or a PWM signal that needs to be measured and translated back into a binary value by the CPU\u2019s timing processing unit. The acceleration sensors we tested were quite sensitive to positional noise (for example servo jitter in walking robots). For this reason we used additional low-pass filters for the analog sensor output or digital filtering for the digital sensor output. 3.8.2 Gyroscope The gyroscope we selected from HiTec is just one representative of a product range from several manufacturers of gyroscopes available for model airplanes and helicopters", + " The electromagnet has to be switched on after detection and close in on a can, and has to be switched off when the robot has reached the collection area, which also requires on-board localization. Driving Robots 136 8 8.3 Tracked Robots A tracked mobile robot can be seen as a special case of a wheeled robot with differential drive. In fact, the only difference is the robot\u2019s better maneuverability in rough terrain and its higher friction in turns, due to its tracks and multiple points of contact with the surface. Figure 8.7 shows EyeTrack, a model snow truck that was modified into a mobile robot. As discussed in Section 8.2, a model car can be simply connected to an EyeBot controller by driving its speed controller and steering servo from the EyeBot instead of a remote control receiver. Normally, a tracked vehicle would have two driving motors, one for each track. In this particular model, however, because of cost reasons there is only a single driving motor plus a servo for steering, which brakes the left or right track. EyeTrack is equipped with a number of sensors required for navigating rough terrain. Most of the sensors are mounted on the bottom of the robot. In Figure 8.7, right, the following are visible: top: PSD sensor; middle (left to right): digital compass, braking servo, electronic speed controller; bottom: gyroscope. The sensors used on this robot are: \u2022 Digital color camera Like all our robots, EyeTrack is equipped with a camera. It is mounted in the \u201cdriver cabin\u201d and can be steered in all three axes by using three servos. This allows the camera to be kept stable when combined with the robot\u2019s orientation sensors shown below. The camera will actively stay locked on to a desired target, while the robot chassis is driving over the terrain" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001158_bfb0042228-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001158_bfb0042228-Figure1-1.png", + "caption": "Figure 1: Attainable set and approximating cones", + "texts": [ + " 3 M o t i v a t i o n a n d s t a t e m e n t o f t h e t h e o r e m Consider the single input system { ~ = u I \" ( ' ) l <- 1 C n) ~, -- =~ z(0) = 0. .~(t) = { x E R ' : I x l l < t , l x : < x , < l { t + x t ) s - I x : } (12 I while K --~ = R x R +, the closed upper half plane. Thus no nontrivial truncation of K s lles inside the attainable s e t / I ( 0 at any time t > 0. However, for every closed convex cone ~7 (with vertex at zero) such that ~7 \\ {0} C_ intK \"~ there are positive constants C, T > 0 such that ~ N B(0, Ct s) C_ ~(t) for all 0 < t < T, see also figure 1. This picture already contains the main idea, as we have the following T h e o r e m 3.1 If-IO in a closed convez cone (with vertez 0 E It\" such that ~7 \\ {0} C int~ 7-~ for some ra < oo, then there are constants C > O, T > 0 such that -X q n B{0, Ct \"~) C/~(t) for a l l O < t < T . Coro l l a ry 3.2 I f ~ ''g = R n then there are constants C > 0, T > 0 such that B(0, Ct \"~) C .4(t) /or all 0 < t < T. We conclude this section with the following remarks: HSlder continuous with exponent I / m in the direction o f ~ \"-~ (or in a full neighbourhood of zero, in the case of a full tangent cone as in the Corollary)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001520_app.23770-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001520_app.23770-Figure1-1.png", + "caption": "Figure 1 (a) Preparation of the two-layer substrate. (b) Frame to hold the stretched substrate.", + "texts": [ + " The silicone was the two-liquid type RTV silicone mentioned earlier. The conducting rubber was purchased from Kinugawa Rubber Industrial, Chiba, Japan, (No. S60; volume resistivity, 1 /cm) as a 0.5-mm thick sheet. A strip of the conducting rubber (75 15 mm2) was stuck on a Teflon tray using a double-faced adCorrespondence to: M. Watanabe (mwatana@giptc.shinshu-u. ac.jp). Contract grant sponsor: Research Foundation for the Electrotechnology of Chubu. Journal of Applied Polymer Science, Vol. 101, 2040\u20132044 (2006) \u00a9 2006 Wiley Periodicals, Inc. hesive tape [Fig. 1(a)]. A mixture (1.5 g) of the RTV silicone (KE-109-A/B 50/50 wt %) was poured into the tray, which was then placed in an oven controlled at 100\u00b0C for 1 h to vulcanize the silicone. The obtained two-layer substrate was removed from the Teflon tray. It was then cut into a rectangular strip 75 mm long and 15 mm wide. The thicknesses of the upper and lower layers were about 0.8 and 0.5 mm, respectively. Gold deposition onto each substrate was carried out as follows. A homemade plastic frame [Fig. 1(b)] held the above single- or two-layer substrate that was prestretched by 20%. A thin gold film ( 30 nm thick) was deposited onto the substrate by an ion-sputtering technique using an ion coater (IB-3; Eiko Engineering, Hitachinaka, Japan). The surface of the gold film was then observed using a light microscope (BH-2; Olympus, Tokyo, Japan) without relaxing the stretched substrate. I also observed the surface while heating it by applying a DC voltage (24 V) to both ends of the conducting rubber (i" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002698_iros.2007.4399277-Figure8-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002698_iros.2007.4399277-Figure8-1.png", + "caption": "Fig. 8. Inverse kinematics of avoidance of a cylindrical hole in a thick wall.", + "texts": [ + " First, k\u03c6 is defined so as that the plane, on which LM is located, intersects the hole. Second, M\u03c6 is defined so as that the LM penetrates the hole. There are two patterns of solutions; however, only the case in which JM+1 is located after the wall is the real solution. B.6. Avoidance of a cylindrical hole in a thick wall When the link LM is passing through a cylindrical hole in a thick wall, four DOF, which are two lateral movements and two rotational movements, are restricted by the hole (see Fig. 8). Therefore, four additional DOF are necessary. The problem is divided to two sub-problems as follows: 1) JM is located on the hole axis by rotating Ji and Jj, 2) LM is on the hole axis by rotating Jk and JM. Since these two sub-problems are not independent to each other, they are carried out alternatively and iteratively until the satisfactory solution is obtained. Namely, under the state of fixing other joints besides the four joints of Ji, Jj, Jk, and JM, this method is as follows: (1) (Jk, JM) is decided appropriately" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003860_ipec.2010.5543118-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003860_ipec.2010.5543118-Figure1-1.png", + "caption": "Fig. 1. The machine concept of the compound magnetomotive forces", + "texts": [ + " This concept is referred to here as compound magnetomotive forces. First, the proposed compound magnetomotive forces are described and theoretically analyzed using a Fourier series expansion. Second, the proposed pole pair combination for obtaining higher output torque with lower torque ripples than other pole combinations is described. Third, the results of a finite element analysis (FEA) and experiments are presented to verify the proposed design and theory. 2. PRINCIPLE OF A VARIABLE CHARACTERISTIC MOTOR Figure 1 shows the concept of compound magnetomotive forces. A permanent magnet synchronous motor (PMSM) usually generates torque by using unique fundamental MMFs. Harmonics are designed to be as low as possible because they generate iron loss and noise. In contrast, our proposed motor has two different MMFs, pth and sth, because of a special magnet arrangement that is achieved as follows. Ring magnets with different pole numbers are coaxially aligned with a phase difference. Magnets having different polarities can be disabled, leaving only the magnets that have the same polarity" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002065_ecc.2007.7068339-Figure3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002065_ecc.2007.7068339-Figure3-1.png", + "caption": "Fig. 3. Error angles.", + "texts": [ + " NIMC DESIGN In order to define the control problem of under-actuated ship mounted STA devoid of beam sensor, we should know about the operation of beam sensor and its measurement. Beam sensor measures the error angles between satellite signal directional vector xe and the antenna plate vector xp. The sensor outputs are two angles \u03d5e and \u03b8e. \u03b8e is the angle between xp and the projection of xe on xp\u00d7yp surface of F p. Also, \u03d5e is the angle between xe and the projection of xe on xp \u00d7 yp surface of F p as shown in Fig. 3. It\u2019s clear that we can obtain the unit vector of xe in F p by vp = \u23a1 \u23a3 cos(\u03d5e)cos(\u03b8e) cos(\u03d5e)sin(\u03b8e) sin(\u03d5e) \u23a4 \u23a6 . (11) The elements of the vector vp are functions of time because of the motions of the ship and antenna motors. Obviously, we want the angles \u03b8e and \u03d5e to be zero, since we want the antenna to track the satellite. This is equivalent to vp = [ 1 0 0 ]T . (12) On the other hand, we can translate the vector xe in F p by vp = Rpexe , (13) where Rpe is the rotation matrix from Fe to F p due to inputs and disturbances given by Rpe = Rp jR jbRbe, (14) where the orthogonal rotation matrices have the property of Rp j = RT jp, R jb = RT b j, and Rbe = RT eb" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003150_978-90-481-9689-0_4-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003150_978-90-481-9689-0_4-Figure1-1.png", + "caption": "Fig. 1 A RPR-2PRR parallel manipulators with (a = 1, b = 2, L2 = 2, L3 = 2, x = 1/2,y = 1, \u03b8 = 0.2). The actuated joint symbols were filled in gray.", + "texts": [ + " 2010 G. Moroz et al. anties that only true solutions are obtained. Then, we classify the parameter space of a family of RPR-2PRR manipulators according to the number of cuspidal configurations. It is shown that these manipulators have either 0 or 16 cuspidal configurations. The proposed method is based on the notion of discriminant varieties and cylindrical algebraic decomposition, and resorts to Gro\u0308bner bases for the solutions of systems of equations. A RPR-2PRR parallel manipulator is shown in Fig. 1. This manipulator was analyzed in [6]. It has 1 actuated prismatic joint \u03c11, and 2 passive prismatic joints \u03c12 and \u03c13. The two revolute joints centered in A2 and A3 are actuated while the ones centered in A1, B1, B2 and B3 are passive. The pose of the moving platform is described by the position coordinates (x,y) of B1 and by the orientation \u03b1 of the moving platform B1B2. The input variables (actuated joints values) are defined by \u03c11, \u03b82 and \u03b83. The points B1, B2 and B3 are aligned, a= (B1, B2), b= (B1, B3), L2 = (A2,B2) and L3 = (A3,B3)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002979_978-3-540-70534-5_7-Figure11.13-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002979_978-3-540-70534-5_7-Figure11.13-1.png", + "caption": "Figure 11.13: Dynamic walking sequence [Jungpakdee 2002]", + "texts": [ + " 181 The CAD designs following this approach and the finished robot are shown in Figure 11.12 [Jungpakdee 2002]. Each leg is driven by only one motor, while the mechanical arrangement lets the foot perform an ellipsoid curve for each motor revolution. The feet are only point contacts, so the robot has to keep moving continuously, in order to maintain dynamic balance. Only one motor is used for shifting a counterweight in the robot\u2019s torso sideways (the original drawing in Figure 11.12 specified two motors). Figure 11.13 shows the simulation of a dynamic walking sequence [Jungpakdee 2002]. Walking Robots 182 11 11.6 References BALTES. J., BR\u00c4UNL, T. HuroSot - Laws of the Game, FIRA 1st Humanoid Ro- bot Soccer Workshop (HuroSot), Daejeon Korea, Jan. 2002, pp. 43-68 (26) BOEING, A., BR\u00c4UNL, T. Evolving Splines: An alternative locomotion controller for a bipedal robot, Seventh International Conference on Control, Automation, Robotics and Vision, ICARV 2002, CD-ROM, Singapore, Dec. 2002, pp. 1-5 (5) BOEING, A., BR\u00c4UNL, T" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000853_rd.202.0109-Figure7-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000853_rd.202.0109-Figure7-1.png", + "caption": "Figure 7 Asymmetric drive condition: (a) Applied asym metric radial field gradient vs time t for pulsed drive. (b) Net radial force on bubble vs t. (C) Lattice rotation mechanism for asymmetric drive shown in (a). Bubble positions at equal time intervals are shown. The forward and reverse deflection angles differ because the corresponding net forces differ, providing a net displacement X\" per cycle. (d) Lattice-rotation mechanism for symmetric sinusoidal drive. The bubble responds to simulta neous uniform gradient and sinusoidal z fields. The effect of the gyrotropic force is greater when the bubble diameter is smaller, providing a net X ,,-displacementper cycle. Both radial and trans lational amplitudes are greatly exaggerated.", + "texts": [ + " This question is naturally answered in terms of the well-known gyrotropic force F, which causes the bubble-deflection effect [5,31]. It is given by where z is a unit vector normal to the film plane and V is the instantaneous velocity of the domain. Here S is the state, or winding. number of the domain wall. It is given by the number of complete rotations executed by the in-plane component of the magnetic vector within the MARCH 1976 BUBBLE LATTICE MOTIONS \u2022 A symmetric drive Here we consider that the drive coil carries a train of identical current pulses of one sign. producing a similar radial gradient dHz/dp as shown in Fig. 7(a). Since a steady component of radial motion is not possible, the net force PI' acting radially with respect to the coil axis, including the effect of interbubble interactions and re straining barriers, must have both signs, as indicated in Fig. 7 (b). Indeed, under simple assumptions the time average of PI' would vanish. If the pulse width is not equal to one-half of the cycle time, then a steady component of velocity V\", orthogonal to PI' arises from the velocity dependence of the bubble deflection angles I) arising from coercivity and other non linear effects. For velocities below the critical instability value V p = 24 A / hK 1 / 2 , 8 is given by the expression (for S = 1) [5J: This expression varies from 8 = 0 at V~ 0 to a maximum value at large V", + " Al though the V dependence of 8 has not been tested experi mentally for such a small value of S, the corresponding expression for large S is well established in hard bubbles (large X) for velocities below that required for Bloch line annihilation [33]. However Eq. (5) cannot be relied on at drives exceeding that required to reach V/I' In any case, it is clear that 8 does depend significantly on drive and that Eq. (6) represents its maximum value. This fact combines with the asymmetry in F p to provide a net dis placement per pulse orthogonal to PI' because of the dif ference in 8 values for the two signs of PI\" as indicated schematically in Fig. 7 (c). The sign of the gyrotropic force is such that in a deflec tion experiment the sign of S H b ia s . F X V is always posi tive. If the pulse duration in our lattice rotation experi ment is less than the time between pulses, then the average of IFpl is greater during the pulse than otherwise. Equation (5) shows that 181 is then also greater during the pulse, if V < VI' holds. Under this restriction, the condition F p > 0 would imply a right-hand screw relation of lattice rotation to H b ia s ' Actually lefi-hand lattice ro tation is observed in the weak-field-well experiment with the negative pulsed field configuration of Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000887_1.1897746-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000887_1.1897746-Figure1-1.png", + "caption": "Fig. 1 The Bennett linkage in sc", + "texts": [ + " The generator of a linkage\u2019s axode, however, is an instantaneous screw axis ISA , represented by a screw motor or vector which incorporates the pitch of the ISA. Although it is possible to extract the pitch value, thereby reducing the screw vector to a line vector, it might be that the latter is a more awkward expression than the former. Such is the situation with the Bennett linkage. Consequently, in preference to the established method, we adopt some of the procedures outlined by Parkin 5 in employing screw vectors directly in our analysis. 2 The Kinematic Tools The Bennett linkage is depicted skeletally in Fig. 1 and, just as in Refs. 2,4 , we choose its frame of reference to comprise link DA and the pin at A. Notation relating to the loop is sufficiently explicated by the diagram and we make use of Bennett\u2019s index p = a/s = b/s . We also employ the abbreviations s for sine and c for cosine. The axis of symmetry of the joint-screw velocities 4 is of great relevance here; its disposition relative to a pair of alternate axes is displayed in Fig. 2. We rely upon screw vector algebra in a dual format and represent a generic screw by Contributed by the Mechanisms and Design Committee for publication in the JOURNAL OF MECHANICAL DESIGN" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002706_sice.2007.4421503-Figure3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002706_sice.2007.4421503-Figure3-1.png", + "caption": "Fig. 3 master/slave robot model", + "texts": [ + " , m (27) The above problem reduces to a standard quadratic programming min g g(t)\u2032\u03a8g(t) \u2212 2xm(t \u2212 T )\u03a8g(t) (28) subject to [ Aineq \u2212Aineq ] g(t) \u2264 [ Bineq \u2212Bineq ] (29) where \u2016x\u20162 \u03a8 := x\u2032\u03a8x, \u03a8 = \u03a8\u2032 > 0, Aineq and Bineq are Aineq = \u23a1 \u23a2\u23a2\u23a2\u23a2\u23a3 L 0 \u00b7 \u00b7 \u00b7 0 HcG . . . . . . ... ... . . . . . . 0 Hc\u03a6m\u22121G \u00b7 \u00b7 \u00b7 HcG L \u23a4 \u23a5\u23a5\u23a5\u23a5\u23a6 Bineq = \u23a1 \u23a2\u23a2\u23a2\u23a3 Cmax Cmax ... Cmax \u23a4 \u23a5\u23a5\u23a5\u23a6\u2212 \u23a1 \u23a2\u23a2\u23a2\u23a3 Hc Hc\u03a6 ... Hc\u03a6m \u23a4 \u23a5\u23a5\u23a5\u23a6 . In this section, the performance of the bilateral teleoperation proposed in section 3 is verified. The simulations are compared with the bilateral teleoperation without the CG in section 2. Consider the single-degree of freedom master/slave robots shown in Fig. 3. The environment is assumed to be a damper system. The model parameters are defined as Mm = Ms = 0.1911 kgm2, Bm = Bs = 2.0173 Nms. The control parameters is designed KP = 2.5, KD = 5, m = 3, \u03a8 = 1. It is assumed that the constant time delay T = 0.5 s exist and the input of the slave robot is subject to the saturation Cmax \u2264 3 Nm. The simulation results are obtained by using MATLAB. Figs. 4-8 and Figs. 10-14 report the responses under the PD-type control law and the PD-type control law with the CG action, respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001205_pime_proc_1972_186_090_02-Figure13-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001205_pime_proc_1972_186_090_02-Figure13-1.png", + "caption": "Fig. 13. Section through first design of apex seal", + "texts": [ + " The circular hub seal is located in the casing, and seals by inward acting forces as it slides against the rotor hub sphere. The apex seals are located by grooves in the rotor and slide on the casings. All the seals are spring-loaded to ensure correct seating for starting but are mainly loaded by gas pressure in the conventional manner. The tip seal has not changed much during development and has been made from a piston ring grade of cast iron. Hub seals were originally cast iron but more recently have been made in En 31 steel. Apex seals have been tried in several materials including carbon and cast iron. Fig. 13 is a section through the apex seal of the first design. Radial support of the apex seal is on the rotor. The components are drawn in the extreme positions they may occupy during operation. This movement between extremes can arise from a combination of bearing clearances, different thermal expansions, load deflections, surface form and position errors and timing gear backlash. In the first design, Fig. 13a, when the gas pressure forces the rotor away from its casing leakage areas of up to 6 mm2 can be opened up at each apex seal. The third design is shown in a similar fashion in Fig. 14. The apex seal is now supported radially by sliding contact on the casing and is made smaller and lighter. This simplifies the corner problems between the tip and apex seals. The hub ring secondary sealing surface has been moved out of the locating groove so that the latter no longer provides a circumferential leakage path" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001979_50037-5-Figure7.69-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001979_50037-5-Figure7.69-1.png", + "caption": "Figure 7.69 Model after meshing process.", + "texts": [ + " Click OK afterward. The frame shown in Figure 7.66 appears. In the box [A] No. of element divisions type 4 this time and press [C] OK button. In the frame MeshTool (see Figure 7.67) pull down [A] Volumes in the option Mesh. Check [B] Hex/Wedge and [C] Sweep options. This is shown in Figure 7.67. Pressing [D] Sweep button brings another frame asking to pick volumes to be swept (see Figure 7.68). Pressing [A] Pick All button initiates meshing process. The model after meshing looks like the image in Figure 7.69. Pressing [E] Close button in MeshTool frame ends mesh generation stage. After meshing completed, it is usually necessary to smooth element edges in order to improve graphic display. It can be accomplished using PlotCtrls facility in the Utility Menu. From Utility Menu select PlotCtrls \u2192 Style \u2192 Size and Shape. The frame shown in Figure 7.70 appears. In the option [A] Facets/element edge select 2 facets/edge and click [B] OK button to implement the selection as shown in Figure 7.70. In solving the problem of contact between two elements, it is necessary to create contact pair" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000254_(sici)1098-111x(200007)15:7<657::aid-int6>3.0.co;2-p-Figure7-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000254_(sici)1098-111x(200007)15:7<657::aid-int6>3.0.co;2-p-Figure7-1.png", + "caption": "Figure 7. Side view of helicopter\u2019s axis system.", + "texts": [ + " \u02d9 \u017e /FF trim, FF x , FF trim, FF u , FF trim, FF\u02d9 q M d y d\u017d .d , FF e e , trim, FFe \u00a1 < <1 if x - 3\u02d9 < <0 if x y 17 - 3\u02d9~m shov x y 14\u02d9 y if 3 F x F 14\u02d9\u00a2 11 \u00a1 < <0 if x - 3\u02d9 < <1 if x y 17 - 3\u02d9~m sFF x y 3 if 3 F x F 14\u02d9\u00a2 11 \u00a8 2\u017d .where x, u , and d represent the forward acceleration ftrs , pitch angle\u00a8 e \u017d 2 . \u017d .acceleration radrs and longitudinal cyclic input rad , respectively. X represents the aerodynamic force along the \u2018\u2018X axis\u2019\u2019 and M represents the pitching moment about the \u2018\u2018Y axis.\u2019\u2019 Figure 7 shows the axis system of the helicopter with respect to the sideview. The aerodynamic parameters and corresponding trim values for the hover and forward flight are given in Table I. These constant values were calculated by a trim analysis program using physical parameters from a Xcell 300 helicopter in hover and forward flight. The state vector of the T \u02d9 Tw x w xhelicopter model is x x x x s x x u u . It is assumed that the\u02d9 \u00a81 2 3 4 output vector of the model is the same as the state vector. In order to perform sensitivity analysis the model can be transformed into the form X x s y g " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003453_00368790910953640-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003453_00368790910953640-Figure1-1.png", + "caption": "Figure 1 Journal bearing experimental setup", + "texts": [ + " Of these two points is taken as origin the point where the thickness of the oil film is greater, and ismeasured ant clockwise to plot the Sommerfeld pressure curve-after determining graphically the values of n from: cos um \u00bc 23n 2\u00fe n2 \u00f05\u00de and the value of khas the sameunits of dimensions as p, n is nondimensional. In this study, the effects of shaft surface profiles, which enhance the load carriage capacity of journal bearing on lubrication are examined under variable revolutions and working conditions. An experimental setup has been designed to examine the performance of journal bearing, in various shaft surface profiles, which affect the load carrying capacity of journal bearing. The technical drawing of experiment setup is shown in Figure 1 (Sinanog\u0306lu, 2006). Experimental set-up consists of three main units. There are the steel journal shaft, the motor shaft and the journal bearing: 1 The radial bearing consists of a clear perspex journal bearing mounted freely on a steel journal shaft (A). 2 The large diameter journal shaft is directly fixedonto amotor shaft (B).The speedof this shaft is controlled by the standard tecquipment control unit, which is mounted within and in front of themain framework.With this system, a speed range of between 500 and 3,000 rev/min can be obtained", + " The positive pressure difference values correspond to local bearing load capacity. MobilSHC-629 synthetic oil was used as lubricant. Kinematics viscosity is 18.3 cSt. (1008C). Hub diameter is 54.8mm, diameter of the bearing bush is 55mm, bearing width is 70mm and weight of bearing is 650 g. Initially, oil tank was fixed 735mm levels (oil supply head p0 \u00bc 735mm). The pressure, which is constant due to on axial direction, indicated by 1, 2, . . . , 5 tubes are placed along the bearing axis. At the experimental work (Figure 1), the masses (H) on the (J) shaft can be placed ondifferent positions. The pressure values on the journal bearing were measured from 3, 6-15 and 16th tubes. The tube number also indicates the angles, i.e. 3 (0 and 3608), 6 (308), 7 (608), 8 (908), 9 (1208), 10 (1508), 11 (1808), 12 (2108), 13 (2408), 14 (2708), 15 (3008) and 16 (3308). Pressure values from the tubes 3 (0 and 3608), 6 (308), 7 (608), 8 (908), 9 (1208), 10 (1508), 12 (2108), 13 (2408), 14 (2708), 15 (3008) and 16 (3308) are used to testing data for ANN", + " tanh function is as follows: y \u00bc f \u00f0x\u00de \u00bc tanh\u00f0x\u00de \u00f010\u00de Its derivative is: \u203ay \u203ax \u00bc 12 y2 \u00f011\u00de Linear function is taken for input layer. The linear function is: y \u00bc f \u00f0x\u00de \u00bc x \u00f012\u00de The structural and training parameters of the proposed network are given in Table I. Moreover, average root mean square (RMS) errors for used training algorithm is shown in Table II. The journal bearing, which has a weight of 650 g, was run unloaded (only bearing weight) and loadings of 200 g each on the front and back loading rods, respectively (Figure 1). The surface profile of the shaft with longitudinal profile is seen in Figure 2. Figure 3(a) (Case 1) and (b) (Case 2) shows the variations of pressure differences versus the angular position for Mobil SHC-629 lubricant and as the loads of 650 g (bearing load) and 650 \u00fe 400 g were applied, respectively. In the Figure 3(a), the measured pressure difference values on the 3, 6, 7, and 12-15 pressure tubes were positive since the narrower oil wedgewas between 0 and 908, and between 240 and 3308.Theother pressuredifferencesvaluesonthe tubes8-11 thus between 120 and 2108 were negative" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002781_wcica.2008.4592987-Figure6-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002781_wcica.2008.4592987-Figure6-1.png", + "caption": "Fig. 6 A six-wheeled all-terrain mobile robot developed by SIA.", + "texts": [ + " If the robot is equipped with visual odometer or other sensors used for observing the velocity of the robot body including x , y , z , x\u03c6 , y\u03c6 and z\u03c6 , the wheel slip velocity can be directly obtained as follows. Let \u03c1\u03b2 iii iz iy ix i V V V 21 EEDXV +++= = , then: iiiiiii r \u03b7\u03be\u03b8 NMMV ++= (10) Left multiply the transpose of Mi, Ni and Oi, respectively, we obtain: iiii r\u03b8\u03be \u2212= VM T (11) iii VN T=\u03b7 (12) Actually, the sinking velocity i\u03bc if involved can also be obtained as the following: iii VO T=\u03bc (13) To testify the availability of the method proposed in this paper, three physical experiments were performed on the allterrain mobile robot developed by Shenyang Institute of Automation (SIA) as shown in Figure 6: (a) The robot straightforwardly moved on a hard even floor; (b) The robot straightforwardly moved upward on a soft sand slope; (c) The robot straightforwardly moved downward on a soft sand slope. Conventional encoder-based method and model-based method proposed in this paper were respectively used to calculate the odometry of the robot. Table I compares the odometry results of the two methods calculated by using the field data. For experiment (a), the ground truth distance that the robot was planned to move was 401" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000291_ias.1995.530343-Figure7-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000291_ias.1995.530343-Figure7-1.png", + "caption": "Fig. 7: 3D material mesh built by extrusion in axial direction", + "texts": [ + " Due to symmetry, only one fourth of one end region has to be modelled. The 3D model consists of a material mesh and a set of coil meshes required for the currents. Both meshes are generated separately allowing a different extrusion direction for material mesh and coil meshes. The material mesh can be built by rotating a base plan similar to figure 1 around the center line of the shaft or by shifting a base plan similar to figure 6 in the axial direction. All outlines required in the material mesh have to be present in the base plane. Figure 7 shows the material mesh when the extrusion is performed in axial direction. From this figure it can be seen that a part of the iron core 11s modelled as well. 51 7 The stator end winding is not incorporated in the material mesh. The end winding is modelled as a set of current driven coils in air. This is feasible since current redistribution due to skin effect is negligible in the stranded stator end winding. In the end ring and the rotor bars, skin effect cannot be neglected. Therefore, they are incorporated in the material mesh" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002616_cca.2007.4389318-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002616_cca.2007.4389318-Figure1-1.png", + "caption": "Fig. 1. A subsea template with relevant frames", + "texts": [ + " The geographic reference frame, (n-frame), is chosen with xn, yn and zn axis directed towards the North, East and Downward normal to the earth\u2019s surface respectively and target site assigned as origin. The configuration in the n-frame is \u03b7 , [xn, yn, \u03c8n] T , where xn, yn describe the distance from the target location and \u03c8n denotes the rotation about the zn axis. The body fixed reference frame (b-frame) is a moving coordinate frame with the origin attached to the Center of Gravity and axes corresponding to the principle axis of inertia. The frames assigned are represented in Fig. 1 for a subsea template to be installed at a target location. The velocity is defined in the b-frame as v , [ub, vb, rb] T where u, v \u2208 R are components of the absolute velocity in the xb and yb directions, and rb \u2208 R describes the angular velocity about the zb axis. The vectors, \u03b7 and v, are related by the kinematic equation, \u03b7\u0307 = J(\u03b7)v (1) where J(\u03b7) , cos\u03c8n \u2212sin\u03c8n 0 sin\u03c8n cos\u03c8n 0 0 0 1 . (2) By taking into account the inertial generalized forces, hydrodynamic effects, gravity, buoyancy, and the thrusters, the nonlinear dynamic equations of motion for a subsea module undergoing installation can be expressed in the canonical form for robotics, Mv\u0307 + C(v)v +D(v)v + g(\u03b7) = \u03c4 (3) where M \u2208 \u211c3\u00d73 is the system inertia matrix, M = MRB+ MA, including completely known rigid body inertia MRB and MA includes uncertainties depending on the operating conditions; C(v) \u2208 \u211c3\u00d73 is the coriolis-centripetal matrix including added mass, given by C(v) = CRB(v) + CA(v); D(v) \u2208 \u211c3\u00d73 represents the damping matrix, g(\u03b7) \u2208 \u211c3 is the vector of gravitational/bouyancy forces and moments and \u03c4 \u2208 \u211c3 representing the vector of forces and moments acting on the payload" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001682_iembs.2005.1616229-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001682_iembs.2005.1616229-Figure1-1.png", + "caption": "Fig. 1. 3D TRUS guided and robotic assisted prostate brachytherapy system. (a) Prototype system; (b) Schematic diagram.", + "texts": [ + " is with the Imaging Research Laboratories, Robarts Research Institute, London, Ontario N6A 5K8, Canada (Phone: (519)663-3834; fax: (519)663-3900; e-mail: afenster@ imaging.robarts.ca). To achieve dynamic intraoperative prostate brachytherapy, we developed a 3D TRUS guided and robot assisted system with new software. In this paper, we describe the development of the system and the related algorithms, and report on the targeting accuracy and variability achievable with the system. Our prototype system (see Fig. 1) consists of a commercial robot, and a 3D TRUS imaging system including a B&K 2102 Hawk US system (B&K, Denmark) with a side-firing 7.5MHz TRUS transducer coupled to a rotational mover for 3D imaging [3]. The mover rotates the transducer about itsT 0-7803-8740-6/05/$20.00 \u00a92005 IEEE. 7429 long axis, while 2D US images are digitized and reconstructed into a 3D image while the images are acquired. A one-hole needle guide is attached to the robot arm, so that the position and orientation of the needle targeting can be changed as the robot moves", + " The automatic seed segmentation algorithm is composed of following five steps: 1) 3D needle segmentation to obtain the needle position when implanting the seeds; 2) reducing the search space by volume cropping along the detected needle, as the implanted seeds are close to the needle; 3) non-seed structure removal based on tri-bar model projection; 4) seed candidate recognition using 3D line segment detection; 5) localization of seed positions using a peak detection algorithm described in [5] to localize the center of the seeds. To assess the performance of the 3D TRUS guided and robotic assisted system, we used tissue-mimicking prostate phantoms made from agar and contained in a Plexiglas box. A hole in the side allowed insertion of the TRUS transducer into the phantom, simulating the rectum (Fig. 1). Each phantom contained of two rows of 0.8mm diameter stainless still beads. The bead configurations formed a 4 4 4 cm polyhedron to simulate the approximate size of a prostate and the robot was controlled to target these beads. The displacements between the preinsertion bead position and the needle tip after the needle has been inserted into the phantom were used to analyse the needle insertion accuracy. A 3D principal component analysis (PCA) was performed to analyze needle targeting accuracy" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001989_978-1-84800-239-5_34-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001989_978-1-84800-239-5_34-Figure1-1.png", + "caption": "Figure 1. a. Generation of a face-gear; b.c. Coordinate systems applied for generation of a face-gear", + "texts": [ + " In this paper, the mathematical model of loaded tooth contact analysis (LTCA) for helical face gears is established. This model can allow for the loads and can solve the real contact ratio, the loaded contact path and the loaded transmission errors et al. More importantly, the new model needs much less computation time than the contact method of FEM. In addition, a longitudinal modification method different from the one proposed by Litvin[3] is presented for improving the stability of contact pattern. Figure 1 shows the coordinate systems applied for the generation of the face gear surface. Sa is the global fixed system. System Ss and S2 are rigidly connected to the shaper and the face gear, respectively. Sp is the auxiliary coordinate system. m is the angle between the axes of rotation of the shaper and the face gear, Zs and Z2 respectively. s and 2 are the rotation angles of the shaper and face gear respectively, and 2= sNs/N2, where Ns and N2 are the tooth numbers of the shaper and face gear. L1 and L2 are the limit inner radius and the limit outer radius, as shown in Figure 1-a. The face-gear tooth surface is calculated as the envelope to the family of the shaper\u2019s surfaces. 2 2 s s, s 2 s s s ( ) 2 s s s ( , ) ( ) ( , ) ( , , ) 0 s s s s s s u l u l f u l Mr r n v (1) here, rs(us,ls) is the surface of the shaper, and it is a modified involute helical surface in this paper. us and ls are the shaper surface parameters. f2s(us,ls, s) is the meshing equation for the generation of the face gear. ( 2)s sv is the relative velocity between the shaper and the face-gear in system Ss" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002979_978-3-540-70534-5_7-Figure3.7-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002979_978-3-540-70534-5_7-Figure3.7-1.png", + "caption": "Figure 3.7: Sharp PSD sensor and sensor diagram (source: [Sharp 2006])", + "texts": [ + " Laser sensors Today, in many mobile robot systems, sonar sensors have been replaced by either infrared sensors or laser sensors. The current standard for mobile robots is laser sensors (for example Sick Auto Ident [Sick 2006]) that return an almost Sensors 56 3 perfect local 2D map from the viewpoint of the robot, or even a complete 3D distance map. Unfortunately, these sensors are still too large and heavy (and too expensive) for small mobile robot systems. This is why we concentrate on infrared distance sensors. Figure 3.7 shows the Sharp sensor GP2D02 [Sharp 2006] which is built in a similar way as described above. There are two variations of this sensor: \u2022 Sharp GP2D12 with analog output \u2022 Sharp GP2D02 with digital serial output The analog sensor simply returns a voltage level in relation to the measured distance (unfortunately not proportional, see Figure 3.7, right, and text below). The digital sensor has a digital serial interface. It transmits an 8bit measurement value bit-wise over a single line, triggered by a clock signal from the CPU as shown in Figure 3.2. In Figure 3.7, right, the relationship between digital sensor read-out (raw data) and actual distance information can be seen. From this diagram it is clear that the sensor does not return a value linear or proportional to the actual distance, so some post-processing of the raw sensor value is necessary. The simplest way of solving this problem is to use a lookup table which can be calibrated for each individual sensor. Since only 8 bits of data are returned, the lookup table will have the reasonable size of 256 entries" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000687_eej.4390950213-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000687_eej.4390950213-Figure1-1.png", + "caption": "Fig. 1. Attractive electromagnet and ground rail.", + "texts": [ + " However, the numerical solutions are not useful for understanding the physical signi5cances of the phenomena and clarifying the influence of important parameters on the eddy current effect, The present authors have succeeded i n deriving the analytical solutions for the electromagnetic equations of the air gap. The solutions are expressed in terms of Fourier series and they can be calculated more easily than the numerical solutions. The analytic eolutions derived in this paper also make it possible to analyze the effects of important parameters and to determine their optimum values. 2. The Electromagnetic Equations and Their Solutions Figure 1 shows the arrangement of attractive electromagnets and ground rails discussed in this paper. The theory developed in this paper holds equally for the flat ground rail. Although the exciting current varies in practice to keep the air gap length constant [3], we assume in this paper that both the exciting current and the air gap length are kept constant; problems of controlling the exciting current and air gap length are beyond the scope of this paper. To facilitate analysis, we consider an analytical model as shown in Fig", + " Therefore it is practically very useful to reduce the core width. For instance, i f we let 2a = 0.01 m in the above example, 1/01 = 0.13 m and therefore we can obtain a large attractive force and a small braking force even at 360 km/h. A similar effect is achieved (without reducing the total width 2 4 by putting insulated iron boards of 1 0 mm thickness one on top of the other. We need not use thin laminated iron sheets. 4. Calculation Formulas for Attractive Force and Braking Force The attractive force exerted on each electromagnet in Fig. 1 is given by where b(x, z) is given by Eq. (34). The attractive force for null velocity is given by Fao=2a L Bo2/po ( 40) From the orthogonality of cosine function, Eq. (39) is rewritten as where and Heme, F a ( 1 ) =-[ 8 1 ---(l-e-uJ))F.O 1 no alL (45) From Eqs. (41, (5) and (91, the braking force is expressed as ab b is = L ( b e + b i ) L dP0 a x The total braking force is therefore given by Equations (47) and (48) indicate that the braking force is greatly affected by g/L. For g/L = 10-2, the braking force is of the order of 10-2 of the attractive force for the null speed" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002238_1.2748775-Figure5-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002238_1.2748775-Figure5-1.png", + "caption": "Fig. 5 \u201ea\u2026 Physical model and \u201eb\u2026 mathematical model", + "texts": [ + " 129, AUGUST 2007 https://vibrationacoustics.asmedigitalcollection.asme.org on 12/07/2018 Terms o 5 Impact Between Tape Edge and Flange The tape/flange impact can be modeled as a forced, single degree of freedom, mass spring system with \u201cdry\u201d friction. Since the direction of the friction force is always opposite to the direction of motion, the friction force is a piecewise constant function with respect to time. Fundamental research on so-called \u201cimpact oscillators\u201d has been performed by several researchers 15\u201317 . Figure 5 a shows the physical system of a tape sliding over a roller while Fig. 5 b illustrates the mathematical model used to simulate the physical system. A point mass, representing the tape, is connected with a spring to a fixed base and slides over the surface that represents the roller, thereby creating a piecewise constant friction force Ff, opposite to the direction of the lateral tape motion dx /dt. The spring effective tape stiffness k introduces a force Fs. Furthermore the Transactions of the ASME f Use: http://www.asme.org/about-asme/terms-of-use m r o s h p fl s fl t w r w t e r f t h f s fl I t t B a w m f t T t o r t a J Downloaded From: ass m is actuated by a harmonic force Fa, representing the reel un-out that forces the tape to move up and down when coming ff the supply pack" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003411_s11012-010-9339-3-Figure6-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003411_s11012-010-9339-3-Figure6-1.png", + "caption": "Fig. 6 Example of two-link manipulator", + "texts": [ + " Optimal controls and state variables (generalized coordinates and velocities) have been numerically evaluated in time. This procedure is recursive between all adjacent points which embrace the starting point I , all points of the A and B type on the MVC and the end point F . The number of A and B points depends on a concrete task (23)\u2013(28). In that way, the problem is finally solved. The presented procedure can be illustrated by the minimum travel time problem of the two-degree-offreedom manipulator [5] having the geometry of a double pendulum (Fig. 6). The manipulator has two massless links with a mass of m = 1 kg attached to the end of each link. The length of each link is l = 1 m. Both links lie on the horizontal plane, and the joint of the first link is located at the origin of the plane coordinate frame. The contacts between the masses and the plane are assumed frictionless. Generalized coordinates are angles q1 and q2, where Q1 and Q2 are actuation torques. Point M is moving along a given straight line xM + yM \u2212 2l = 0. (46) The expression (28) takes the following form cosq1 + cos ( q1 + q2 ) + sinq1 + sin ( q1 + q2 ) \u2212 2 = 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002979_978-3-540-70534-5_7-Figure9.6-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002979_978-3-540-70534-5_7-Figure9.6-1.png", + "caption": "Figure 9.6: Mecanum principle, turning clockwise (seen from below)", + "texts": [ + " all four wheels being driven forward, we now have four vectors pointing forward that are added up and four vectors pointing sideways, two to the left and two to the right, that cancel each other out. Therefore, although the vehicle\u2019s chassis is subjected to additional perpendicular forces, the vehicle will simply drive straight forward. In Figure 9.5, right, assume wheels 1 and 4 are driven backward, and wheels 2 and 4 are driven forward. In this case, all forward/backward veloci- Omni-Directional Robots 150 9 ties cancel each other out, but the four vector components to the left add up and let the vehicle slide to the left. The third case is shown in Figure 9.6. No vector decomposition is necessary in this case to reveal the overall vehicle motion. It can be clearly seen that the robot motion will be a clockwise rotation about its center. Kinematics 151 The following list shows the basic motions, driving forward, driving sideways, and turning on the spot, with their corresponding wheel directions (see Figure 9.7). \u2022 Driving forward: all four wheels forward \u2022 Driving backward: all four wheels backward \u2022 Sliding left: 1, 4: backward; 2, 3: forward \u2022 Sliding right: 1, 4: forward; 2" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001938_ias.2006.256753-Figure26-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001938_ias.2006.256753-Figure26-1.png", + "caption": "Figure 26. Prototype axial flux machine (on the far left) at the test bench", + "texts": [ + " PERFORMANCE EVALUATION OF THE PM MACHINE The drive prototype is composed of a 500 W axial-flux PM machine, a low-cost full bridge IGBT inverter and a 16-bit fixed point DSP. The inverter switching frequency is set to 15 kHz and the dead time is 2.5 [ts. A dc machine is coupled to the motor shaft and it's armature is connected to a resistive load, so that at 100 rad/s the motor is operating at nominal load conditions. An incremental encoder has been mounted on the drive's shaft to allow a comparison between the estimated angle and a high-resolution measurement. The main motor characteristics are shown in Table 1 and Fig. 26 shows a photograph of the motor and load at the test bench. Various tests have been performed to evaluate the performance of the drive when the observer estimates are used as state feedback. In order to better explain improvements that can be achieved using this technique, results are compared with the zero-order Taylor algorithm [1], used for state feedback. The observer eigenvalues (at the upper speed limit) were set to 80 Hz, 8 Hz and 0.8 Hz. The various harmonic decoupling waveforms were calculated offline and implemented in stored 256 word look up tables" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003812_003-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003812_003-Figure2-1.png", + "caption": "Figure 2. Forces acting on a bicycle.", + "texts": [ + " By also assuming that a top athlete can reach a speed of 10 m s\u22121 (see [6]) we obtain a value of hmax = 102 2 \u00d7 10 + 1 = 6 m. This is close to the records achieved by the top athletes of this event. More accurate models do take into account that energy losses will in fact occur (see for example [7]). Many young people spend much of their free time cycling and this activity offers ample opportunities for illustrating basic dynamics principles. When travelling horizontally, the bike is under the influence of five forces (figure 2). In the vertical direction gravity and the normal reaction force (F) cancel each other out. In the horizontal direction there are two retarding forces, one due to the aerodynamic drag (which is commonly referred to as air resistance) and one due to the rolling friction experienced by the wheels. Finally a propulsive force due to the reaction of the tyres of the bike pushing back on the road also acts in the horizontal direction. Therefore, we can assume that if the bike is moving with a steady velocity, Fb = D + Frf (7) where Fb is the propulsive force, D is the aerodynamic drag and Frf is the rolling friction" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000154_mmcs.1999.778211-Figure6-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000154_mmcs.1999.778211-Figure6-1.png", + "caption": "Fig 6. Composite force vector", + "texts": [ + " The color marker is displayed on Hyper Class to show the direction of 3D pointer. To accomplish cooperative work, we have developed a method where a 3D object is moved or rotated by a composite force given by participants. A fraction of movement of the 3D object caused by an operator is translated to an element force vector at each side, and a consequent movement of the object is determined according to the composite force vector which is calculated by summing up all of the element force vectors applied to the object. Fig.6 shows the image of cooperative work. The element force vectors a and b is translated to a composite force vector c. ( 5 ) Intelligent coding An intelligent coding is used to reduce the amount of data transmitted over the network. This coding enables the use of Internet for the operation in Hyperclass in real time. In this system, the human image and object data used for synthesizing a Hyperworld is broadcast from the server to every client before a session starts, and only the movement information such as hand gesture or the object motion transmitted from a client to every other client via the server in an event-driven manner during the session" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003021_jae-2010-1280-Figure6-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003021_jae-2010-1280-Figure6-1.png", + "caption": "Fig. 6. The FE model of Coiled tube with Velcro strips.", + "texts": [], + "surrounding_texts": [ + "The initial tube of telescopically-folded configuration is divided into two connected finite volumes, exterior airbag and inside airbag, by a dummy partitioned membrane allowing for a full continuity of the gas flow. The telescopically folded tube model is shown in Fig. 7. The spacing of tube wall between exterior airbag and inside airbag is lower than 3 mm. The boundary condition is similar as the models of above two tubes. The volume of folded cavity in initial stage of the inflation is 0.0012 m3, nearly one thirds of the whole volume of the tube." + ] + }, + { + "image_filename": "designv11_69_0000295_robot.1997.614414-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000295_robot.1997.614414-Figure1-1.png", + "caption": "Figure 1: System description", + "texts": [ + " F is the resultant external force/moment vector at the object\u2019s center of mass, A is the object\u2019s generalized inertia matrix, constructed from its mass m and inertia tensor I , and B represents Coriolis and gravitational effects. A contact frame C; is attached to the object at the contact point with the ith manipulator. Its third axis zi is the local normal vector to the object\u2019s surface. The position and orientation of Ci with respect to the object\u2019s frame R, are represented by the constant 3 x 1 vector ri and the constant 3 x 3 rotation matrix Ai, respectively (figure 1). Pure forces are applied to the object through the n. contact points. These forces, expressed in their corresponding contact frames, are grouped into a 3n x 1 vector \u201cf which is related to F through the grasp matrix c w: F = \u201cW(C2) \u201cf (5) Note that W depends on the object\u2019s orientation R because F is expressed with respect to the base frame Rb. The object\u2019s internal loading comes from the component of \u201cf that belongs to the null space of \u2018W(C2). Assuming that the vectors of a base for this null space are grouped to form the columns of the matrix \u2018N(R), we have: where fl designates the generalized inverse" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002841_gt2008-51086-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002841_gt2008-51086-Figure2-1.png", + "caption": "Figure 2: Model of kinematics relationship", + "texts": [], + "surrounding_texts": [ + "In figure 1 one can to observe a waterfall diagram related to one of the no contact probes installed upon the pinion of gearbox during the FSFL condition. In the plot above synchronous speed of High Speed (86 Hz) and Low Speed shaft (50Hz) are clearly visible and sometimes the vibration at the first torsional resonance component (9Hz) appears. Several researches have been published in order to provide a better understanding of lateral and torsional interaction, and presently it has been widely demonstrated that torsional and lateral motion are coupled inside the Gear Box, due to the offset of the of the shaft centerline. Many authors have developed a mathematical formulation to predict critical frequencies taking into account torsional-lateral coupling finding a good matching with field data. A work published in 1980 by Wachel and Szenasi showed that proximity probes were used to measure lateral vibrations in the fourth stage compressor as well as the lateral vibrations in the gear box. The authors observed presented a predominant subsynchronous vibration and noticed that the frequency of these lateral vibrations coincided with the third torsional natural frequency of the compressor train. Moreover, in 1996 Viggiano and Schmied issued a paper describing the coupling of the torsional and lateral motion in the gear and formulated a model in according to the following scheme: 2 ownloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 01/30/2016 Te It assumes that the gear wheel and the pinion are rigid and the torsional vibration is affected only by the gear journal bearings, with the kinematics relationship, described in the following formula: 222111 yryr The journals of the other component of the train have no influence on the variability of this phenomenon. The force on the gear bearings are given by: y x kk kk y x dd dd F F yyyx xyxx yyyx xyxx y x The damping and stiffness parameters in the above equation are function of bearing load, transmitted gear load and bearing type. Thus it wouldn\u2019t be possible to find out an univocal transfer function that correlates torsional stress with radial vibration. As described above, the results are heavily influenced by the dynamic behavior of the gear box bearings; in fact, in case of instability, subsynchronous vibrations are not necessarily inducted by torque ripple excitation. Therefore, before attempting at one of their test campaigns, the authors of this paper have carried out an extensive stability analysis of each train analyzed to define the rotordynamic behaviors of gear box In order to deeply inquire into this phenomenon, a decision was taken to arrange some test campaigns on different train to measure lateral vibration, torsional stress and electrical parameters coming from electrical machine. The primary objectives of the torque measurement were: Identification of train\u2019s Torsional natural frequencies (data-matching with the analytical prediction) Copyright \u00a9 2008 by ASME rms of Use: http://www.asme.org/about-asme/terms-of-use D Determination of the Amplification factor corresponding to each critical frequency Verification of mechanical Torque ripple presence .STANDARD EXPERIMENTAL SETUP The standard experimental setup was arranged as it follows: - regarding lateral vibration the train was fully instrumented with normal no contact probe; additional velocimeters and axial probes on gas turbine were also available; - 3-phases Electrical parameters of the generator were acquired by means of dedicated probes connected inside the VFD control panel, corresponding to 3 voltages and 3 currents for Armature and 3 voltages and 3 currents for Field since this generator is a synchronous type. - torsional stress measurement systems; here below, a description of the instrumentation is offered: the strain gage torque meter technology is based on a 90\u00b0 standard full bridge rosette applied in the 45\u00b0 standard configuration for torque measurement. Inductive sensor telemetry with dedicated signal amplifier, built on flexible pcb, has been connected to transmit signals to the receiver located on statoric part. ownloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 01/30/2016 The data acquisition system was able to capture torsional stress data up to 86 kNm (maximum range) and the sensitivity of this instrument, coming from calibration, was 8.6kNm/V. Regarding angular deflection measure based on optical probe system, authors adopted this technology because the field environment didn\u2019t allow a machining of a tooth wheel. This one replaces the use of tooth wheels and eddy current sensors. The angular deflection measure technology is based on common angular vibration measurement chain (TK 17 demodulator). The principle of this measure is the frequency modulation: since torsional vibration is simply a cyclic variation of shaft speed, it produces a variation of carrier signal frequency generated by a tooth wheel, which can be a gear with N teeth or an optical tape with black and white serigraphy. In the first case, the sensor type involved is an eddy current proximity probe, while in the second one it is an optical probe alternatively reading a reflective/not reflective part of the installed optical tape. The type of optical tape chosen, is based on polyester support built by means of thermal transfer, after studying different characterization of several models Two optical probes for each reflective tape are involved in order to minimize lateral vibrations. TK15 boxes amplify the signal and the TK17 demodulates the carrier frequency and it produces an output proportional to the angular vibration. To obtain torsional stress data it is necessary to calculate the difference between the two signals installed on the respective ends of the coupling. 3 Copyright \u00a9 2008 by ASME Terms of Use: http://www.asme.org/about-asme/terms-of-use Figure 7: Instrumentation installed on the coupling All the above dynamic signals were continuously recorded with a bandwidth of 20 kHz, in addition to several static signals such as temperatures, pressures and so on, coming from HMI gas turbine control panel. Real-time analysis through a dedicated software platform provided Trend, RMS, THD, Torque THD, Torque Ripple, Power, and Power Factor for a better investigation of the problem. In details, a powerful waterfall diagram was used to provide root cause diagnostic capabilities: with this kind of plot users can monitor frequency harmonics propagation (Y axis), versus time (X axis), identifying amplitude through a specific color scale (fig.8). Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 01/30/2016 T EXPERIENCE #1 The following experience was developed in a power generation plant, where a Ms61B were coupled with a 4 pole Generator (40MW) by a Gear Box (double helical) supported by bearings with fixed profile (4-lobes). Unfortunately, due to a premature failure of the instrumentation, the test was not thoroughly developed. It has been however possible to identify a strict correlation between electrical grid THD and subsynchronous radial vibration @1\u00b0Torsional Natural Frequency as exactly predicted by analytical calculation: vibration completely disappears when the generator switches in Island mode (fig.9). Simultaneously, torsional stress has been identified at the same 9Hz frequency through the s/g torque meter installed on load coupling recognizing the chance to build a transfer function involving GB radial vibration and pulsating torque. The captured data show an alternating torque of 800 Nm (fig.10) in no load condition @ 9 Hz and being the vibration level of 2-3 m at the same frequency, a possible correlation of roughly 260\u00f7400 Nm/m has been outlined 4 Copyright \u00a9 2008 by ASME erms of Use: http://www.asme.org/about-asme/terms-of-use For this reason the authors decided to continue the investigation, extending this kind of measurement to other geared train configuration with electrical motor." + ] + }, + { + "image_filename": "designv11_69_0003813_ijmms.2010.036064-Figure4-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003813_ijmms.2010.036064-Figure4-1.png", + "caption": "Figure 4 Single schemes of triangular fingerprints for (a) a new tool (b) a BUE affected tool", + "texts": [ + " On the other hand, BUE can be withdrawn due to the continuous dragging of the chip. Then, it can carry on particles of the insert surface causing tool wear. As it has been said, these alterations of the geometrical ideal cutting conditions facilitate a decreasing of the surface finishing. According to Rubio et al. (2005), the material accumulation on the tool during the cutting process may affect directly to the edge and counter-edge position angles of the tool. This influence can be reflected in an effective decrease of the value of these angles, Figure 4. The height (p\u2019) of the triangular fingerprint developed onto the workpiece surface diminished, Figure 4. It is well-known that there is a direct relationship between the surface roughness \u2013 measured in terms of the arithmetical average roughness parameter, Ra \u2013 and the edge and counter-edge position angle through the height p\u2019. This leads to an effective decrease of the Ra value, as it was demonstrated by Rubio et al (2005) for dry turning process of aerospace aluminium alloys. In this way, even though the ideal cutting conditions are modified, the quality of the surface finishing may be apparently improved due to these causes" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003709_icinfa.2010.5512052-Figure6-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003709_icinfa.2010.5512052-Figure6-1.png", + "caption": "Fig. 6 Disc brake model with contact forces and inertial flywheel", + "texts": [ + " A total of 288 massless dummy parts were appended to the rotor and pads at circles of four different radiuses. Massless dummy parts can transmit forces and don\u2019t add degrees of freedom to the model system. The geometric shape of the dummy parts can be defined at will, such as cube and spherule. At each circle, 30 dummy part spherules were fixed onto the rotor and distributed evenly, and 6 dummy part planes were fixed to the inner surface of unilateral pad. Then contact forces were established between every dummy part plane and every dummy part spherule at the same circle, as shown in Fig. 6. The model shown also contains the inertial flywheel mentioned in section III. D. Pressing force was applied to backplate through our program. For example of maximum pressing force of 84000N, the sinusoidal IF function of the program is: if(time-0.2:84000 *SIN(0.5*time*90d), 84000 * SIN(0.5 * time * 90d), 84000). The force curve is shown in Fig.7. D. Inertial flywheel modeling and equivalent rotational inertia calculation of entire vehicle Relevant equivalent rotational inertia (including rotational and moving inertia of an entire vehicle) was applied onto rotor through inertial flywheel. Inertial flywheel was connected with rotor through screw bolt joint, as shown in Fig. 6, and its inertial value can be changed based on different vehicle type. The disc brake is fitted on King Long minibus (high 7~8m) for front wheel braking. The parameters used for calculating equivalent inertial are as follows: ma =10500kg, m0=7050kg, =0.07, s=0.1, R=0.449m, =0.5. Substitution of above values into (1) given in the Ref. [9] gives If=499kgm2. If =(ma+ m0)(1-s) R2/2 (1) Where If is equivalent rotational inertial of unilateral front brake, ma is maximum total mass of vehicle, is rotational mass transformation coefficient of vehicle, m0 is unladen mass of vehicle, s is sliding rate, is braking force distribution ratio of front wheels and rear wheels, R is rolling radius of wheel" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002870_s0001924000004802-Figure10-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002870_s0001924000004802-Figure10-1.png", + "caption": "Figure 10. Definition of the target lateral manoeuver in the y-z plane of the body frame.", + "texts": [ + " m (m/s) 119\u22c559 188\u22c568 112\u22c545 153\u22c529 128\u22c504 200\u22c553 143\u22c546 131\u22c521 150\u22c506 157 \u03c8m (deg) 17\u22c5636 30\u22c58 24\u22c5059 \u201329\u22c5951 \u201333\u22c5921 \u201329\u22c5693 \u201342\u22c5693 \u201333\u22c5195 \u201326\u22c5994 \u201317\u22c5138 \u03b8m (deg) 23\u22c5494 38\u22c5971 22\u22c5014 30\u22c5728 25\u22c5265 41\u22c5946 28\u22c5568 25\u22c5937 30\u22c5013 31\u22c5556 \u03d5a (deg) 127\u22c52 283\u22c52 183\u22c581 136\u22c516 28\u22c5469 160\u22c58 83\u22c5297 153\u22c521 107\u22c571 26\u22c5922 tgo,man(s) 6\u22c57996 5\u22c57585 9\u22c55913 8\u22c54281 6\u22c53441 9\u22c55381 6\u22c58227 8\u22c58922 5\u22c58699 7\u22c54291 A step target manouever(3) in the form of a lateral acceleration command with constant magnitude, random direction and random initiation time is utilised in the simulations. The positive direction of the target lateral acceleration is shown in Fig. 10. The pitch and yaw autopilot dynamics are chosen as the second-order time invariant. The detailed data, used in simulations, are listed in Table 6. The cost functions, corresponding to each design point, are evaluated based on the average performance obtained over ten randomly generated engagement scenarios. These scenarios are pre-generated because the same situations are needed to evaluate different design With the design problem and parameters completely defined, PCACS was executed several times" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001550_iros.2005.1545550-Figure3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001550_iros.2005.1545550-Figure3-1.png", + "caption": "Fig. 3. System configuration. Rotation and translation are realized by friction wheel mechanism. Gimbals mechanism drives the forceps around the incision hole to determine the direction of the insertion.", + "texts": [ + " As for the mechanical performance such as speed, torque, force, and accuracy, they depend on the intended organ, surgical procedure, and so on. Especially in the case of image-guided surgery, the resolution of the imager defines the accuracy required to the manipulator. In this study, we set above values assuming that cholecystectomy is conducted by this manipulator. To satisfy the abovementioned requirements, we adopted following two mechanisms; \u201dFriction wheel mechanism\u201d (FWM) realizes the rotation around the forceps shaft, and translation along the shaft (circled number 1 and 2 in Fig.3). Gimbals mechanism is used to determine the direction of the forceps (3 and 4 in Fig.3). We set this manipulator above the incision hole. This is because mechanisms and actuators should be mounted near the operating field so that they require less torque or force and miniaturization of them are realized [3]. 1) Friction wheel mechanism (FWM): a) Friction wheel: We used \u201dfriction wheel\u201d. It consists of three titled idle rollers and outer case (Fig. 4(a)). Among the three rollers, we insert the forceps, and rollers hold the shaft of forceps. When outer case rotates, rollers travel spirally on the shaft of forceps (Fig", + " As we reported in [12], it is not a problem because abdominal muscle of a patient under anesthesia gets relaxed, and manipulator will not damage the abdominal wall by driving the forceps. The rotational center is located above the trocar at the incision hole. We newly implemented DC servomotors (ENC-1858011000/3CH-SAP1/576, Chiba Precision Co.,Ltd, Japan) for actuation. This motor has the special controller unit that simplifies the building up of closed feedback loop. The motor for pitch motion was located at some distance from the incision hole, and linkage mechanism was added for transmission (circled number 3 in Fig.3). We intended to keep sterilization around incision hole by separating sterilized and nonsterilized part via linkage mechanism. Linkage mechanism also works as a mechanical stopper to limit the working range of pitch for safety. The new prototype is shown in Fig. 6. Weight is 1.7 [kg]. FWM was 62\u00d752\u00d7150[mm3], 0.6[kg], and gimbals mechanism was 135\u00d7165\u00d7300[mm3], 1.1[kg]. III. EVALUATION EXPERIMENTS We conducted mechanical performance evaluation of our new prototype. Torque of pitch, roll, and rotation, and force of translation were measured by 6-DOF strain gauge force/torque sensor (MINI sensor 8/40, BL Auto Tech, Japan)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001806_00207720601014180-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001806_00207720601014180-Figure1-1.png", + "caption": "Figure 1. Single link robot.", + "texts": [ + " Choosing b\u00f0 , ih\u00de \u00bc 2P 1 \u00f0 , ih\u00deL\u00f0 , ih\u00de and assuming that w(t) and v(ih) are sufficiently small, Theorem 1 will reduce to Theorem 1 given in Guillard (1997). Remark 5: In general, it is very difficult to find a global solution \u00f0 ~x, t\u00de which satisfies all the conditions given in Theorem 1. In Huang and Lin (1995), a numerical method has been proposed for solving Hamilton\u2013Jacobi equations locally, where \u00f0 ~x, t\u00de is of the form \u00f0 ~x, t\u00de \u00bc ~xTP ~x\u00fe Xn k\u00bc3 Pk ~x \u00bdk , where 4. Example A schematic representation of a single link robot is given in figure 1. The definition of variables associated with the figure is given as follows: \u00bcAngle \u00f0rad\u00de B\u00bc Viscous bearing friction coefficient Nm sec rad 1 D \u00bc Diameter of the payload (m) L \u00bc length of the link (m) M \u00bc Mass of the payload (kg) m \u00bc Mass of the link (kg) \u00bc Torque (Nm) J \u00bc Inertia of the actuator\u2019s rotor \u00f0kgm2\u00de For this system, the dynamics equations are given by : _x1\u00f0t\u00de \u00bc B Mt x1\u00f0t\u00de \u00fe Nt sin\u00f0x2\u00f0t\u00de\u00de Mt \u00fe 0:1f\u00f0t\u00dex1\u00f0t\u00de Mt \u00fe 0:1w\u00f0t\u00de Mt \u00fe u\u00f0t\u00de Mt _x2\u00f0t\u00de \u00bc x1\u00f0t\u00de, 8>>>>>>< >>>>>>: \u00f04:1\u00de where x1\u00f0t\u00de \u00bc _ \u00f0t\u00de, x2\u00f0t\u00de \u00bc \u00f0t\u00de, u\u00f0t\u00de \u00bc is the motor\u2019s torque, Mt \u00bc J\u00fe 1 3mL2 \u00fe 1 10M\u00f0L\u00fe 0:5D\u00de 2, Nt \u00bc mgL \u00feMgL, g represents the gravitational constant, w(t) is the bounded motor torque disturbance, and f(t) represents the uncertain part of the viscous bearing friction coefficient and is assumed to be jf\u00f0t\u00dej 1" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000779_j.crvi.2004.05.006-Figure3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000779_j.crvi.2004.05.006-Figure3-1.png", + "caption": "Fig. 3. If we consider a little element of the surface, where a ring passes, we see that the torque exerted by the ring for a given rotation d\u03b8 of the membrane is higher when the ring is more curved. This induces a behavior of the surface tension such that the surface is stiffer wherever the pattern of fibers or cells forms a smaller ring, or a drawing with a smaller radius of curvature. In this image, we have represented the bundle as a rope of circular cross-section; it could as well be a ring of cuboidal cells, or a ring of fibers.", + "texts": [ + " In this first case, the fibers run azimuthally. Our question is: can the corresponding equilibrium shape be reconstructed from simple mathematics, such as the one coming from the Wulff construction? Fibers embedded in a surface may have an intrinsic curvature \u03ba completely different from the curvature of the surface. For the azimuthal distribution of rings (Fig. 2), \u03ba = 1/r(\u03b8) = 1/R cos\u03b8 . Suppose, now, that we look locally at a small element of the interface, close to a ring, and suppose we wish to rotate this element of interface (Fig. 3). We see that the ring exerts a torque on the surface, which is all the greater as the ring is more curved. This curvature should not be confused with the curvature of the surface itself. In the Hookean limit (linear elasticity), and neglecting the other deformations, we may simply consider the elementary torque dC as proportional to the curvature \u03ba (Eq. (6) and Fig. 3): (6)dC = \u03b1 \u03ba d\u03b8 We see that this torque is bigger closer to the poles. This is very familiar: straw disks, hats or baskets are stiffer in the center than in the periphery. So, when rotating an element of surface away from the reference configuration, the fiber resists locally by dC = \u03b1 \u03ba d\u03b8 , and this should overcome the other mechanical contributions in the pole region. The same term arises for bundles of meridian fibers that converge towards a pole and are arranged in rings. This torque corresponds to a variation in \u2018free energy\u2019 due to an equivalent surface tension \u03b3 such that (Eq" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001039_bf02637105-Figure5-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001039_bf02637105-Figure5-1.png", + "caption": "Fig. 5. The one-layer sealing process: The use of low-temperature melting material to close an opening at arbitrary chosen pressure and ambient", + "texts": [ + " This technique is self-aligned and seals the device from the environment. It allows a great flexibility with respect to the choice of the sealed-in gas and pressure and it offers compatibility with low-cost high-throughput batch fabrication techniques. Also, as it can be used for sealing openings directly above the device, it allows for a smaller die area and thus reduced device dimensions and costs. The basic idea of the proposed technique is to deposit a material with low melting temperature on top of openings forming a cusp until they are almost closed (Fig. 5a). This material is then reflowed in a furnace with controlled atmosphere and pressure to close the final openings (Fig. 5b). Obviously, a one-layer sealing process as shown in Fig. 5 has the risk that the film relaxes during reflow, thereby removing the cusp and opening the etch channels 366 Fig. 7. Processing l-layer of 6 ~tm PECVD BPSG on Si trenches (ca. 6 ~tm wide and 10 ~tm deep) before and after reflow 367 Fig. 9. 0.5 ~tm deposited BPSG on 9 ].tm PECVD SiO2 on Si trenches (6 pm deep) before and after reflow Fig. 10. 1.5 lam evaporated A1 on 2.2 t~m PECVD SiO 2 on Si trenches (10 pm deep, ca. 3 ~tm wide) before and after flow sealing process, a non-conformal silicon dioxide (Si02) layer was deposited first" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000210_iros.1996.571043-Figure10-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000210_iros.1996.571043-Figure10-1.png", + "caption": "Figure 10: Reference point of mobile manipulator", + "texts": [ + " First, we select the reference point on the first joint axis of manipulator. Note that we assume the first joint to be revolute. We also assume that its axis is vertical, intersects the symmetry axis of mobile robot, and dose not intersect the wheel axis. Next, we fix the desired position of the reference point on an extended line from the origin to the contact point with a distance 1, from the contact point, where I, is the distance from the first joint to the arm tip on XOYO plane when the manipulability[l5] is maximum (see Figure 10). Then the desired trajectory of the reference point is obtained from that of the object. 7 Experiment 7.1 Experimental system In this section, experimental result of pushing with a mobile manipulator will be presented to show the effectiveness of the proposed method. The robot used for the experiment is a mobile robot with two independent driving wheels on which a horizontal paralleldrive two-degrees-of-freedom manipulator is mounted. Figure 11(a) shows an overview of the mobile manipulator, and Figure l l ( b ) shows its parameters" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003234_vppc.2009.5289623-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003234_vppc.2009.5289623-Figure2-1.png", + "caption": "Figure 2. Magnetic field simulation of EVT", + "texts": [ + " This paper researches on the change law of the inductance when EVT is working,by detecting different inductances based on different excitation currents and angles between two motors\u2019 flux,builds EVT model,and then realize a certain extent of the decoupling control. II. MODELING OF EVT A. Inductance Change Rules Of EVT The structure of EVT in shown in Fig.1. Previous research[2] shows that EVT can\u2019t be treated as a simple mechanical connection of two general induction motors. First do the research on EVT\u2019s magnetic field analysis, field distribution is showed as figure1-2 From Fig.2 it can be seen that the flux generated by inner and outer machine will penetrate through the interrotor and enter into the counterpart, thus the magnetic field will coupled together, and from the contrast with the three groups of field distribution in the figure we can see that EVT\u2019s inner field distribution condition changed a lot when the inner and outer machines\u2019 excitation currents and their flux linkage angle change, especially to the magnetic flux lines distributed between stator and inner rotor, inner rotor and the inner cage of interrotor, stator and the outer cage of interrotor" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000662_s00366-004-0269-3-Figure11-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000662_s00366-004-0269-3-Figure11-1.png", + "caption": "Fig. 11 The linkage dual to the truss of Fig. 4 and its transformation; a The dual linkage; b The transformed dual linkage", + "texts": [ + " On the other hand, linkages composed of higher order modular groups correspond to compound trusses in which all the analysis equations are to be solved simultaneously. Thus, in accordance with the results of the section Transferring the method to linkages, the transformed Willis method in trusses enables to replace a compound truss with a simple one using the proposed technique. To clarify this idea, the compound truss given in Fig. 10 is analysed using the transformed Willis method. The truss of Fig. 10 is compound, thus it is reasonable to attempt solving it using the proposed procedure. The linkage dual to this truss is shown in Fig. 11a. The linkage of Fig. 11a, which actually is the same as the one treated in the section \u2018\u2018An example for the analysis of linkage by means of the transformed Willis method\u2019\u2019 (Fig. 8), is composed of a driving link 1\u2019 and a tetrad, thus according to the previous section it would be efficient to fix link 3\u2019 and change the driving link to 4\u2019, as shown in Fig. 11b. The transformed truss, dual to the transformed linkage of Fig. 11b, appears in Fig. 12a. The truss of Fig. 12 is a simple truss and thus it can be solved by one of the efficient methods available for solution of such trusses. One of such methods is the wellknown graphical method, called Maxwell-Cremona diagram [12], shown in Fig. 12b. One can see that consistently with the results reported in [2], the Maxwell Cremona diagram of the truss of Fig. 12a is identical to the image velocity diagram of its dual mechanism, as appears in Fig. 8. The solution of the transformed truss can now be substituted into Eq" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003122_978-90-481-3141-9_4-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003122_978-90-481-3141-9_4-Figure2-1.png", + "caption": "Fig. 2 DCB cylinder specimen", + "texts": [], + "surrounding_texts": [ + "Delamination specimens were machined from composite cylinders consisting of E-glass fibers in an epoxy resin. The internal diameter of the cylinders was 160 mm and the nominal wall thicknesses were 6 and 12 mm (12 and 24 layers). The lay-ups of the cylinders were \u0152\u02d9\u2122 6 and \u0152\u02d9\u2122 12, where \u2122 D 30\u0131; 55\u0131 and 85\u0131. A 58 mm long and 13 m thick, release agent coated aluminum film, was inserted at the mid-plane of the cylinders during the filament winding process to define a starter delamination crack, see Fig. 1. The film insert was wrapped around the circumference of the cylinder to enable machining of multiple test specimens from each cylinder. After filament winding, the cylinders were cured at 160\u0131 C for 3 h. The average fiber volume fraction was 0.61 for all cylinder lay-ups. Beam fracture specimens of a nominal length of 200 mm and a nominal width of 18 mm were cut from the cylinder wall for the subsequent delamination tests as schematically illustrated in Fig. 1. The beam axis was parallel to the cylinder axis producing straight beams with a curved cross-section. Figures 2\u20134 show the DCB, ENF and MMB cylinder specimens and loading principles. In order to accommodate the curved cross-section of the beam fracture specimens, contoured aluminum loading tabs were fitted to the DCB and MMB specimen and contoured loading pins and supports were attached to the ENF and MMB test fixtures as shown in Figs. 2\u20134. Further experimental details are provided in Refs. [10\u201312]. Glass/epoxy cylinders of lay-up \u0152\u02d955 n were manufactured for the external pressure tests using the filament winding process. Cylinder internal diameters were 55 and 175 mm, with wall thicknesses of 6.5 and 19 mm. The fiber volume fraction was 0.68. E-glass fibers in these cylinders were impregnated with the same epoxy resin as the cylinders used for delamination specimens. After winding, the cylinders were cured at 125\u0131 C for 7 h. In order to simulate fabrication defects in the form of delaminations that may arise during cure of thick cylinders due to exothermic heating, 50 mm square aluminum foil layers of 13 m thickness, coated with release agent on both sides, were introduced at different thickness locations during filament winding into some of the 55 mm diameter cylinders. The wall consisted of 12 layers and defects were placed between the third and fourth (referred to as 1/4 thickness), sixth and seventh (mid-thickness), and ninth and tenth layers (3/4 thickness) where ply #1 is the inner layer and ply #12 is the outer layer of the cylinder. Carbon fibre (T700) reinforced epoxy cylinders, (the same epoxy resin as for the glass reinforced cylinders), were filament wound with the same dimensions and cure cycle as for the glass/epoxy cylinders. Carbon (AS4)/PEEK (poly-ether-ether-ketone) cylinders were produced with the same dimensions by tape laying." + ] + }, + { + "image_filename": "designv11_69_0002257_s0263574707003426-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002257_s0263574707003426-Figure1-1.png", + "caption": "Fig. 1. Algorithm of motion selection for online hand\u2013eye calibration.", + "texts": [ + " According to the above three observations, the following \u201cgolden rules\u201d are used for motion selection. Rule 1: Try to make (ka,i , ka,i+1) (which is equal to (kb,i , kb,i+1)2) large, which is no less than \u03b10. Rule 2: Try to make \u03b8i large, which is no less than \u03b20. Rule 3: Try to make \u2016ta,i\u2016 small, which is no bigger than d0. Let us denote the ith sample of hand\u2013eye pose and motion by (Pi, Qi) and (Ai, Bi), respectively, in the following paper, and \u03b10, \u03b20, d0 are thresholds determined by experience. Also (A\u2032, B \u2032) and (A\u2032\u2032, B \u2032\u2032) are selected motion pairs for the calibration (see Fig. 1). In motions A\u2032 and A\u2032\u2032, the rotation axis, rotation angle and translation are denoted by (k\u2032 a, \u03b8 \u2032 a, t \u2032 a) and (k\u2032\u2032 a , \u03b8 \u2032\u2032 a , t \u2032\u2032a ), respectively. At the beginning of the calibration, we need to estimate (A\u2032, B \u2032), which is first recovered from (P1, Q1) and (P2, Q2). If \u03b8 \u2032 \u2265 \u03b20 and \u2016t \u2032a\u2016 \u2264 d0, we can claim that (A\u2032, B \u2032) has been found. Otherwise, continue to compute (A\u2032, B \u2032) from (P1, Q1) and (P3, Q3) and judge the value \u03b8 \u2032 and \u2016t \u2032a\u2016 in the same way as before. Repeat this procedure until \u03b8 \u2032 and \u2016t \u2032a\u2016 fulfill the given conditions" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002608_iembs.2008.4649685-Figure5-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002608_iembs.2008.4649685-Figure5-1.png", + "caption": "Fig. 5. Marking point of the fingertip.", + "texts": [ + " The ratio I of Weber\u2019s Law is 0.046. During the experiment under the pressed condition, we monitor how much blood flow is, while we start the experiment with roughly 30 % of the blood flow under the non-pressed condition. All of subjects are 24 persons, the male with the age distribution of 21 \u223c 25, where the blood flow rate under the non-pressed condition is 18 \u00b1 7 ml/min/100g and that under the pressed condition is 6 \u00b1 3 ml/min/100g. We executed all experiments under the room temperature of 26 \u25e6C. Fig.5 shows the test point where all tests are done at the marking point indicated by \u00d7. This is for coping with the different size of fingers where L and W denote the length between the distal phalange and the width of fingertip, respectively. Fig.6 shows the main result of this work to confirm how the touch sensitivity changes under the pressed condition, where each point is computed with the time average during every 1 min. The horizontal and the vertical axes denote the time and the touch sensitivity, respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002407_3.3921-Figure5-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002407_3.3921-Figure5-1.png", + "caption": "Fig. 5 Root locus.", + "texts": [ + " 3, the ordinate is K**, which is defined by \u00a3** = (T2K)(TL/T)/T2F (13) and the abscissa is rX, where TL/r appears as a parameter. These curves show the borders between stability and instability of the closed system. For example, the system is stable on the shaded side for TL/r = 4.0. In Fig. 4, the ordinate is TCO and the abscissa is rX, where TL/r appears as a parameter. The dotted line in the figure shows the boundary of the stable closed system. The effect of gain K** on the stability boundary also is shown in Figs. 5a and 5b by the root-locus method applied to the characteristic equation (8). Figure 5a indicates that as K** is increased for a large value of TL/T, an unstable root enters into the stable region and then returns to the unstable region again. On the other hand, Fig. 5b indicates that as 7\u00a3** is increased for a small value of TL/T, the unstable root enters into the stable region, while another stable root enters into the unstable region. Returning to Fig. 3, we find that for a set of r2F, rX, and TL/T, there exists for the value of K** a region that makes the xL/T = 2.5 3.0 4.0 5.0 6.0 closed system stable. We shall denote the maximum and minimum values of K** which bound the stable region by Kmax** and Kmin**, respectively and corresponding values of K by Kmax and J\u00a3min, respectively, where K** and K are related by Eq" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002819_6.2007-3813-Figure10-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002819_6.2007-3813-Figure10-1.png", + "caption": "Figure 10. Rear-Blowing Slot.", + "texts": [ + " Tangential Rear-Blowing Slot As the experiments progressed it was decided to also experiment with tangential blowing over the tail of the projectile, thus additional hardware was created to blow air over the boat tail of the model. In order to affect this change without creating a new model a cover was designed to fit closely over 10 pressure taps in the rear of the 2:1 Page 8 bullet. The device was designed to leave only a slot of height 0.254 mm between the cover and the surface of the bullet which has an azimuthal extent of 60 degrees. Figure 10 shows the cover attached to the rear of the bullet over the pressure taps. Pressure taps 44-51, 56, and 59, shown in Figure 11, were used to inject air into the slot cover section out of the small slot directed tangential to the surface of the bullet. This mass injection device was tested for a range of jet pressures between 25 psi and 85 psi. GTRI\u2019s success with guiding supersonic rounds with mechanical pins that rotate into the flow (ref__) prompted the testing of similar devices for the present subsonic round" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000147_thc-2003-11202-Figure4-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000147_thc-2003-11202-Figure4-1.png", + "caption": "Fig. 4. A typical programmed elbow trajectory (left) at a certain fixed shoulder angle q1 and shoulder trajectory (right) at a fixed elbow angle q2.", + "texts": [ + " In both cases the wrist was not fixed and was allowed to move freely since the deviation from the neutral position was found to be only a few degrees. Before the particular measurements, ten different circular trajectories (not shown here) were programmed into the robot for each subject. The first five measurements concentrated on the elbow angle smooth variation from one boundary angle to the other and backwards, with the shoulder fixed at different angles (\u221268\u25e6,\u221240\u25e6,+16\u25e6,+10\u25e6,+36\u25e6). The shoulder angle was kept constant by programming an appropriate trajectory, using no additional fixation mechanisms Fig. 4\u2013left side). The second set of trials focused on movements of the shoulder joint, with the elbow kept at constant angles (20\u25e6, 30\u25e6,41\u25e6,49\u25e6,59\u25e6). For fixating the elbow angle, an orthosis was used, which allowed angle adjustments from extension to a flexion angle of 85 degrees (Fig. 4\u2013right side). The mass of the orthosis utilized for shoulder movements was included into the calculation of the G(q) matrix in Eq. (2), which describes the new upper and forearm masses and center of gravity locations as mi and li: m1 = mua + muo, m2 = mfa + mlo l1 = a1 \u2212 luamua + luomuo mua + muo (9) l2 = lfamfa + lfomfo mfa + mfo Here the ua and fa indices refer to the upper arm and forearm, whereas uo and fo describe the upper and lowerorthosis parts. The orthosis masses and centers of gravity were accurately determined before the experiment" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001979_50037-5-Figure7.48-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001979_50037-5-Figure7.48-1.png", + "caption": "Figure 7.48 Pull-out stress contours on the pin.", + "texts": [ + " From Utility Menu choose Select \u2192 Everything. Next, from ANSYS Main Menu select General Postproc \u2192 Read Results \u2192 By Load Step. The frame shown in Figure 7.47 appears. As shown in Figure 7.47, [A] Load step number = 2 was selected. Pressing [B] OK implements the selection. From ANSYS Main Menu select General Postproc \u2192 Plot Results \u2192 Contour Plot \u2192 Nodal Solu. In appearing frame (see Figure 7.40), the following are selected as items to be contoured: [A] Stress and [B] von Mises (SEQV). Pressing [C] OK implements selections made. Figure 7.48 shows stress contours on the pin resulting from pulling out the arm. 7.2.2 Concave contact between cylinder and two blocks Configuration of the contact between cylinder and two blocks is shown in Figure 7.49. This is a typical contact problem, which in engineering applications is represented by a cylindrical rolling contact bearing. Also, the characteristic feature of the contact is that, nominally, surface contact takes place between elements. In reality, this is never the case due to surface roughness and unavoidable machining errors and dimensional tolerance" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000131_melcon.1994.380908-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000131_melcon.1994.380908-Figure1-1.png", + "caption": "Figure 1: Moving items of typical backhoe excavator, and assignment of their coordinate frames", + "texts": [ + " A backhoe excavator is made of three main units: (i) a mounting or travel unit which may be a cmwler with heavy-duty chassis, or a heavy framed rubber-tired chassis; (ii) a revolving unit or superstructure which carries engine, transmission, and operating machinery; (iii) an arm which includes a bucket at its end. An important distinguishing feature of a backhoe is its working space that suits far digging below its own base level. -Backhoes for heavy loads are always equipped by hydraulic actuators. The arm of a backhoe consists of three strong structural members; a boom, a stick, and a bucket as seen in Figure 1. The automation of these machines are needed to reduce the uncomfortable operating conditions in mining and construction, as well as planetary excavation requirements [I]. The automation of the excavators require both position and force control along the trajectory of the bucket [2]. A pre-planned excavation requires the representation of the physical shape and properties of the excavation location, and generation of the bucket trajectories to dig and remove the soil [3]. The movement of the bucket against the soil is one of the most critical action in controlling the backhoe, because of the unpredicted soil properties and unexpected soil slides", + " The dynamic control of bucket position against an unpredicted soil resistance is simulated, and the control of the digging forces through the bucket trajectory by modifying the pre-planned trajectory of the bucket is tested on a typical digging trajectory. KINEMATICS AND DYNAMICS OF BACKHOE During digging at a certain point on the excavation trajectory, both the crawler and the rotational super-structure bodies are stationary, and thus the kinematic and dynamic model is reduced to 3 degree of freedom. Kinematic solution of the arm is accomplished in the form of homogeneous transformation matrix by using Denavit-Hartenberg (DH) notation [SI. The assignment of the joint displacement variables and the coordinate frames are as shown in Figure 1 and (D-H) parameters are as shown in Table 1. The forward kinematic transformation T = AlA2A3 which is obtained using the homogeneous transformation matrices A, a = 1,2 ,3 of the boom, stick and bucket, respectively. It converts the coordinates in the bucket frame into the fixed superstructure frame. using the Lagrange-Euler formulation, the dynamics where T = [ T I , T Z , ~ 3 3 1 ~ is the vector of joint torques applied to the boom, stick and bucket by the hydraulic actuators; Td = [ T d J , Tdd,2 , Td,3IT is the loading torque vector due to digging, D(q) E R 3 x 3 , h(q,q) E R3 and g(q) E R3 are coefficient matrices and vectors of the arm dynamics in Lagrange-Euler form" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000669_1350650042128076-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000669_1350650042128076-Figure1-1.png", + "caption": "Fig. 1 Coordinate system of the shaft and connecting rod", + "texts": [ + " With the expression of the pressure formulation of Rhode and Li, and by integration over the axial direction of the bearing, the Reynolds equation becomes 2 3 q qx G qp qx 8 L2 p \u00bc U qF qx \u00fe qh qt In this expression, p is the pressure in the centre plane of the film, h is the film thickness, U is the relative surface velocity of the shaft with respect to the housing, and F and G are functions of the viscosity m and are given by F \u00bc J1 J0 , G \u00bc J2 J2 1 J0 with Ji \u00bc \u00f0h 0 xi m\u00f0x\u00de dx i \u00bc 0, 1 For a circular bearing (Fig. 1), the film thickness h at point M located at angle y on the internal bearing housing is given by h\u00f0y\u00de \u00bc ho\u00f0y\u00de \u00fe het\u00f0y\u00de Proc. Instn Mech. Engrs Vol. 218 Part J: J. Engineering Tribology J04603 # IMechE 2004 at East Carolina University on April 24, 2015pij.sagepub.comDownloaded from where het\u00f0y\u00de is the summation of elastic and thermal deformations of the solids and ho\u00f0y\u00de is the nominal film thickness which depends on the radial clearance C and the eccentricity e for a rigid bearing ho\u00f0y\u00de \u00bc C ex cos\u00f0y\u00de ey sin\u00f0y\u00de It is supposed that the lubricant viscosity varies only with temperature", + " On the film reformation boundaries, the pressure is zero and the mass flow conservation conditions developed by Bonneau and Hajjam are assumed [23]. The temperature in the film is obtained by solving the transient energy equation rCp qT qt \u00fe u qT qx \u00fe v qT qy \u00bc kf q2T qy2 \u00fe m qu qy 2 where T is the temperature and r, Cp and Kf are respectively the density, the specific heat and the thermal conductivity of the lubricant. The velocity component u depends on the circumference pressure gradients and journal velocity and is given by (Fig. 1) u\u00f0x, y\u00de \u00bc qp qx I1 I0 J1 J0 \u00feU 1 I0 J0 where Ii \u00bc \u00d0 y 0 xi=\u00bdm\u00f0x\u00de dx. The velocity component across the film, v, is expressed from the mass conservation equation by integration through the film thickness. Shaft and connecting rod temperatures are obtained by solving the transient heat equation which can be reduced for the present case rsCs qT qt \u00bc ks q2T qx2s \u00fe q2T qy2s In this equation, rs, Cs and Ks are respectively the density, the specific heat and the thermal conductivity of the solid, and xs and ys are the solid coordinates", + " The analogue signal obtained from the pressure gauge is digitized and transmitted as infrared to the receiver which is located outside the oil chamber. Figure 8 shows the position of the pressure gauge, thermocouples and strain gauges on the connecting rod. The pressure transducer is located on the middle plane of the bearing. The crankpin design makes it possible to change the gauge position in order to measure the pressure in the oil film for every 158 Proc. Instn Mech. Engrs Vol. 218 Part J: J. Engineering Tribology J04603 # IMechE 2004 at East Carolina University on April 24, 2015pij.sagepub.comDownloaded from housing angle y (Fig. 1), and also to locate the feeding hole in the best position (about 458 crankpin). Oil film pressure diagrams, obtained for different positions, 608 shifted on the bearing at 150 r/min, are presented in Fig. 9. In Fig. 9a, the oil film pressure reaches its maximum at about 3608 crankshaft angle, an angle for which the value of the force is maximum. The maximum oil film pressure is about 0.615MPa. This simulates the firing phase in the vicinity of 3608 crank angle j. Figures 9b to d present the pressure evolution for other positions of the pressure gauge on the crankpin" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003285_1.39707-Figure4-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003285_1.39707-Figure4-1.png", + "caption": "Fig. 4 Quadrotor reference frames.", + "texts": [ + " The bias instability coefficient, B, is proportional to Allan standard deviation of the horizontal segment [12] B = \u03c3A \u221a \u03c0 2 n 2 (5) The noise coefficients extracted from the Allan variance analysis are listed in Table 1 for each of the three gyroscopes. The measured coefficients are found to be in reasonable agreement with the specifications provided by the manufacturer, though the z-axis gyroscope exhibits a larger than expected bias instability. The platform attitude is the orientation of a body fixed-reference system with respect to inertial space or a predefined navigation frame. The body reference frame, Fig. 4, is defined analogous to that of a typical fixed-wing aircraft with the origin located at the center of gravity; the x-axis is aligned with one arm, the z-axis pointing down, and the y-axis aligned with a second arm consistent with a right-hand coordinate system. For the planned short duration flights, the rotation of the earth can be ignored and the axes of the inertial reference frame are set to north, east, and down, corresponding to the navigation frame. The attitude is described by three Euler angles, \u03c8 , \u03b8 , and \u03c6" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001231_sice.2006.314734-Figure3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001231_sice.2006.314734-Figure3-1.png", + "caption": "Fig. 3. Mounting structure for bearings and related motors.", + "texts": [ + " A laboratory set is prepared to examine theoretical results, which consists of a vibration sensor and its driver circuit, amplifying and filter circuits, a data acquisition card and a digital computer. To obtain the vibration of defective bearings an apparatus is made consisting of a shaft which two bearings are assembled on its ends and two housings to retain the bearings, accompanied by two induction machines working as motor and generator. By proper selection of pulleys and belts the shaft can rotate in different velocities. All of these parts are mounted on a frame (Fig. 3). The applied bearing can be easily disassembled. The complete bearing and its defective parts are shown in Fig. 4. Bearing has 28 balls in two rows, ball diameter is 7.9mm, inner raceway diameter is 35mm, and outer raceway diameter is 57mm. The vibration sensor is an accelerometer, with a bandwidth of more than 10 kHz, the selected data acquisition card, is a 16-bit card with a maximum sampling rate of 250ks/s for all channels. The circuits consist of a current source to drive the vibration sensor and a low-pass 6th order Butterworth filter with a cut-off frequency of 1 kHz" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002400_s12239-008-0037-2-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002400_s12239-008-0037-2-Figure2-1.png", + "caption": "Figure 2. Rear rigid-axle movement (Watanabe et al., 2002).", + "texts": [ + " OVERVIEW OF THE SIMULATION 2.1. Vehicle Model The CarSim (Version 6) simulation model (MSC Co., USA) was used as the vehicle model in this study. Details are provided in (Watanabe et al., 2002). Figure 1 shows a schematic diagram of the vehicle model. Table 1 lists the number of primary bodies and the number of degrees of freedom in the vehicle model. For example, the rear axle is *Corresponding author. e-mail: nozaki@mech.kindai.ac.jp rigid and has vertical movement and rotation of the axle (see Figure 2). Table 2 lists the parameters for the vehicle used in this simulation. Vehicles having four different front and rear weight distribution ratios (40:60, 45:55, 50:50, 55:45) were used in the simulation, as shown in Table 2. Figure 3 shows the cornering force characteristics for four different axle loads per road wheel. By adding the slip angle and the slip ratio for a specific time, the concept of a friction circle based on the calculation of the combined characteristics of CarSim was applied to the model" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003176_cca.2010.5611314-Figure9-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003176_cca.2010.5611314-Figure9-1.png", + "caption": "Fig. 9: 3D scan coordinates D p i to camera coordinates Dc i .", + "texts": [ + " By combining this information with a 3D scan of the puppet in a perspective projection, a 2D mask of each of these puppets can be determined. For this the pixel coordinates D p i in the 3D scan have to be transformed to 3D pixel coordinates in the camera frame Dc i , according to Dc i = Rr s(O s p +D p i )+Oc r (8) where Rr s is the rotation matrix, Os p is the transformation vector from the rod to the heart of the puppet and Oc r is the transformation vector from the heart of the puppet to pixel coordinates Di. A corresponding graphical interpretation is given in Fig. 9. constant switching between the selected puppet. Tk = 1, if i yw k < 2l ; 1, if i yw k < 1r and Tk\u22121=1; 3, if i yw k > 2r; 3, if i yw k > 3l and Tk\u22121=3; 2, else (15) Create rod reference point Now that a puppet is selected the equation for the reference point of the rod is r yw k = i yw k +Pb(1\u2212Tk) (16) where Pb is the equidistant pitch between the puppets. Kick ball A kick is simply performed when the estimated position of the ball is near the rod. In the low level scheme the references for the rods are executed" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003355_winvr2010-3730-Figure3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003355_winvr2010-3730-Figure3-1.png", + "caption": "FIGURE 3. Haptic interaction through phantom omni. A grasper tool was implemented to randomly grasp a point in the brick and pull or push the brick through manipulating the grasped point.(a) and (b) are the results rendered by descretized points. The point dragged by the grasper is rendered with a different color; (c) is the result rendered by polygon meshes. The green vector in (c) is the user input force vector, Young\u2019s modulus is set to be 107Pa.", + "texts": [], + "surrounding_texts": [ + "The first experiment was the gravity test. Where we set up a rectangular object composed by 90 points (3 by 3 by 10). The TABLE 1. Analysis of cut off distance in weighting function. The first column shows the different h values; the second column is the corresponding Volume plot of each particle. The size of each cube which represents each point is proportional to the volume of the point. In case (a), (b), and (c), h is set to 0.10 meters, 0.15 meters and 0.20 meters respectively. Investigation of the effect of cut off distance h/m (a) h=0.10 (b) h=0.15 (c) h=0.20 points were evenly distributed; that is, the spatial distance between each point was constant. In this case, we set the distance to be 0.015 m, so the total dimension of the object was 0.135m by 0.030m by 0.030m. Young\u2019s Modulus was set to 105Pa and Poisson\u2019s ratio to 0.4. The cut off distance, h, in equation (7) was set to 0.1 m. Figures (2(a)) to (2(c)) illustrate that When object is affected by gravity, the points in the middle of the body are dragged downward. Figures (3(a)) and (3(b)) illustrate the haptic interaction with the point-based rectangular model. Figure (3(a)) displays the pulling up and figure (3(b)) illustrates pushing down. The entire brick model deforms accordingly. Figures (4(a)) and (4(b)) illustrate the haptic feedback force when the user is interacting with the object. Figure (4(a)) shows the force when the user is continuously pulling the object; figure (4(b)) shows the feedback force when the user is interacting with the object at random locations." + ] + }, + { + "image_filename": "designv11_69_0000105_s0263574702004630-Figure3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000105_s0263574702004630-Figure3-1.png", + "caption": "Fig. 3. The Coordinate system of the elastic space bar at static in Von Mises2 \u2013 (a) compared with that of the arm we discussed here \u2013 (b).", + "texts": [ + " It is not too difficult to work out the stiffness matrices for different systems but here we are only interested in the stiffness matrix of a simple beam. In reference [10] von Mises shows that the stiffness matrix of a uniform beam is given by, EJx l 0 0 0 0 0 0 EJy l 0 0 0 0 0 0 GJ l 0 0 0 K = 0 0 0 12EJy l 3 0 0 0 0 0 0 12EJx l 3 0 0 0 0 0 0 EA l (4) where, as usual, E is Young\u2019s modulus, G the shear modulus, A the cross sectional area of the beam, J, Jx, Jy are the relevant polar and second moments of area and l is the length of the beam. This result hold, is the coordinate frame fixed in the middle of the beam (see Figure 3(a)). For our purposes we need to know the stiffness matrix in the fixed hub coordinates (Figure 3(b)). In general, if H is the active rigid transformation which moves the old http://journals.cambridge.org Downloaded: 11 Mar 2015 IP address: 169.230.243.252 coordinate frame to the new one, then the stiffness matrix in the new coordinates will be given by the transformation, K=HTK H (5) This can be done in two stages: first, we translate to the hub, 4EJx l 0 0 0 6EJx l 0 0 4EJy l 0 6EJy l 0 0 0 0 GJ l 0 0 0 K0 = 0 6EJy l 0 12EJy l 3 0 0 6EJx l 0 0 0 12EJx l 3 0 0 0 0 0 0 EA l (6) Then we rotate about the hub; however, it seems to be more useful to leave this transformation un-evaluated and simply write, K=Hr K0H T r (7) where, Hr = cos 0 sin 0 0 0 0 1 0 0 0 0 sin 0 cos 0 0 0 0 0 0 cos 0 sin 0 0 0 0 1 0 0 0 0 sin 0 cos (8) The momentum of a rigid body can also be represented as a co-screw, P= j p (9) i" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003196_s1064230710050151-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003196_s1064230710050151-Figure1-1.png", + "caption": "Fig. 1. Schematic diagram of injection engine.", + "texts": [ + " An interesting predic tion of the development of control systems for auto mobile industry is given in the survey [29]. In Russian papers, however, this important direc tion of modern theory and practice of automatic con trol has not been adequately presented. This paper is aimed at filling this gap and attracting attention of Russian experts to the topical problem of automatic control of internal combustion engines. OF INJECTION ENGINE The simplified typical schematic diagram of an injection engine is shown in Fig. 1. The engine oper ates as follows. Air flows through throttle gate 1 to intake manifold 2. The amount of supplied air is con trolled by changing the inclination angle of the throttle gate which is determined by the driver by pressing throttle pedal 3 or by a special controller. Injector 5 which injects fuel into the intake manifold is situated near intake valve 4. During the intake cycle (when pis ton 6 moves to the bottom dead point, the intake valve opens and air together with fuel get to cylinder 7, forming combustible mixture" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000834_j.matcom.2004.05.010-Figure5-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000834_j.matcom.2004.05.010-Figure5-1.png", + "caption": "Fig. 5. Double pendulum with flexible upper link and orthogonal joint axes.", + "texts": [ + " The position of its centre of mass relative to where it is attached to joint 2 (and with respect to the frame fixed in the bar) is [0,\u2212L/2, 0]T , and its three principal moments of inertia are I1 = (1/12)ML2, I2 = 0 and I3 = (1/12)ML2. Consequently, the final input is: IsFlexible[2] = False; m[2] = M; d[2, 2] = {0,\u2212L/2, 0}; I1[2] = 1 12 \u2217 M \u2217 L2; I2[2] = 0; I3[2] = 1 12 \u2217 M \u2217 L2; Once the input data is ready, MultiFlex then reads it from a file and returns the following equation of motion (with respect to the coordinate q[2]): LM(3 Sin[q[2][t]](g \u2212 Cos[3[1] + t [1]]P[1] [1]2)+ 2Lq[2]\u2032\u2032[t])+ 6c[2]q[2]\u2032[t] = 0 This problem is shown in Fig. 5. The top of the flexible beam is attached to the inertial frame via a revolute joint which rotates about the z0 = z1 axis. Also, the top of the rigid bar is attached to the bottom of the flexible beam via a revolute joint which rotates about the x2 axis. Note, the x2 axis is parallel to the x0 axis when the multibody system is in its reference/rest configuration. As with the previous example, the number of bodies in the present system is 2. Furthermore, as before, the \u201cparent\u201d of body 1 is body 0 and the position of the attachment point of joint 1 (the top revolute joint) on body 0 is the origin of the inertial frame", + " The first thing to declare is the direction in which the flexible beam is pointing when the multibody system is in its reference/rest configuration. This is defined by the MultiFlex input variable WhichDirection, which takes the value 5 when the beam is pointing in the negative y direction when the system is at rest. For a flexible beam it is also necessary to state the distance down the undisturbed beam to the attachment point of the joint linking it to its parent body. In the present case, this is s11 (see Fig. 3) and from Fig. 5 it can be seen that this distance is zero. The final input for a flexible beam is to give the number of modes used for each of the translational and rotational deformations of the beam. That is, the user must supply the number of modes Nt1, Nt2, Nt3, Nr1, Nr2 and Nr3 used for u1, u2, u3, \u03b81, \u03b82 and \u03b83, respectively (see Eq. (9)). For the present problem, 2 modes are used for each of the 6 deformations of the flexible beam. Therefore, to summarise, the required input for body 1 is: IsFlexible[1] = True; WhichDirection[1] = 5; s[1, 1] = 0; Nt1[1] = 2; Nt2[1] = 2; Nt3[1] = 2; Nr1[1] = 2; Nr2[1] = 2; Nr3[1] = 2; A consequence of the above is that, with the 1 degree of freedom associated with each of the 2 joints, the present example represents a 14 degree of freedom problem" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002609_icma.2008.4798870-Figure5-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002609_icma.2008.4798870-Figure5-1.png", + "caption": "Fig. 5. The end positions lie on different calibration objects (4. Case).", + "texts": [ + " \u2022 If the robot base frame is the reference frame of the connecting vector, the computations slightly differ: u\u0302DH,s,e(\u03bd ,q j) = 0T DH,s,e n (\u03bd ,q j)nTc cu\u0302, v\u0302DH,s,e(\u03bd ,qk) = 0T DH,s,e n (\u03bd ,qk)nTc cv\u0302, where cu\u0302 represents the direct measurement of the position of point U within the camera frame. The connecting vector \u2212\u2192 UV is specified w.r.t. the base frame, thus resulting in the nonlinear regression function below: min \u03bd H \u2211 h=1 \u2225\u2225\u2225(uDH,s,e(\u03bd ,q j)+ \u2212\u2192 UV )\u2212vDH,s,e(\u03bd ,qk) \u2225\u2225\u22252 2 . (8) The scenario is shown in Fig. 5. As you can see, different tuples of configurations can be selected. Some cases might not support an accurate determination of the geometric and nongeometric parameters. The selection of tuples must be made regarding the size and geometry of the robot, and the set of identification parameters. By considering configurations that result in end positions located not too close to each other, the robot\u2019s workspace can be enlarged. The importance of these case differentiations becomes clear in Section VI" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000700_j.amc.2003.11.041-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000700_j.amc.2003.11.041-Figure1-1.png", + "caption": "Fig. 1. Conical bearing gap.", + "texts": [ + " Assuming the ratio of the radial component of lubricant velocity to the peripheral velocity of the journal bearing to be of order of the relative radial clearance and hence, neglecting the terms having the order of the latter, than equations of: motion, continuity and energy, become q\u00f0V r\u00deVai \u00bc 1 ha oP oai \u00fe K o oa2 oV a1 oa2 2 2 4 \u00fe oVa3 oa2 2 n 1 2 oVai oa2 3 5 rB2Va1 for i \u00bc 1; 3; \u00f06\u00de oP oa2 \u00bc rB2Va2; \u00f07\u00de 1 ha1 o oa1 \u00f0qVa1\u00de \u00fe o oa2 \u00f0qVa2\u00de \u00fe 1 ha3 o oa3 \u00f0ha1qVa3\u00de \u00bc 0; \u00f08\u00de q\u00f0V r\u00deCvT \u00bc o oa2 v oT oa2 \u00fe k oVa1 oa2 2 \u00fe oVa3 oa2 2 n 1 2 oVa1 oa2 2 \" \u00fe oVa3 oa2 2 # \u00fe rB2V 2 a1: \u00f09\u00de For 0 < n < 1, Eqs. (6)\u2013(9) describe the magneto-hydrodynamic flow of a nonNewtonian power law lubricant through the curvilinear (in width direction) gap of a slide bearing. For n \u00bc 1, the equations listed above hold for a Newtonian gap flow [6]. The unknown functions Vai, p and T may be found by solving Eqs. (6)\u2013(9). In the special case of a conical bearing gap, the curvilinear coordinates ai\u00f0i \u00bc 1; 2; 3\u00de become: a1 \u00bc /, a2 \u00bc y, a3 \u00bc x respectively, see Fig. 1. Thus, the Lame\u2019s coefficients are hu \u00bc x cos a , hy \u00bc hv \u00bc 1 where a denotes the slope of the generating line of conical surface. The components of the local lubricant velocity vu, Vy , Vv, the hydrodynamic pressure p and temperature T are now assumed to be of the following forms: Vu \u00bc xl cos a Vu1; Vy \u00bc ~w1 v l Vy1; Vx \u00bc v l Vx1 P \u00bc q0 v 2 P1; T \u00bc T0 \u00bc Ec Pr T0T1 ) ; \u00f010\u00de where Vu1, Vy1 and Vx1 are the dimensionless components of the local lubricant velocity in the u, y, and x directions, respectively, P1 is the dimensionless hydrodynamic pressure T1 is the temperature, l is the length of the cone generating line, e denotes the height of the gap, q0 is the dimensional characteristic value of lubricant density, T0 denotes the ambient temperature, x is the angular velocity of the journal, Ec and Pr are the Eckert and Prandtl numbers, respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000529_s0043-1648(02)00034-0-Figure3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000529_s0043-1648(02)00034-0-Figure3-1.png", + "caption": "Fig. 3. Shape of a Rockwell indentation on a large roller.", + "texts": [ + " This study focused on the case of double roller testing with Rockwell indentations artificially induced on the large roller as shown in Fig. 1. Contact pressure variation and the internal stresses below the surface of a defect-free small roller were calculated for when an indentation on a large roller surface moved through the contact area as shown in Fig. 2. The 3D contact problem analysis program TED/CPA\u00ae (TriboLogics Corporation) was employed for the calculations. This program is designed to solve elastic contact problems by using the boundary element method [14,15]. Fig. 3 displays the top and cross-sectional views of a Rockwell indentation induced with a load of 147 N on a large roller surface. As indicated in Fig. 3, it had a 0.18 mm width, a spherical indentation depth of 0.009 mm, and a 0.5 mm radial curvature. The circumferential shoulder edge had a radial curvature of 0.025 mm. A Rockwell indentation was modeled by using these measured values for the calculations. Due to the limitation that contact models must be defined as rotationally symmetric bodies, some modification was necessary in order to create the contact model of a large roller surface with an indentation. The indentation is a circular dent which can be assumed to be rotationally symmetric about its central axis", + " The \u03b8 value in each figure is the angle between the center of the indentation and the center position of the contact as indicated in Fig. 2. In the cases with indents, the pressure distributions show sharp pressure peaks at the edges of the indentation. Fig. 8 shows the relationship between Pmax and \u03b8 , indicating that Pmax is almost triple, 12 GPa, in the range of \u03b8 = \u00b10.5\u25e6. The results of Figs. 7 and 8 were obtained by using the actual configuration of a Rockwell indentation where the edge curvature was set at r = 0.025 mm (as shown in Fig. 3). The relationship between the values of Pmax and r with \u03b8 = 0\u25e6 is shown in Fig. 9, exhibiting that the magnitude of Pmax varies inversely to the value of r. Thus, Pmax increases as r decreases. As r increases, Pmax decreases to asymptotically approach 4.0 GPa, the pressure seen with no indentation present. Accordingly, it should be noted that the sharper an indentation edge, the greater the influence it has on a contact area. Fig. 10 shows the calculated equivalent stress \u03c3 c at the various subsurface depths under the defect-free small roller contact zone for each contact angle shown in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003365_amr.148-149.1741-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003365_amr.148-149.1741-Figure1-1.png", + "caption": "Fig. 1 The axisymmetric and 3-D models", + "texts": [ + " No part of contents of this paper may be reproduced or transmitted in any form or by any means without the written permission of Trans Tech Publications, www.ttp.net. (ID: 128.6.218.72, Rutgers University Libraries, New Brunswick, USA-04/06/15,11:24:43) Finite element model. Three standard ISO M14 bolts with three different pitches P=1, P=1.25 and P=2 are chosen to study the helical effect and the elastic-plastic behavior of threaded connection. The finite element analyses are implemented using the software ABAQUS. The axisymmetric and 3-D models are shown in Fig. 1. Nine threads are engaged. One pitch of internal thread protrudes from the nut. The threaded connection with 8 elements on the curved profile of thread root is employed. It is fine enough for the analysis. No residual stresses in the thread are considered. A simpler 3-D model is generated by rotating the axisymmetric thread cross section model with one pitch height helically around the bolt axis. Thirty-six elements are used around the circumference of the model. So the thread profile is constant and continuous along the helix as shown in Fig. 1(b). To avoid geometric singularity, a small hole around the bolt axis of 0.09D (D-bolt nominal diameter) has been modeled as shown in Fig. 1(c). Material Properties and Constraints. The material properties used in the analysis for the nut and bolt are the same. The stress-strain relationships of the material are determined by the uniaxial tension tests of specimens that are manufactured from the objective bolts. The Young\u2019s modulus E is 174 GPa, the Poisson\u2019s ratio \u03c5 is 0.3 and the initial yield stress \u03c30 is 413MPa. The coefficient \u00b5 between engaged threads is taken as 0.1. The nut was constrained along its upper surface in the axial direction" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001961_s1387-2656(06)12004-9-Figure9-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001961_s1387-2656(06)12004-9-Figure9-1.png", + "caption": "Fig. 9. Schematic of the different approaches taken to assemble tethered bilayer architectures on solid substrates. Method 1 showed the direct adsorption of vesicles onto the quasi-dried polymer on a solid support. Method 2 shows the polymer adsorption on the lipid bilayer formed on the bare solid support. Method 3 is referred to as the top-down concept with a lipid\u2013lipocopolymer layer being transferred from the water\u2013air interface to a solid support, pre-coated with a reactive monolayer capable of covalently binding to some of the polymer units. Method 4 illustrates the bottom-up layer-by-layer approach with the substrate being first coated by a reactive monolayer to which a polymer \u2018\u2018cushion\u2019\u2019 binds after adsorption from solution. The final monolayer, also pre-organised at the water\u2013air interface contains some reactive \u2018\u2018anchor\u2019\u2019 lipids, able to bind to the tethering polymer. Both methods 3 and 4 yield a polymer-supported monolayer from which the tethered bilayer is obtained by a Langmuir\u2013Scha\u0308fer transfer of the distal lipid monolayer (method 3) or by vesicle fusion (method 4).", + "texts": [ + " An alternative approach to stabilisation of a polymer-supported lipid bilayer is based on controlled covalent tethering between the polymer cushion and the solid substrate and between the lipid bilayer and the polymer cushion. The tethering density can be controlled in a manner by adjusting the density of cross-linker molecules at the substrate through different selfassembly conditions and the density of actual tethering points through different reaction times. There are several simple but versatile techniques for covalently tethering polymers to solids (Fig. 9). The linkage is mediated by monolayers of alkyl silanes (for Si/SiO2 or indium-tin-oxide surfaces) or alkylmercaptanes (for gold and GaAs surfaces) carrying functional groups at the opposite ends, which can be covalently coupled to polymer chains. These functional groups can be: (1) epoxy groups, which are covalently bound to carboxyl groups; (2) amines, which bind covalently to carboxyl groups of chains (activated via coupling of succinimide or imidazole esters) and (3) photocross-linking groups (benzophenone silane-functionalised glass), which can be photochemically linked to any polymer segment" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002034_2007-01-2860-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002034_2007-01-2860-Figure1-1.png", + "caption": "Figure 1: Semi-Independent Suspension Twist-beam Type", + "texts": [ + "VI Congresso e Exposi\u00e7\u00e3o Internacionais da Tecnologia da Mobilidade S\u00e3o Paulo, Brasil 28 a 30 de novembro de 2007 AV. PAULISTA, 2073 - HORSA II - CJ. 1003 - CEP 01311-940 - S\u00c3O PAULO \u2013 SP FILIADA \u00c0 SAE TECHNICAL 2007-01-2860 PAPER SERIES E Twist Beam Rear Suspension - Influencies of the Cross Section Member Geometry in the Elastokinematics Behavior Vinicius Leal Fiat Automoveis S/A Rudinixon Moreira Bitencourt Comau do Brasil Janes Landre Junior Pontificia Universidade Cat\u00f3lica de Minas Gerais 2 torsional bar called crossbar, forming a typical \u201cT\u201d for this type of solution - Figure 1 illustrates the common geometry of twist-beam suspensions. The crossbar is the most important component for this type of suspension. Factors like area moment of inertia, polar moment of inertia and material used completely modify the suspension behavior. The crossbar profile is essential to control parameters such as suspension roll center, vertical reaction of the wheels and toe. At the same time, the crossbar has a critical structural aspect. Since it is subject to high rotation, which leads to great strain, fatigue dimensioning becomes a complex stage to define geometry, mainly in the connection with the trailing arm, where variation in profile section generally occurs" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001950_1-84628-179-2_5-Figure5.1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001950_1-84628-179-2_5-Figure5.1-1.png", + "caption": "Fig. 5.1. Helicopter model representation.", + "texts": [ + "4, both the Euler\u2013Lagrange and Newtonian model are presented and some control algorithms are proposed. Finally, Section 5.5 presents some simulations in order to validate the model and controller proposed. There exist different configurations of helicopters. The most popular are basically three: standard configuration, tandem rotor configuration and coaxial rotor configuration. We will develop the three configurations by using Newton laws. They were obtained in hover flight conditions. This model was taken basically from [36, 43]. Consider Figure 5.1. Denote I = {Ex, Ey, Ez} as a right-hand inertial frame, stationary with respect to the earth and let C = {E1, E2, E3} be a right-hand body fixed frame, where C is fixed on the position of the centre of mass (CG) of the helicopter. In the first case, take Ez in the direction downwards into the heart, and in the second case, E1 in the normal direction of helicopter flight, while E3 should correspond with Ez in hover flight conditions (see Figure 5.1). Let R : C \u2192 I be an orthogonal rotation matrix of type R \u2208 SO(3): R(\u03b7) = \u239b\u239d c\u03b8c\u03c8 s\u03c6s\u03b8c\u03c8 \u2212 c\u03c6s\u03c8 c\u03c6s\u03b8c\u03c8 + s\u03c6s\u03c8 c\u03b8s\u03c8 s\u03c6s\u03b8s\u03c8 + c\u03c6c\u03c8 c\u03c6s\u03b8s\u03c8 \u2212 s\u03c6c\u03c8 \u2212s\u03b8 s\u03c6c\u03b8 c\u03c6c\u03b8 \u239e\u23a0 (5.1) denoting the helicopter orientation with respect to I, and where \u03b7 = (\u03c8, \u03b8, \u03c6) describes the yaw, pitch and roll angles respectively. We have used the shorthand notation c\u03b2 = cos(\u03b2) and s\u03b2 = sin(\u03b2). The dynamic model was obtained by taking the following assumptions: 5.1 The blades of the two rotors are considered to hinge directly from the hub, that is, there is no hinge offset associated with rotor flapping. In this way, the coning angle is assumed to be zero, As a consequence each rotor will always lie in a disk termed the rotor disk. 5.2 The main rotor blades are assumed to rotate in an anti-clockwise direction when viewed from above and the tail rotor blades rotate in a clockwise direction, see Figure 5.1. 5.3 It is assumed that the cyclic lateral and longitudinal tilts of the main rotor disk are measurable and controllable. That is the flapping angles are used directly as control inputs. 5.4 The only air resistance modelled are simple drag forces opposing the rotation of the two rotors. 5.5 The aerodynamic forces generated by the relative wind are not considered. 5.6 The effects of operating the helicopter close to the ground are neglected. 5.7 The effects of the aerodynamic forces generated by the stabilizers are not taken into account", + "89) \u0393Q = |QA|E3 \u2212 |QD|E3 (5.90) where a = \u2212\u03c42 AD \u2212 l1A|TA| \u2212 l1D|TD| l3A|TA| + l3D|TD| (5.91) b = \u2212\u03c41 AD + l2A|TA| + l2D|TD| l3A|TA| + l3D|TD| (5.92) Another way to represent a dynamic model is by using the equations of motion of Euler\u2013Lagrange. Here, the model of [9, 43, 94] works. Define the generalized coordinates of the helicopter as q = (\u03be, \u03b7)T = (x, y, z, \u03c8, \u03b8, \u03c6)T \u2208 R 6 (5.93) where \u03be and \u03b7 represent the position and orientation of the helicopter with respect to the inertial-fixed frame respectively (see Figure 5.1). The translational and rotational kinetic energy of the helicopter are Ttrans = m 2 \u2329 \u03be\u0307, \u03be\u0307 \u232a = m 2 (x\u03072 + y\u03072 + z\u03072) (5.94) Trot = 1 2 \u2126T I\u2126 (5.95) The angular velocity in the body-fixed frame C is related to the generalized velocities (\u03c8\u0307, \u03b8\u0307, \u03c6\u0307) [59]: \u2126a = \u239b\u239d \u03c6\u0307 \u2212 \u03c8\u0307s\u03b8 \u03b8\u0307c\u03c6 + \u03c8\u0307c\u03b8s\u03c6 \u03c8\u0307c\u03b8c\u03c6 \u2212 \u03b8\u0307s\u03c6 \u239e\u23a0 (5.96) which can also be written as \u2126a = W\u03b7 \u03b7\u0307 (5.97) where W\u03b7 = \u239b\u239d \u2212s\u03b8 0 1 c\u03b8s\u03c6 c\u03c6 0 c\u03b8c\u03c6 \u2212s\u03c6 0 \u239e\u23a0 (5.98) Therefore \u03b7\u0307 = \u239b\u239d \u03c8\u0307 \u03b8\u0307 \u03c6\u0307 \u239e\u23a0 = W\u22121 \u03b7 \u2126a (5.99) The total kinetic energy of the system is given by T = Ttrans + Trot = 1 2 M \u2329 \u03be\u0307, \u03be\u0307 \u232a + 1 2 \u3008\u2126a, IC\u2126a\u3009 (5" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002979_978-3-540-70534-5_7-Figure9.9-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002979_978-3-540-70534-5_7-Figure9.9-1.png", + "caption": "Figure 9.9: Wheelchair simulation and CAD design [Woods 2006]", + "texts": [ + " Inverse kinematics The inverse kinematics is a matrix formula that specifies the required individual wheel speeds for given desired linear and angular velocity (vx, vy, \u03c9) and can be derived by inverting the matrix of the forward kinematics [Viboonchaicheep, Shimada, Kosaka 2003]. 9.4 Omni-Directional Robot Design We have so far developed three different Mecanum-based omni-directional robots, the demonstrator models Omni-1 (Figure 9.8, left), Omni-2 (Figure 9.8, right), and the full size omni-directional wheelchair (Figure 9.9). The first design, Omni-1, has the motor/wheel assembly tightly attached to the robot\u2019s chassis. Its Mecanum wheel design has rims that only leave a few millimeters clearance for the rollers. As a consequence, the robot can drive very well on hard surfaces, but it loses its omni-directional capabilities on softer surfaces like carpet. Here, the wheels will sink in a bit and the robot will then drive on the wheel rims, losing its capability to drive sideways. \u03b8\u00b7 FL \u03b8\u00b7 FR \u03b8\u00b7 BL \u03b8\u00b7 BR 1 2\u03c0r -------- 1 1 d e+( ) 2\u2044 1 1 d e+( ) 2\u2044 1 1 d e+( ) 2\u2044 1 1 d e+( ) 2\u2044 vx vy \u03c9 \u22c5= Omni-Directional Robot Design 153 The deficiencies of Omni-1 led to the development of Omni-2" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001967_978-1-4302-0088-8-Figure3-11-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001967_978-1-4302-0088-8-Figure3-11-1.png", + "caption": "Figure 3-11. The Lynxmotion Aluminum Arm", + "texts": [ + "java: A servo position data structure to assist in implementing the GroupMoveProtocol in the MiniSSC. \u2022 MiniSscGM: The implemented GroupMoveProtocol for the MiniSSC. In the next section, we\u2019ll discuss how to use the LM32 and the GroupMoveProtocol with a robotic arm. Moving your robot on the ground is just one type of motion. The second type is motion from a fixed position. To demonstrate this, I\u2019m going to use a robot arm. If you don\u2019t have a robot arm, you can purchase the components from Lynxmotion, Inc. at www.lynxmotion.com (see Figure 3-11) or make them yourself. I have included a class diagram of these classes in Figure 3-12. C H A P T E R 3 \u25a0 M O T I O N 91 The objective in this example is to create a simple model of a robot arm. The fields in this class are mostly static constants that will define the range of motion of its two axes: the shoulder and elbow. Of the remaining fields, ssc of type MiniSSC is the worker, and shoulderPos and elbowPos are in the class to maintain state. The constructor of the class takes the JSerialPort, and the move() method is just a passthrough to the MiniSSC" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002054_bfb0119386-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002054_bfb0119386-Figure1-1.png", + "caption": "Fig. 1: Modular concept of the fire fighting robot", + "texts": [ + " Due to the environmental conditions of the robot while heading to the scene of fire, the sensors must detect objects through smoke at high heat radiation. Currently there is no reasonable sensor system available on the market which fulfils these tasks sufficiently. 2 R e q u i r e m e n t s a n d b o u n d a r y c o n d i t i o n s 2.1 Requirements for the robot system Due to the experiences of the past with fire fighting robots, a new concept has been developed [1]. A small chain-driven remote-controlled vehicle is used as a tractor to pull a wheeled hose reels with an extinguishing monitor (fig. 1). The tractor is a tracked vehicle called MF4 and is manufactured by telerob (Germany) with the following technical data: dimensions: speed: total weight: umbilical: 1300.850-400 mm (L-W-H) 50 m/min 350,0 kg sensor data transfer control signals The robot will be used for fire fighting in plants or apartment houses. In that case the tractor drives into the building, navigates through bad visibility conditions caused by heavy smoke, pulls the reel nearby the fire, drops it and drives back to the starting point" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002690_robio.2007.4522277-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002690_robio.2007.4522277-Figure1-1.png", + "caption": "Fig. 1. Manipulating a 2D object by a single DOF robot finger with rigid tip in 2D plane", + "texts": [ + " The third is to show that this control signal is effective in case of manipulating even a thin and light rigid object by two coordinated robot fingers with soft hemispherical ends. First consider a simple problem of immobilization of rotational motion of a 2D rigid object pivoted at a fixed point Om by means of a single DOF planer finger robot (see Figs.1 and 2). It is assumed that the xy-plane in those figures is horizontal and therefore the effect of the gravity can be ignored. At first, consider the finger-object system of Fig.1, which has two independent position variables q1 and \u03b8, because \u03c6 is dependent on q1 and \u03b8 in such a way that \u03c6 = \u03c0 + \u03b8 \u2212 q1. Second, contact of the finger-tip with the 978-1-4244-1758-2/08/$25.00 \u00a9 2008 IEEE. 870 object surface can be expressed as a holonomic constraint of the form Q = r + l \u2212 {(xm \u2212 x0) cos \u03b8 \u2212 (ym \u2212 y0) sin \u03b8} = 0 (1) (see Fig.1). In addition, if there arose rolling between the finger-tip sphere and the object without slipping, then it would give rise to the equality \u2212r d dt \u03c6 = d dt Y (2) that means the zero relative velocity of the contact point between one along the finger sphere and another on the object surface. Equation (2) implies \u03b4Y + r\u03b4\u03c6 = 0, which decreases the DOF of the system. Thus, the overall DOF of the system becomes zero, that is, motion would stop and the finger-end should be stacked at the contact position" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003664_eacm.2010.163.2.83-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003664_eacm.2010.163.2.83-Figure1-1.png", + "caption": "Figure 1. Revolute joint schematisation", + "texts": [ + " In the following subsections, the internal loss coefficients expressions for mechanisms with revolute and contoured joints are first presented. These expressions will be incorporated in Engineering and Computational Mechanics 163 Issue EM2 Mechanical efficiency of a spur gear system Chaari et al. 83 the spur gear pair model in order to compute the whole mechanical efficiency of the gear system. To model the internal loss coefficient in bearings supporting the gear system, revolute joints between shafts and gears are used. Let us consider a revolute joint between links i and i + 1 as presented in Figure 1. It is assumed that the input power Pe is transmitted over the link 1 by a punctual force ~Fe having a linear velocity ~V e. The partial internal loss coefficient in the joint is expressed by i,i\u00fe1 \u00bc P f i,i\u00fe1 Pe 1 where Pf i,i\u00fe1 is the power loss owing to sliding friction. i,i\u00fe1 can then be expressed by i,i\u00fe1 \u00bc i,i\u00fe1 ~Ri,i\u00fe1 ~Fe ~V i,i\u00fe1 ~V e \u00bc i,i\u00fe1 ~Ri,i\u00fe1 ~Fe di=2\u00f0 \u00de\u00f8i,i\u00fe1 ~V e 2 with i,i\u00fe1 is the sliding friction coefficient, ~Ri,i\u00fe1 the reaction force between links i and i + 1, \u00f8i,i\u00fe1 is the angular speed between links i and i + 1 and di is the diameter of the revolute joint" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000254_(sici)1098-111x(200007)15:7<657::aid-int6>3.0.co;2-p-Figure4-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000254_(sici)1098-111x(200007)15:7<657::aid-int6>3.0.co;2-p-Figure4-1.png", + "caption": "Figure 4. Stability envelope of mode -to-mode trajectory.p q", + "texts": [ + "\u02d9 0 0 Define the error between the nominal trajectory and perturbed trajectory as e t s e t , a , x\u017d . \u017d .0 s x ta , x y x t , a , xU\u017d . \u017d .0 0 0 s x t y x* t 16\u017d . \u017d . \u017d . \u017d .Since x* t converges asymptotically from the equilibrium of mode to thep equilibrium of mode , then the sensitivity analysis of the mode -to-modeq p q controller involves examining how close the perturbed trajectory remains to the nominal trajectory when the system is subjected to small perturbations of plant \u017d .parameters. The stability envelope of x* t , shown in Figure 4, denotes the \u017d .region about x* t in which the perturbed trajectories must be constrained in order for the closed loop system to be considered stable and to have acceptable performance. The boundaries of the stability envelope satisfy the relationship x t s x* t q D x ) x* t ) x* t y D x s x t 17\u017d . \u017d . \u017d . \u017d . \u017d . \u017d .max min where D x is a positive vector. Therefore, the error between the nominal and perturbed trajectory satisfies the condition < 0 is the forward friction coefficient. Similarly, when the scales slide backwards over its environment the total friction is FBW = \u2212\u00b5BW v, where \u00b5BW > 0 is the backward friction coefficient. The friction in the transversal or lateral direction is Ft = \u2212\u00b5tv, where \u00b5t > 0 is the transversal friction 978-1-4244-2079-7/08/$25.00 \u00a92008 AACC. 862 coefficient. In the differential friction model, it is assumed that the friction coefficients satisfy the following ordering: \u00b5BW >> \u00b5t > \u00b5FW > 0, where \u00b5BW is much larger than the other friction coefficients. This implies that the backwards friction FBW is much larger when the scales slide backwards, which agrees with the geometry of the scales of the body. As a shorthand notation, we introduce a function \u00b5A(v), which describes the axial friction coefficient with the positive direction shown in Fig. 1. Thus, \u00b5a(v) = { \u00b5FW if v > 0; \u00b5BW if v < 0. (1) In this section we consider our first simple toy creature. Consider the flapper system in Fig. 2. Two (inflexible) rods are hinged at O with the scales orientations as shown. We assume that the instantaneous velocity of the flapper aOb is directed towards the left. The half-opening angle is \u03b8. Let \u03b8\u0307 = \u03c9 be the angular velocity of rod Oa. At a point P, which is a distance s away from the hinge O, the resulting linear velocity is \u03c9s. The combined velocity component in the axial and transversal direction of the section of the rod Oa at P is va = v cos \u03b8 and vt = \u03c9s + v sin \u03b8" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001293_rob.20099-Figure10-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001293_rob.20099-Figure10-1.png", + "caption": "Figure 10. a Long-range positioning of vehicle. b Initialization phase proceeding long-range motion.", + "texts": [ + " Let us assume that for the generalized 6-dof docking problem considered herein, the vehicle\u2019s shape is a cube with array type detectors on at least three of its orthogonal faces, Figure 9. Three spatial LOSs are sufficient to specify any 3D relocation of the vehicle, Table I. Two actions are carried out concurrently: While the galvanometer mirrors are in the process of align- ing the three LOS, the vehicle moves to its desired pose, Td xd ,yd ,zd , d , d , d , with respect to the world coordinate frame, Fw, but only achieves an actual pose defined by Ta xa ,ya ,za , a , a , a Figure 10 a . It is assumed that the LOS can be aligned in a significantly shorter amount of time than that required to move the vehicle itself . Ta and Td are defined as follows: Td = cos d cos d cos d sin d sin d \u2212 sin d cos d cos d sin d cos d + sin d sin d xd sin d cos d sin d sin d sin d + cos d cos d sin d sin d cos d \u2212 cos d sin d yd \u2212 sin d cos d sin d cos d cos d zd 0 0 0 1 1 and Ta = cos a cos a cos a sin a sin a \u2212 sin a cos a cos a sin a cos a + sin a sin a xa sin a cos a sin a sin a sin a + cos a cos a sin a sin a cos a \u2212 cos a sin a ya \u2212 sin a cos a sin a cos a cos a za 0 0 0 1 . 2 In order to estimate Ta, the equations of intersections between each LOS and its corresponding detector are obtained, in world coordinates, Fw: li Ta oDi = Tadi , 3 where li, i=1\u20133, represent the three LOS vectors; oDi are the initial Cartesian planes of the detectors, when the vehicle is in its initial pose e.g., before its long-range motion ; and di are the measured detector offsets defined as vectors with respect to the center of the vehicle at its actual pose, aFc Figure 10 b . The left-hand side of Eq. 1 is derived from the notion that the intersection between each of the three LOS vectors and the three detector vectors, Di, anywhere in space can be represented by, pi, a vector depicting the point of intersection with respect to the world coordinate frame: l1 D1 = p1, 4 l2 D2 = p2, 5 and l3 D3 = p3. 6 However, since the locations of the detectors, e.g., Di , are unknown in Fw, after the long-range positioning of the vehicle, the above equations can be expanded to incorporate Ta: l1 D1 = l1 Ta oD1 , 7 l2 D2 = l2 Ta oD2 , 8 l3 D3 = l3 Ta oD3 , 9 or, alternatively, pi can also be represented in terms of the offsets measured along the detectors: p1 = Tad1 , 10 p2 = Tad2 , 11 and p3 = Tad3 ", + " 15 Each equation above can be further expressed with respect to the x, y, and z coordinates of Fw. Three more equations are required in order to have a unique solution. To obtain these three additional equations, another movement transformation , M1, of the vehicle is introduced. M1 is referred to as a \u201cprobing\u201d motion, a simple translation that provides the necessary equations to solve for Ta. This probing motion also attempts to reduce the maximum of the three measured PSD offsets, emax e.g., e1, e2, or e3 , in each detector\u2019s frame, Fdi Figure 10 b . After M1 is applied to the vehicle, the equations of intersections between each LOS and corresponding detector in the world coordinates, Fw, are again solved for in the same manner as above: li M1Ta oDi = M1Tadi . 16 A nonlinear mathematical solver can then be used to estimate xa, ya, za, a, a, and a using the aforementioned relationships. Once Ta is estimated, the proposed guidance algorithm requires an initial corrective movement of the vehicle by M2 with respect to Fw in order to minimize the difference between Ta and Td, where M2 = Ta \u22121Td" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002489_elan.200703951-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002489_elan.200703951-Figure2-1.png", + "caption": "Fig. 2. Planar-chip sensor.", + "texts": [ + " 42, washed with cold deionized water several times to remove any impurities adsorbed on the surface of the precipitate, dried at room temperature for 24 h, and ground to a fine powder.Amembrane cocktail was prepared by mixing 2.1 mg portion of the precipitate with 124.4 mg of o-NPOE and 64.5 mg PVC. The mixture was dissolved in 3 mL THF and thoroughly mixed to obtain a homogeneous transparent mixture. Aplanar gold base electrode (3 5 mm)was sputtered on a (13.5 3.5 mm) flexible polyimide (Kapton, DuPont) substrate (125 mm thick), as shown in Figure 2; single site electrode (area\u00bc 0.06 cm2) (used for all the optimization and characterization studies), and used as previously described [27]. An electrical wire was connected to the electrode by means of Ag-epoxy (Epoxy Technology).Insulation of the electrical contact wasmade using silicon rubber coating seal (Dow Corning 3140 RTV). Themembrane cocktailmixturewas directly coated to the sputtered gold layer using microsyringe to apply few microliters of the sensing solution (Typically 10 mL of membrane cocktail is dispersed), left to dry in the air for oneminute before repeating further addition (i" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003995_powereng.2009.4915233-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003995_powereng.2009.4915233-Figure1-1.png", + "caption": "Fig. 1. Three-dimensional sketch of the AFDM.", + "texts": [ + " In general, AFDMs have an axial length much smaller than the length of a conventional machine of the same rating. Such a system is characterised by a great computational complexity, which should benefit by fast simulation methods, possibly based on conventional software. A number of 3D or quasi-3D, FEM and/or analytical models [1-11] and optimisation procedures [12-15] have been proposed in the literature. In this paper, a single-stator, twin rotors machine structure is considered. A sketch of the machine is shown in Fig. 1. This machine is taken as a reference to determine the equivalent circuit parameters. However, the described method can be easily applied to any other configuration. A cylindrical section of a three-phase machine, with a magnet for each rotor disc and three windings is shown in Fig. 2. The section is straightened to obtain a 2D scheme [16,17]. In a machine with two poles (p=1) there are six windings at a distance of 60 mechanical/electrical degrees, while in a machine with two pole pairs there are twelve windings at a distance of 30 mechanical degrees and 60 electrical degrees" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001605_robot.2005.1570127-Figure5-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001605_robot.2005.1570127-Figure5-1.png", + "caption": "Fig. 5. Graphical characterization of critical horizonal force: (a) fd = fgtan(\u03b2 1 ) and (b) fd = fgtan(\u03b2 2 ).", + "texts": [ + " Also, let fg be the gravitational force, which is the mechanism\u2019s constant weight. The external horizontal force, denoted fd, is variable. The total applied force is rotated by the angle \u03b2 = tan\u22121(fd/fg) about the vertical direction, where fg and fd are now scalars. Although the static contact reaction forces are indeterminate, the two possible critical cases of contact breakage or slippage are determinate and can be obtained graphically as follows. The first critical case occurs when the action line of the net external force intersects the contact point x2, as shown in Figure 5(a). In this case the contact reaction force at x1 vanishes, resulting in contact breakage at x1 and rolling about x2. The corresponding critical force angle is \u03b21 = tan\u22121(l/2h). The second critical case occurs when the action line of the net external force intersects the point puu , as shown in Figure 5(b). In this case the contact reaction forces lie on the edges of their respective friction cones, and sliding starts at both contacts. Using geometric relations, the corresponding critical force angle is \u03b22 = tan\u22121(sin2\u03b3/(2hsin2\u03b3/l\u2212 cos2\u03b1\u2212cos2\u03b3)), where \u03b3 = tan\u22121(\u00b5). The critical force fd for which the equilibrium conditions are violated, is given by fd = fgtan(\u03b2), where \u03b2 = min{\u03b21, \u03b22}. The experimental results are presented in Figure 6. For each force application height h, five experiments were conducted and the average critical force fd was measured with its corresponding angle \u03b2" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002115_s00170-007-1350-z-Figure8-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002115_s00170-007-1350-z-Figure8-1.png", + "caption": "Fig. 8 Experimental device (a: Rod, b: Laser, c: Shield gas).", + "texts": [ + " The rod length was chosen to be able to consider the rod as a semi-infinite solid. The steel rod was placed in a semi-closed chamber filled with inert gas in order to avoid the formation of oxides that could generate additional heat and could change the surface absorption coefficient. Moreover, a gas flux directly impinged on the top surface of the rod. A plate was used to hold the rod and to protect the cylindrical surface of the rod from the forced convection due to the gas (the cylindrical surface was modelled as adiabatic). In Fig. 8 a representation of the experimental setup is reported. A direct diode laser Rofin DL022 was used in the experiments. The laser is made up of stacks of separate diodes delivering radiation at wavelengths of 808\u00b110 nm and 940\u00b110 nm. For each power level, half power is delivered at 808 nm and half at 940 nm. The maximum nominal power of the laser was 2200 W. The spot adopted was rectangular and 4.8\u00d77.9 mm in size, which is the area of the laser spot was larger than the exposed surface of the rod. It should be noticed that the power distribution inside the spot of a diode laser is fairly flat, and the hypothesis of uniform heat input over the exposed surface is well verified" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001979_50037-5-Figure7.89-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001979_50037-5-Figure7.89-1.png", + "caption": "Figure 7.89 Cylinder with surface elements (174).", + "texts": [ + "88, the following selections are made: [A] Elements (first pull down menu); [B] By Elem Name (second pull down menu); and [C] Element Name = 174. The element with the number 174 was introduced automatically during the process of creation of contact pairs described earlier. It is listed in the Preprocessor \u2192 Element Type \u2192 Add/Edit/Delete option. Selections are implemented by pressing [D] OK button. From Utility Menu select Plot \u2192 Elements. Image of the cylinder with mesh of elements is produced (see Figure 7.89). It is seen that the gap equal to 0.05 units exists between two half of the cylinder. It is the result of moving half of the cylinder toward the block (by 0.05 cm) in order to create loading at the interface. From ANSYS Main Menu select General Postproc \u2192 Plot Results \u2192 Contour Plot \u2192 Nodal Solu. The frame shown in Figure 7.90 appears. In the frame shown in Figure 7.90, the following selections are made: [A] Contact and [B] Pressure. These are items to be contoured. Pressing [C] OK implements selections made" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000881_robot.2004.1308085-Figure7-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000881_robot.2004.1308085-Figure7-1.png", + "caption": "Fig. 7. Planar plots of contact point time history during sideways locomotion produced by Gait 1 in (a) RRRobot-on-a-plane simulation, (b) RRRobot-on-a-plane experiment, (c) Pivoting Dynamics model simulation. The solid arrow gives robot motion direction, and the dotted lines indicate the robot position at the specified time.", + "texts": [], + "surrounding_texts": [ + "Home position Displaced position\nA planar eccentricmass wheel performs harmonic oscilla-\nC. Single Axis Rotation Models If we consider body attitude changes only about one axis, say, the X or^ Y axis, then RRRobot on a plane is similar to a planar wheel with an.eccen*c mass (see Fig. 5). The 1ocatio.n of the mass and the inertia of the system is determined by the weight distribution on the robot. If T is the wheel radius, &I is the lumped mass of the system, and p is the radius of gyration of the system with respect to an axis passing through the contact point and perpendicular to the plane, the time-period for small amplitudes~ is\nwhere g is gravity. Note that T, decreases as T increases, and T, increases as p increases.\nIf the Pivoting Dynamics model is restricted to oscillate about the X or Y axis, then the Pivoting Dynamics model is similar to a simple pendulum (see Fig. 6). whose time-period is Tap = 2i8, where p is the radius of gyration, and g is gravity. The time-period Tsp decreases as p decreases.\nNote that to get similar oscillatory behavior between the eccentric mass wheel and the simple pendulum, a rearrangement of masses may he required. Table I shows the time-periods for X and Y rotations for the RRRoboton-a-plane model and the Pivoting Dynamics model. Since the time periods of the two models are close to each other, we do not rearrange the masses.\nMODEL A N D THE PIVOTING DYNAMICS MODEL\nI X Rotations (sec) I Y Rotations (sec) RRRobot-on-a-olane I 1.29 I 1 .w", + "Y, Sideways (m)\nX, Forward (m)\nFig. S. Planar plots of contact point time history during forwards Iocomotion produced by Gait 2 in (a) RRRobot-on-a-plane simulation, (b) RRRobot-on-a-plane experiment, (c) Pivoting Dynamics model simulation. The solid mow gives robot motion direction, and the dotted lines indicate the robot position at the specified time.\nthe horizontal. Gait 2 does not produce much translation, because the XZ oscillations are small and surface stickiness restricts motion. Our experience indicates that this gait is the least reliable of the gaits explored in this paper.\nGait 3 Gait 3 produces counter-clockwise circular translation (see Fig. 9) due to a combination of XYZ body attitude oscillations. The robot completes a circle in the RRRobot-on-a-plane simulation, completes one and a half circles in the Pivoting Dynamics Model simulation, and almost completes a half circle in experiment.\nNote that in all three gaits, swapping the relative phase between the two legs produces translation in the opposite direction. The paths followed by the contact point in simulation and experiment match well, but there is one clear difference- the robot in experiment moves slower than in simulation. This may be due to unmodelled surface friction, slip between the body and the surface, or a deformed spherical shape at the contact point.\nThe translation produced in the Pivoting Dynamics model and in the RRRobot-on-a-plane model match", + "well; the contact point follows similar paths, but the Pivoting Dynamics model moves faster, especially for. Gaits 1 and 3. This is because the Pivoting Dynamics Model is pivoted at its geometric center, while in the RRRobot-on-a-plane Model, the robot has a rolling contact. Thus, for a given change in attitude, the point of contact moves faster in the Pivoting Dynamics model .than in ,the RRRobot-in-a-plane Model. In summary, we can use the Pivoting Dynamics model to approximate RRRobot planar translation.\nIV. CONCLUSION\nWe explored locomotion for a high-centered roundbodied legged robot, the RRRobot, using experiments and simulation. We presented sinusoidal leg trajectories that produce forward, sideways, and rotational translation and explored simplified models to understand the\nlocomotion. Future work will include using more legs to perform richer motion, understanding the influence of body shape on the translation, gait search, and kinematic reduction of RRRobot dynamics.\nACKNOWLEDGMENT This work was supported under NSF IIS 0082339, NSF IIS 0222875, and DAWNONR NW014-98-1-0747 contracts. Devin Balkcom gave insightful comments on the dynamics and nonholonomy of mechanical systems, Brendan Meeder helped build the robot and tracking code, and Siddhaxtha Srinivasa gave useful comments on the paper.\nREFERENCES [I I R. Balawhramanian, A. A. Rizzi, and M. T. Mason. Legless\nlocomotion for legged robots. In Proceedings of rlte Inremarional Conferrnce on Rohrs and lnrelligenr Systems, volume 1. pages 88&885. 2003. 121 A. Bicchi, A. Balluchi, D. Pranichino, and A. Gorelli. Introducing the sphericle: an experimental testhed for research and teaching in nonholonomy. In Pmceedings of the IEEE lriremarional Conference on Roborics and Auromarion. Dazes . _ 262&2625, 1997.. 131 A. Bloch. J. Baillieul. P.Crouch. and I. Marsden. Nonholonomic . . Mechnnics and Conrml. Springer, 2003. 141 F. Bullo, A. D. Lewis, and K. M. Lynch. Controllable kinematic reductions for mechanical systems: concepts, computational tools. and examples. In Mathematical Theory of Nerworkr and Sjsterns. aug 2002. 151 C. Camicia. F. Conticelli, and A. Bicchi. Nonholonomic kinematics and dynamics of the sphericle. In Pmceedings of r11e IEEE lnternarional Conference on lnrelliqenr Robors and Sjsrems, pages 805-810, 2ooO. 1. 1. Craig. Inrmduerion ro Roborics. Addison Wesley. 1989. C. Ferandes, L. Gurvits, and Z. Li. Near optimal nonholonomic motion planning for a system of coupled rigid bodies. IEEE Transactions on Auromric Conrml, 39(3), 1994. A. Lewis, 1. Ostrowski, R. Murray, and 1. Burdick. Nonholonomic mechanics and locomotion: The snakeboard example. In Pmceedings of the lnremorional Conference on Roborics and Aurornarion, volume 3, pages 2391-2397, 1994. 2. Li and 1. Canny. Motion of two rigid bodies with rolling constraint. IEEE Transactions on Rohorics and Auromarion, 6(1):62-72, Feb. 1990. D. I. Montana. The kinematics of contact and grasp. m e lnremarional Journal of Robotics Research, 7(3):17-32, June 1988. R. M. Murray, 2. X. Li, and S. S. Sastry. A Marhemrical lnrmducrion ro Rohric Manipularion. CRC Press, 1994. J. P. Ostrowski. The Mechanics and Conrml of Undularory Rohoric Locornorion. PhD thesis, California Institute of Technology, 1996. C. Rui, 1. V. Kolmanovsky, and N. H. McClamroch. Nonlinear attitude and shape control of spacecrafl with articulated appendages and reaction wheels. IEEE Tmnsacrions on Auromric Conrml, 45(8):145549, Aug. 2000. D. Zenkov, A. Bloch, and 1. Marsden. The energy-momentum method for the stability of nonholonomic systems. Technical repon, California Institute of Technology, 1997." + ] + }, + { + "image_filename": "designv11_69_0003808_jrc2009-63011-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003808_jrc2009-63011-Figure1-1.png", + "caption": "Figure 1. Finite element model of a 36\u201d freight car wheel with a rim thickness of 1.5 in (38.1 mm).", + "texts": [ + " The authors are currently working on estimating the condemning rim thickness limit and the allowable defect size to prevent shattered rim cracking considering residual stress as initial stress. In this paper, the residual stresses that develop during both the manufacturing process and thermal brake loading under service conditions are estimated using three-dimensional decoupled thermal-structural finite element analyses. Two sets of analyses are performed. In the first set, the manufacturing process is simulated and in the second set, the thermal brake loading under service conditions is simulated. Figure 1 shows a three-dimensional finite element model of a 36\u201d freight car wheel with a rim thickness of 1.5 in (38.1 mm) built in ANSYS [12]. The wheel profile is chosen according to AAR standards [13]. The finite element model is meshed using SOLID70 elements, which have 8 nodes with temperature as the only degree of freedom, for thermal analysis, and using SOLID185 elements, which have 8 nodes with three translational degrees of freedom, for structural analysis. The finite element model contains 7584 elements and 11041 nodes", + " To determine the converged mesh density, initial analysis is performed to estimate the as-manufactured residual hoop stresses with different mesh densities. The residual hoop stress values computed on the taping line are compared for all the mesh densities. The mesh density which has the residual hoop stress value very close (less than 5%) to that of the immediate finer mesh is chosen as the converged mesh and is used for all the analyses performed in this paper. This converged mesh is shown in Figure 1. However, the mesh shown in Figure 1 is not acceptable for modeling shattered rim cracking. To model the shattered rim cracking much finer mesh is needed. In the literature, Liu et al. [16] have considered much finer mesh in the rim portion compared to the other regions in order to accurately estimate the contact stresses. Shattered rim Copyright \u00a9 2009 by ASME ms of Use: http://www.asme.org/about-asme/terms-of-use Dow cracking considering residual stresses can be modeled through two stages. In the first stage, rolling contact loading can be simulated using a full model of rail and wheel without considering the shattered rim crack. The mesh shown in Figure 1 is acceptable for full model analysis. In the second stage, a sub model can be built, including a shattered rim crack below the tread surface, focusing on the region close to the contact area. The rolling contact loading for the sub model can be simulated by applying cut boundary displacement boundary conditions obtained from the full model. The mesh density shown in Figure 1 is not acceptable for the sub model; therefore, much finer mesh needs to be considered for the sub model. The wheel manufacturing process is simulated using two steps: non-linear transient thermal analysis and non-linear elastic-plastic structural analysis including creep effects. The thermal and structural analyses are performed using the same finite element model meshed with different types of elements. The analyses considered temperature dependent material properties. In the non-linear transient thermal analysis, various steps of the manufacturing process are simulated using different convection boundary conditions on the wheel surfaces" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000710_robot.2004.1307395-Figure4-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000710_robot.2004.1307395-Figure4-1.png", + "caption": "Fig. 4. The snakeboard indel", + "texts": [ + " SNAKEBOARD: INVERSE KINEMATICS, COMPLETENESS We apply the ideas of the previous section to the snakeboard model. We show that the switch-optimal inverse kinematics procedure satisfies the convergence property, and therefore \u2018Here a neighborhood is in he seme of the standard topology. e.g.. thal induced by the Euclidean metric. the motion planner for the snakeboard among obstacles is complete. A. The Snakeboard Model FIrst, we briefly describe a model of the snakeboard taken from [6], (171 based on the original work [24]. As shown in Fig. 4, the snakeboard has a five-dimensional C-space described by the coordinate vector q = (x:y)t\u2019,$>+) E Q = SE@) x S\u2019x [-$,:I. Specifically, (z,y) represents the Cartesian position of the center of the snakeboard body, B is its angle, and 1c. and 4 are the angle of the momentum rotor and the steering angle of the wheels, respectively, expressed in the body frame. The two inputs are the torque to spin the steering wheels and the torque to spin the rotor. As the rotor spins with the steering wheels fixed, by conservation of angular momentum about the rotation center chosen by the wheels and their noslip constraints, the snakeboard body rotates in the opposite direction" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002979_978-3-540-70534-5_7-Figure4.1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002979_978-3-540-70534-5_7-Figure4.1-1.png", + "caption": "Figure 4.1: Motor\u2013encoder combination", + "texts": [ + "1 DC Motors Electrical motors can be: AC motors DC motors Stepper motors Servos DC electric motors are arguably the most commonly used method for locomotion in mobile robots. DC motors are clean, quiet, and can produce sufficient power for a variety of tasks. They are much easier to control than pneumatic actuators, which are mainly used if very high torques are required and umbilical cords for external pressure pumps are available \u2013 so usually not an option for mobile robots. Standard DC motors revolve freely, unlike for example stepper motors (see Section 4.4). Motor control therefore requires a feedback mechanism using shaft encoders (see Figure 4.1 and Section 3.4). The first step when building robot hardware is to select the appropriate motor system. The best choice is an encapsulated motor combination comprising a: Actuators 74 4 \u2022 DC motor \u2022 Gearbox \u2022 Optical or magnetic encoder (dual phase-shifted encoders for detection of speed and direction) Using encapsulated motor systems has the advantage that the solution is much smaller than that using separate modules, plus the system is dust-proof and shielded against stray light (required for optical encoders)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001793_gt2006-90280-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001793_gt2006-90280-Figure2-1.png", + "caption": "Fig. 2 Coordinate systems", + "texts": [ + " This paper analyzes the variations of the stiffness and damping coefficients for the tilting pad journal bearings with the frequency of excitation and describes the analytical and experimental techniques used to evaluate these properties. Ax, Ay Fourier transforms of accelerations in the horizontal and vertical directions, respectively Fx, Fy Fourier transforms of excitation forces in the horizontal and vertical directions, respectively Ip mass moment of inertia of pad about pivot L bearing width N number of data points T sampling time X, Y Fourier transforms of displacements in the horizontal and vertical directions, respectively cxx, cxy, cyx, cyy bearing damping coefficients in x-y coordinate system (see Figure 2a) )()()()( ,,, eqeqeqeq cccc \u03b7\u03b7\u03b7\u03be\u03be\u03b7\u03be\u03be equivalent pad damping coefficients in \u03be-\u03b7 coordinate system (see Figure 2b) )()()()( ,,, iiii cccc \u03b7\u03b7\u03b7\u03be\u03be\u03b7\u03be\u03be fixed pad damping coefficients in \u03be-\u03b7 coordinate system (see Figure 2b) )()()()( ,,, itititit cccc \u03b7\u03b7\u03b7\u03be\u03be\u03b7\u03be\u03be tilting pad damping coefficients in \u03be-\u03b7 coordinate system (see Figure 2b) d bearing nominal diameter fd,x,, fd,y components of the dynamic force (excitation) kxx, kxy, kyx, kyy bearing stiffness coefficients in x-y coordinate system (see Figure 2a) )()()()( ,,, eqeqeqeq kkkk \u03b7\u03b7\u03b7\u03be\u03be\u03b7\u03be\u03be equivalent pad damping coefficients in \u03be-\u03b7 coordinate system (see Figure 2b) )()()()( ,,, iiii kkkk \u03b7\u03b7\u03b7\u03be\u03be\u03b7\u03be\u03be fixed pad stiffness coefficients in \u03be-\u03b7 coordinate system (see Figure 2b) )()()()( ,,, itititit kkkk \u03b7\u03b7\u03b7\u03be\u03be\u03b7\u03be\u03be tilting pad stiffness coefficients in \u03be-\u03b7 coordinate system (see Figure 2b) mb bearing mass mp mass of pad mz equivalent pad mass, /r I m ppz = nd number of records rp distance from pivot to pad centre x,y shaft center coordinates in the rectangular system with the origin at the bearing equilibrium position t\u2206 time increment for sampling data \u03c9exc frequency \u03c9 angular speed of the shaft EXPERIMENTAL INVESTIGATION The NRC\u2019s test rig, which is shown in Figure 1, utilizes the concept of a fixed rotating shaft (1) and a free vibrating test 2 ownloaded From: http://proceedings", + " Each shaker is attached to the bearing housing through a long steel rod (5). The shaft is supported on high precision, angular ball bearings. A tensioned cable (6) applies a static load. Soft springs (7) minimize the effect of bearing vibration on the applied static load. The shakers have been programmed to provide a multifrequency excitation, which consisted of frequencies ranging between 20 and 500 Hz with an increment of 10 Hz. In the presence of external excitation, equations of bearing motion with respect to the shaft are as follows (Figure 2a illustrates the coordinate system): (1) ydyyyxyyyxb ,( ) ( ) ( ) ( ) ( ) ( )tftyctxctyktxktym =++++ &&&& ( ) ( ) ( ) ( ) ( ) ( )tftyctxctyktxktxm xdxyxxxyxxb ,=++++ &&&& a. attached to shaft centre b. attached to pad centre \u00a9 National Research Council of Canada rl=/data/conferences/gt2006/71120/ on 02/13/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use D Table 1 Rig specifications Shaft speed 16,500 rpm Journal diameter 0.09843 m (3.875 in) Static load 20 kN (4,500 lbf) Lubricant flow 0", + " The model also calculates both the thermal and elastic distortions of the individual pads. P iv ot s tif fn es s, N /m The above hydrodynamic considerations allow for calculation of the stiffness and damping coefficients for fixed pads. They have been used to calculate the coefficients for individual tilting pads applying the technique described by Lund [15]. This technique takes into considerations the mass and the excitation frequency of the tilting pad. For example, in the coordinate system associated with the pad (Figure 2b), the direct coefficients of stiffness and damping in radial direction for the \u201ci\u201d pad can be written as (6) ( )( ) ( ) + \u2212\u2212+ += )()()()()(2 )()()()(22)()( )( )()( 1 iiiii exc iiii excexc i z i i iit ckckc kkccmk A kk \u03be\u03b7\u03b7\u03be\u03b7\u03be\u03be\u03b7\u03b7\u03b7 \u03b7\u03be\u03be\u03b7\u03b7\u03be\u03be\u03b7\u03b7\u03b7 \u03b7 \u03be\u03be\u03be\u03be \u03c9 \u03c9\u03c9 (7) c ( )( ) ( ) \u2212 +++ \u2212= )()()()(2)( )()()()(2)()( )( )()( 1 iiii exc i iiii exc i z i i iit kkccc ckckmk A c \u03b7\u03be\u03be\u03b7\u03b7\u03be\u03be\u03b7\u03b7\u03b7 \u03be\u03b7\u03b7\u03be\u03b7\u03be\u03be\u03b7\u03b7\u03b7 \u03b7 \u03be\u03be\u03be\u03be \u03c9 \u03c9 where ( ) )(222)()()( i excexc i z ii cmkA \u03b7\u03b7\u03b7\u03b7\u03b7 \u03c9\u03c9 ++= (8) )( )( )( i p i pi z r I m = (9) Relationships for the remaining six coefficients are given in Appendix A" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002690_robio.2007.4522277-Figure3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002690_robio.2007.4522277-Figure3-1.png", + "caption": "Fig. 3. Manipulating a 2D object by a two DOF robot finger with soft tip in 2D plane", + "texts": [ + " Equation (3) is no more a constraint, because it can be regarded that the reproducing force F of deformation arises in the normal direction to the object surface as a nonlinear function of \u0394x together with viscous-like force. Therefore, even if rolling of contact is taken into account, the net DOF of the system of Fig.2 becomes at least of one, that enables to stabilize rotational motion of the system at a state of force/torque balance. The details are discussed in the following by illustrating a more general case of use of a 2 DOF finger shown in Fig.3. III. IMMOBILIZATION BY A ROBOT FINGER WITH SOFT TIP The kinetic energy of the system of Fig.3 is expressed as K = 1 2 q\u0307TH(q)q\u0307 + 1 2 I\u03b8\u03072 (4) where q = (q1, q2)T, H(q) stands for the inertia matrix of the robot finger, I the inertia moment of the object around the z-axis at the fixed point Om. Obviously it follows from geometrical meanings of symbols specified in Fig.3 that \u0394x = r + l \u2212 (xm \u2212 x0) T rX (5) x1 = x0 + (r \u2212\u0394x) rY (6) xm = x1 + lrX \u2212 Y rY (7) Y = (x0 \u2212 xm)T rY (8) where rX = ( cos \u03b8 \u2212 sin \u03b8 ) , rY = ( sin \u03b8 cos \u03b8 ) (9) We assume that rolling between the finger-tip and the object through movement of the contact area can be expressed as the zero relative velocities of the center of contact area in the polar coordinates of the finger-end and the object coordinates in such a way that (r \u2212\u0394x) d dt \u03c6 = \u2212 d dt Y (10) where \u03c6 = \u03c0 + \u03b8 \u2212 q1 \u2212 q2 = \u03c0 + \u03b8 \u2212 qTe (11) and e = (1, 1)T and \u03b8 denotes the angle of inclination of the object to the x-axis", + " In other words, equation (28) shows that force/torque balance is achieved. Therefore, by the simulation, we confirmed that stable grasping in a dynamic sense was realized by using a single DOF robot finger with hemispherical soft end. TABLE II PARAMETERS OF CONTROL SIGNALS fd internal force 0.250[N] c Damping Coefficient 0.001[msN] \u03b30 regressor gain 0.001 N\u03020(0) initial estimate value 0.0 y x F f O =(x ,y)m m m P =(x ,y)1 1 Y l Fig. 5. Robot finger manipulating an object with parallel flat surfaces in 2D plane Even in the case of use of a 2-DOF finger shown in Fig.3 that has joint redundancy for the sake of only immobilization, it is confirmed that force/torque balance is realized as seen in Fig.6 (time-scale of the graphs differs from Fig.4) within 0.5 seconds. Physical parameters of the finger and object and control gains of this simulation are shown in Tables III and IV respectively. The noteworthy difference between the non-redundant case of Fig.4 and redundant case of Fig.6 is that the convergence speed of the non-redundant case is faster than the redundant one" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001384_j.ijmachtools.2006.02.011-Figure3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001384_j.ijmachtools.2006.02.011-Figure3-1.png", + "caption": "Fig. 3. Components of hot tool.", + "texts": [ + " The thermoplastic polymer material satisfies low decomposing temperature and viscosity property. The melting temperature of thermoplastic polymer is below 300 1C and the decomposition temperature is approximately 450 1C. For the RHA process, expandable polystyrene (EPS) foam widely used in industry is selected as a material. In addition, EPS foam part can be applied using evaporable pattern casting (EPC) process to convert into the metal part. Table 2 shows thermal properties of EPS foam from Ref. [9]. The hot tool used in the present work, as shown in Fig. 3, consists of a zig, a holder, an insulator, a cartridge heater, and a tool with tangential grooves. (1) Zig: The zig is composed of Vesfels (Dupont), which has good characteristics, such as heat-resistance and heatinsulation. The zig plays an important role in isolating heat flowed from the cartridge heater to protect the apparatus from heat. (2) Holder: A holder with heat-resistance connects the zig to the insulation. The chosen material for the holder, SUS has good stiffness and heat resistance" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001285_cacsd-cca-isic.2006.4776982-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001285_cacsd-cca-isic.2006.4776982-Figure2-1.png", + "caption": "Fig. 2. Flight path", + "texts": [ + " This papers is devoted to describe the first phase in the developing of a tailsitter convertible vehicle, which is the attitude control in vertical mode. The proposed configuration is a trade-off between a rotary-wing aircraft and a fixed-wing aircraft, offering the manoeuvrability of the helicopter and the flight endurance of the airplane [see figure 1]. The main assignment for this vehicle is to perform a vertical take-off and hovering in vertical mode, afterwards it switches from vertical to horizontal mode in order to perform forward flight and finally, the aircraft lands vertically [see figure 2]. There are a few publications concerning this kind of vehicles. In [3] is presented a convertible UAV named Heudiasyc-UTC UMR 6599 Centre de Recherches de Royallieu B.P. 20529 60205 Compiegne France Tel.: + 33 (0)3 44 23 44 23 ; fax: +33 (0)3 44 23 44 77 Corresponding author jescareno@hds.utc.fr rlozano@hds.utc.fr sergio@hds.utc.fr Twing, which is a twin-engine tailsitter UAV that uses a LQR algorithm applied to the linearized hover dynamics. The paper is presented as follows: Section II presents the attitude model of the birotor" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003424_s12206-009-1173-y-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003424_s12206-009-1173-y-Figure1-1.png", + "caption": "Fig. 1. Experimental apparatus.", + "texts": [ + " Further, we hypothesize that piston rings can project and get caught in the ports (except piston ring gaps), causing further scuffing. As far as we can tell, no experiments have reported on the measurement of the extent to which piston rings project and get caught in the ports. In this study, we installed strain gauges, on the bottom sides of piston rings, over the intake and exhaust ports and plotted the variation in strain per cycle, while running the engine. Examining the variation in strain per cycle, we investigated whether piston rings do in fact project and get caught in the ports. Fig. 1 shows our experimental apparatus. We used a twostroke air-cooled single-cylinder gasoline engine with a bore of 62 mm and a stroke of 58 mm. In this engine, the cylinder has an intake port on the thrust side, an exhaust port on the anti-thrust side, and scavenging ports on both the front and rear sides. In this piston, we installed two rings: a barrel-faced half keystone top ring on top and a taper\u2013faced rectangular second ring below it. Both rings had a width of 2.0 mm, a thickness of 2.8 mm, and a tension of 11N" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001245_j.memsci.2005.10.019-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001245_j.memsci.2005.10.019-Figure2-1.png", + "caption": "Fig. 2. (a) Experimental device. (b) Membrane module.", + "texts": [ + " Enzymes We used two kinds of enzymes: a solid enzyme made up of 1,4-beta-xylanase from Sigma\u2013Aldrich (2500 U/g) and a liquid solution made up of a mixture of enzymes (including arabanase, cellulase, -glucanase, hemi-cellulase and xylanase) from Sigma\u2013Aldrich. To obtain the complex with the liquid enzyme, solutions containing the activated carbon or the activated carbon\u2013metal system, and the enzyme solution were agitated for controlled periods. 2.5. Experimental device Enzymatic membrane reactors were tested in an experimental system containing a pump piston, a surge suppressor, a back-pressure controller (to keep the pressure constant) and a circular flat membrane module with an effective membrane area of 15 cm2. The pressure was fixed at 9 bars. Fig. 2 shows the experimental device (Fig. 2a) and the membrane module (Fig. 2b). Two different oligosaccharides solutions were tested. A real sample mixture of oligosaccharides obtained in the laboratory by acid hydrolysis from nutshells for the EMR containing the s o E t ( a w a t f T R C M P P D o lignin of P/L = 0.7\u20131.75 [9]. Surface area and pore size charcterization were performed using a Micromeritics ASAP2020 as adsorption surface area analyzer. The specific surface area f the samples was determined from the nitrogen isotherms at 196 \u25e6C and the BET equation. Micropore volume was deterined from the t-plot, mesopore volume from the BJH equation nd total volume of pores was calculated with a relative pressure p/p0) of 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001979_50037-5-Figure7.114-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001979_50037-5-Figure7.114-1.png", + "caption": "Figure 7.114 Model after meshing process.", + "texts": [ + "109) click [A] Set in the Size Controls: Lines option and pick two lines located on top surface of the rail: one coinciding with the line previously picked and the other at the right angle to the first one. Click [A] OK as shown in Figure 7.110. The frame shown in Figure 7.111 appears again. In the box [A] No. of element divisions type 30 and press [C] OK button. In the frame MeshTool (see Figure 7.112) pull down [A] Volumes in the option Mesh. Check [B] Hex and [C] Sweep options. Pressing [D] Sweep button brings another frame, as shown in Figure 7.113, asking to pick volumes to be swept. Pressing [A] Pick All button initiates meshing process. The model after meshing looks like the image in Figure 7.114. Pressing Close button on MeshTool frame, ends mesh generation stage. After meshing being completed, it is usually necessary to smooth element edges in order to improve graphic display. It can be accomplished using PlotCtrls facility in the Utility Menu. From Utility Menu select PlotCtrls \u2192 Style \u2192 Size and Shape. The frame shown in Figure 7.115 appears. In the option [A] Facets/element edge select 2 facets/edge and click [B] OK button as shown in Figure 7.115. From ANSYS Main Menu select Preprocessor \u2192 Modelling \u2192 Create \u2192 Contact Pair" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001500_sice.2006.315273-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001500_sice.2006.315273-Figure2-1.png", + "caption": "Fig. 2 An inverted cart pendulum model", + "texts": [], + "surrounding_texts": [ + "The inverted pendulum model which is added by the Coulomb friction after linearized at the point of x = 0, 0=0 is E-44j brg (IH i I 0 and u> 0, the time friction estimation error become negative. time derivative means the estimation error o approach zero as time goes. That is, esti friction converges to real value. Now we can compensate the Coulomb f the estimate from this friction observer. Fr section, we construct a stabilizing contr inverted cart pendulum with the Coulomb state feedback controller added by the est friction is given by u -Kx + 4. EXPERIMENT AND RESUThe experimental apparatus to confirm i of tracking performance by the estimat( friction is depicted in Fig. 4. The inverte system in experimental setup is manu Quanser. The Coulomb friction considered (17). iear friction we define an ation of the lin Eq. (16) ,er nynamics. ie estimation The parameters of the inverted cart pendulum used in the experiment are listed in Table 1. The state feedback gain K to stabilize inverted pendulum is determined to satisfy a performance index of the cart position and pole angle. K = [-50 - 50.26 180.163 28.49] (20) Using the Coulomb friction compensator Eqs. (15) - gnOFW2IL7)ySTO.0,~(17) and Eq. (20), the influence of the Coulomb friction is removed efficiently. The results of experiments are sufficient to show the validity of proposed algorithm that is a method to eliminate effects of the Coulomb friction. Note that the convergence rate of estimation changes (18) due to the value of the observer gain k and exponent derivative of p from Eq. (18). Fast convergence makes the The negative performance of the friction observer too sensitive to )f the friction noise and disturbance in the process. In contrast, too imate of the slow convergence influences on the tracking performance of the controller. Therefore we examine riction using the improvement of the tracking performance for om previous various values of k and ,u. The effects of k is less oller for an than p on the performance of compensation of the friction. The Coulomb friction in the system. And as we can see in imate of the the equation of the observer for the Coulomb friction, a large p makes the convergence slow. Because of this (19). reason, large pu doesn't estimate the friction quickly. LTS Therefore the performance of the elimination friction effect on the system gets bad. In contrast to improvement previous example, a small p can estimate the ed Coulomb Coulomb friction fast. And then the oscillation in the d pendulum response of the cart position and pole angle is improved factured by very well. Without the estimation and compensation of in this paper the Coulomb friction, the amplitude of the oscillation of F =as01 the cart position and pole angle are 0.022[m] and 0.6\u00b0 respectively with only the state feedback controller. But with the friction estimator of ,u = 1, k =20, the oscillations decrease below 0.004[m] and 0.3\u00b0 respectively. Another important fact is that the pole is staying at 0\u00b0 and the cart is O[m] after the friction is compensated. This is meaning that the effect of the friction on the system response is eliminated effectively. And the results of tracking control are also depicted in Fig. 6. In the contrast to the case without compensation of the Coulomb friction, the controller with an estimation and compensation of the friction shows better tracking performance about given reference input. (reference input r +0.02 [m])" + ] + }, + { + "image_filename": "designv11_69_0003973_00405000802184862-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003973_00405000802184862-Figure1-1.png", + "caption": "Figure 1. Actual versus predicted fibre length (50:50).", + "texts": [], + "surrounding_texts": [ + "Tables 2\u20137 show the predicted values of length, strength, elongation and micronaire values of blends. The predicted Table 2. The HVI results of mixed cottons (25:75). 25:75 Predicted Tested Actual Difference Predicted Tested Actual Difference S. no. Cotton UHML (mm) UHML (mm) difference (%) micronaire micronaire difference (%) 1. MCU-5 + BUNNY 30.100 30.730 0.630 2.050 3.423 3.650 0.228 6.233 2. DCH-32 + MCU-5 31.667 31.310 \u22120.357 \u22121.140 3.892 3.990 0.098 2.456 3. DCH-32 + BUNNY 31.187 30.870 \u22120.317 \u22121.027 3.105 3.290 0.185 5.623 4. US PIMA + S-6 30.850 31.530 0.680 2.157 3.582 3.870 0.288 7.442 5. MECH-1 + US PIMA 33.897 34.360 0.463 1.347 3.747 3.650 \u22120.097 \u22122.658 6. MECH-1 + BRAHMA 30.612 29.740 \u22120.872 \u22122.932 3.748 3.630 \u22120.118 \u22123.237 7. GIZA-86 + DCH-32 34.045 35.740 1.695 4.743 3.290 3.720 0.430 11.559 8. GIZA-86 + BUNNY 30.302 30.820 0.518 1.681 3.455 3.440 \u22120.015 \u22120.436 9. GIZA-86 + S-6 29.890 27.860 \u22122.030 \u22127.286 3.785 3.770 \u22120.015 \u22120.398 10. MCU-5 + US PIMA 34.067 34.870 0.803 2.303 3.700 3.520 \u22120.180 \u22125.114 11. MCU-5 + BRAHMA 30.782 29.500 \u22121.282 \u22124.346 3.700 3.420 \u22122.280 \u22128.187 12. DCH-32 + MECH-1 31.157 29.410 \u22121.747 \u22125.940 4.035 4.110 0.075 1.825 13. MECH-1 + GIZA-86 31.017 32.320 1.303 4.032 4.355 4.360 0.005 0.115 Note: UHML indicates the upper half mean length. D ow nl oa de d by [ A st on U ni ve rs ity ] at 0 6: 30 3 0 Ja nu ar y 20 14 Table 3. The HVI results of mixed cottons (25:75). 25:75 Predicted Tested Predicted Tested strength strength Actual Difference elongation elongation Actual Difference S. no. Cotton (g/tex) (g/tex) difference (%) (%) (%) difference (%) 1. MCU-5 + BUNNY 24.400 21.700 \u22122.700 \u221212.442 6.275 6.400 0.125 1.953 2. DCH-32 + MCU-5 25.400 23.200 \u22122.200 \u22129.483 6.600 6.500 \u22120.100 \u22121.538 3. DCH-32 + BUNNY 24.800 21.800 \u22123.000 \u221213.761 6.675 6.500 \u22120.175 \u22122.692 4. US PIMA + S-6 25.575 22.700 \u22122.875 \u221212.665 6.775 7.000 0.225 3.214 5. MECH-1 + US PIMA 28.450 27.300 \u22121.150 \u22124.212 7.125 7.300 0.175 2.397 6. MECH-1 + BRAHMA 25.150 21.700 \u22123.450 \u221215.899 5.925 6.400 0.475 7.422 7. GIZA-86 + DCH-32 27.800 24.100 \u22123.700 \u221215.353 7.750 7.200 \u22120.550 \u22127.639 8. GIZA-86 + BUNNY 26.000 27.100 1.100 4.059 6.625 7.600 0.975 12.829 9. GIZA-86 + S-6 25.925 21.700 \u22124.225 \u221219.470 6.850 6.400 \u22120.450 \u22127.031 10. MCU-5 + US PIMA 28.750 27.200 \u22121.550 \u22125.699 7.025 7.100 0.075 1.056 11. MCU-5 + BRAHMA 25.450 22.800 \u22122.650 \u221211.623 5.825 6.500 0.675 10.385 12. DCH-32 + MECH-1 24.500 23.500 \u22121.000 \u22124.255 6.900 6.500 \u22120.400 \u22126.154 13. MECH-1 + GIZA-86 29.500 25.900 \u22123.600 \u221213.900 7.350 7.000 \u22120.350 \u22125.000 Table 4. HVI results of mixed cottons (50:50). 50:50 Predicted Tested Actual Difference Predicted Tested Actual Difference S. no. Cotton UHML (mm) UHML (mm) difference (%) micronaire micronaire difference (%) 1. MCU-5 + BUNNY 30.260 30.730 0.470 1.529 3.685 3.700 0.015 0.405 2. DCH-32 + MCU-5 32.755 32.810 0.055 0.168 3.575 3.520 \u22120.055 \u22121.563 3. DCH-32 + BUNNY 32.435 32.810 0.375 1.143 3.050 3.250 0.200 6.154 4. US PIMA + S-6 32.310 33.830 1.520 4.493 3.565 3.720 0.155 4.167 5. MECH-1 + US PIMA 32.565 32.650 0.085 0.260 3.965 3.730 \u22120.235 \u22126.300 6. MECH-1 + BRAHMA 30.375 30.280 \u22120.095 \u22120.314 3.965 4.340 0.375 8.641 7. GIZA-86 + DCH-32 33.160 34.400 1.240 3.605 3.640 4.020 0.380 9.453 8. GIZA-86 + BUNNY 31.985 31.770 \u22120.210 \u22120.660 4.105 3.740 \u22120.365 \u22129.760 9. GIZA-86 + S-6 30.620 31.080 0.460 0.480 4.145 4.000 \u22120.145 \u22123.625 10. MCU-5 + US PIMA 32.905 33.720 0.815 2.417 3.870 3.570 \u22120.300 \u22128.403 11. MCU-5 + BRAHMA 30.715 30.180 \u22120.535 \u22121.773 3.870 3.610 \u22120.260 \u22127.202 12. DCH-32 + MECH-1 32.415 33.420 1.005 3.007 3.670 3.760 0.090 2.394 13. MECH-1 + GIZA-86 30.645 29.180 \u22121.465 \u22125.021 4.370 4.370 0.000 0.000 Note: UHML indicates the upper half mean length. D ow nl oa de d by [ A st on U ni ve rs ity ] at 0 6: 30 3 0 Ja nu ar y 20 14 Table 6. HVI results of mixed cottons (75:25). 75:25 Predicted Tested Actual Difference Predicted Tested Actual Difference S. no. Cotton UHML (mm) UHML (mm) difference (%) micronaire micronaire difference (%) 1. MCU-5 + BUNNY 30.420 30.840 0.420 1.326 3.947 4.180 0.233 5.574 2. DCH-32 + MCU-5 33.840 33.740 \u22120.100 \u22120.296 3.257 3.040 \u22120.217 \u22127.138 3. DCH-32 + BUNNY 33.682 34.260 0.578 1.687 2.995 3.010 0.015 0.498 4. US PIMA + S-6 33.770 34.260 0.490 1.430 3.547 3.710 0.163 4.394 5. MECH-1 + US PIMA 31.232 31.720 0.488 1.538 4.182 4.080 \u22120.102 \u22122.500 6. MECH-1 + BRAHMA 30.137 29.320 \u22120.817 \u22122.786 4.182 3.590 \u22120.592 \u221216.490 7. GIZA-86 + DCH-32 32.275 32.460 0.185 0.570 3.990 4.290 0.300 6.993 8. GIZA-86 + BUNNY 31.027 31.250 0.223 0.714 4.045 4.170 0.125 2.998 9. GIZA-86 + S-6 30.890 30.470 \u22120.420 \u22121.378 4.155 4.270 0.115 2.693 10. MCU-5 + US PIMA 31.742 32.070 0.328 1.023 4.040 3.620 \u22120.440 \u221210.89 11. MCU-5 + BRAHMA 30.647 30.000 \u22120.647 \u22122.157 4.040 3.80 \u22120.240 \u22126.32 12. DCH-32 + MECH-1 33.672 35.200 1.528 4.341 3.305 3.480 0.175 5.029 13. MECH-1 + GIZA-86 30.272 29.580 \u22120.692 \u22122.339 4.385 4.180 \u22120.205 \u22124.905 Note: UHML indicates the upper half mean length. values were calculated on the basis of the following formula: p1l1 + p2l2 = Lp. (1) where p1 is the proportion of fibre 1; l1 is the length of fibre 1; p2 is the proportion of fibre 2; l2 is the length of fibre 2; and Lp is the predicted length of blend. Similarly strength (Strp), elongation (Ep) and micronaire (Micp) have been calculated from individual cotton properties. The actual values of blends were obtained from HVI by directly testing the blended cotton. The results show that difference in length values varies from 0.2% to 6% averaging 3.15% for 25:75 blends, 1.91% for 50:50 blends and 1.66% for 75:25 blends. The difference in predicted versus actual micronaire values averaged 4%\u20135% for all blends. The difference in strength values is highest amongst the properties considered averaging 6%\u201311% for different blend proportions. The difference in elongation values averaged from 4.5%\u20135.0% for different blend proportions. The difference in length values may be due to difference in distribution of fibres of different length in a sample. When two bunches of fibres of differing extension values or strength values are taken and load is applied on them, first whole load is shared by all fibres, and subsequently fibres that have less breaking strength and less extension start D ow nl oa de d by [ A st on U ni ve rs ity ] at 0 6: 30 3 0 Ja nu ar y 20 14 1 2 3 4 5 6 7 8 9 10 11 12 13 0 5 10 15 20 25 30 Actual strengthPredicted strength Blends S tre ng th (g /te x) Figure 2. Actual versus predicted fibre strength (50:50). to break, putting the whole load on, less number of fibres. This results in lower breaking strength value when mixture of two fibres is taken. The difference in breaking extension may be explained by the same reasoning. The difference in micronaire values in the range of 4%\u20135% is astonishing. The compressed air method is followed by HVI to measure the micronaire of cottons. Variation in diameter or variation of crimps may have caused differing airflow through two fibres and thus affected the result. Figures 1 and 2 show the predicted and tested length values and strength values of the blends. It can be seen that in most of the cases, the tested strength values are lower than predicted values. In the case of length values, some show higher values, whereas some lower. The regression equations (Equations (2)\u2013(13)) obtained are as follows: Influence of blend on properties of cottons For 75:25 blend StrA = 0.978 Strp \u2212 1.143 R2 = 0.81 (2) LA = 0.944 Lp \u2212 2.03 R2 = 0.46 MicA = 0.77 Micp + 0.754 R2 = 0.58 EA = 1.13 Ep \u2212 0.859. R2 = 0.80 For 25:75 blend StrA = 0.9 Strp + 0.241 R2 = 0.69 (10) LA = 1.18 Lp \u2212 5.612 R2 = 0.75 (11) MicA = 0.72 Micp + 1.103 R2 = 0.63 (12) EA = 0.48 Ep + 3.507. R2 = 0.61 (13) From the equations, it can be seen that in the case of 50:50 blend, least correlation was observed. When both constituents are of equal proportion, the variation in distribution is maximum, which may be the reason for the same. Though there are some differences in absolute values, t tests show that only actual strength values are significantly different from calculated values (Table 8). The differences between actual and calculated values of length, micronaire and elongation are not significant. The results discussed so far have shown that there is a difference between calculated or predicted fibre properties and actual fibre properties in a blend. South India Textile Research Association (SITRA) earlier has done a lot of work in deriving prediction equations for yarn properties on the basis of fibre properties. The latest study (Chellamani, Thanabal, Basu, & Ratnam, 2004) reports the prediction of yarn count strength product (CSP) and rupture per kilometre tenacity (gm/tex) on the basis of cotton fibre properties given by HVI (tested in HVI calibration mode). Those equations were derived from single cottons. Yarns have been spun using two cottons at a time with blend proportions 50:50, 75:25 and 25:75. For each blend, yarns of two fineness values were spun. The properties of the yarns have been reported in Table 9. Also, the fibre D ow nl oa de d by [ A st on U ni ve rs ity ] at 0 6: 30 3 0 Ja nu ar y 20 14 quality indices (FQI) calculated on the basis of calculated blend fibre properties and actual test values of blended cottons have been reported. Similarly, CSP and U% have been calculated from those results; these were presented along with actual CSP and U% of the yarns. For calculation of FQI, CSP and U%, Equations (14)\u2013 (16) have been used: FQI = LS f (14) CSP = 165 (\u221a FQI ) + 590 \u2212 13C (15) U 2 = 137 ( f L )2 (d \u2212 1) d Ne and 4.1 (d \u2212 1) + U 2 r . (16) Where L= Mean length of fibre (mm) S = Strength in HVI mode (g/tex) f = Fibre fineness (micronaire) C = Yarn count (Ne) d = Draft used on ring frame Ur = U% of roving The results from Table 9 show that the difference in actual CSP and predicted CSP (both from calculated fibre strength and actual fibre strength) is significant in most of the cases. The difference is more when the difference in properties of constituent fibres is more (especially extension). The overall regression equations are as follows: Actual CSP = 1.05 CSPActF + 242.52 R2=0.63 (17) Actual CSP = 1.10 CSPCalF + 173.22 R2=0.57 (18) Actual U% = 0.568U%ActF + 4.78. R2=0.55 (19) Where CSPCalF is CSP calculated from predicted fibre strength values CSPActF is CSP calculated from actual fibre strength values U%Act is U% calculated based on actual fibre properties The comparative tests (t-test) show that both predicted CSP values from actual blended fibre strength values (t = 4.08) and predicted CSP values from calculated fibre strength values (t = 4.615) are significantly different from the actual CSP values. The CSP values predicted from actual blend strength values are marginally closer to the actual CSP values. The predicted U% values and actual U% values show no significant difference (t = 0.48). This is expected as fibre length and fineness properties of actual blended fibres are not significantly different from weighted average values. Similar to CSP values, the single yarn tenacity values tenacity were predicted using Equation (20) Tenacity (g/tex) = 1.1 \u221a FQI + 4.0 \u2212 ( 13C 150 ) (20) calculated blended fibre strength values and actual blended fibre strength values. These are shown in Table 10. The table shows that the difference in tenacity values is minimum in the case of blends between MECH1 and S6. Individual fibre properties show that they have minimum difference in breaking elongation values of the constituent fibres. The correlation analysis shows the following relationships: Actual tenacity = 0.905 RKMActF + 3.56 R2=0.685 (21) Actual tenacity = 0.955 RKMCalF + 3.008. R2=0.66 (22) Where tenacityActF is the predicted RKm based on actual blend fibre strength, and RKmCalF is predicted RKm based on calculated blend fibre strength. The comparative analysis between actual tenacity values and predicted values shows that these are significantly different. The predicted values based on actual blend fibre strength are marginally closer (t = 3.53) to that based on calculated blend fibre strength (t = 4.03). The above analysis shows that the prediction equations applicable for single fibre yarn do not fit well with yarns made of more than one fibre. After observing the inaccuracies of the prediction equations for single cotton, when applied for blended cottons, further experiments were conducted to find new equations for binary blends. In addition to blends used in earlier experiments, two other mixings USPIMA + BRAHMA and USPIMA + BUNNY with proportions 75:25, 50:50 and 25:75 were used to spin the yarns of two fineness levels (60s and 80s Ne). After combining all the results, a series of equations were derived. Table 11 shows the comparative accuracies against the equations used earlier. It can be seen from Table 11 that although two new Equations (23) and (24) showed nearly the same variance (R2) values, the errors have been reduced considerably in the case of all three new Equations (23), (24) and (25) as compared with the conventional Equations (15), (16) and (20). However, the low variances noted are a matter of concern. Of course, this could be improved by using a larger number of samples, but in practice, that should increase the cost beyond tolerable limit. This means that the predicted values derived from the new equations will be a step closer to the actual values, but we have to work continuously to improve even more. D ow nl oa de d by [ A st on U ni ve rs ity ] at 0 6: 30 3 0 Ja nu ar y 20 14 Ta bl e 9. In te ra ct io n of th e pr op er ti es of in di vi du al co tt on fi br es in a bl en d \u2013 H V I m od e. Fi br e qu al it y in de x C S P D if fe re nc e (% ) M ix C ou nt F m A c F m P r C S P F m A c F m P r F m A c F m P r U % U % D if fe re nc e S .n o S am pl e pr op or ti on C ou nt S tr S tr (A ct ua l) S tr S tr S tr S tr (A ct ua l) (P re di ct ed ) (% ) 1. D C H -3 2 + M C U -5 50 :5 0 60 s 20 1. 20 20 0. 88 27 44 21 50 .4 2 21 48 .5 6 21 .6 0 21 .7 0 13 .5 3 17 .1 9 \u22122 7. 08 2. D C H -3 2 + M C U -5 50 :5 0 80 s 20 1. 20 20 0. 88 24 31 18 90 .4 2 18 88 .5 6 22 .2 0 38 .9 0 15 .5 6 19 .7 8 \u22122 7. 12 3. D C H -3 2 + M C U -5 75 :2 5 60 s 23 5. 17 23 4. 26 28 35 23 40 .3 2 23 35 .4 0 17 .4 0 17 .6 0 13 .6 9 15 .8 5 \u22121 5. 74 4. D C H -3 2 + M C U -5 75 :2 5 80 s 23 5. 17 23 4. 26 25 67 20 80 .3 2 20 75 .4 0 19 .0 0 19 .5 1 15 .5 3 18 .2 0 \u22121 7. 17 5. D C H -3 2 + M C U -5 25 :7 5 60 s 19 6. 76 19 3. 14 23 74 21 24 .4 7 21 03 .0 7 10 .5 0 11 .4 0 15 .7 3 17 .4 1 \u22121 0. 66 6. D C H -3 2 + M C U -5 25 :7 5 80 s 19 6. 76 19 3. 14 21 20 18 64 .4 7 18 43 .0 7 12 .1 0 13 .1 0 17 .1 7 20 .0 3 \u22121 6. 65 7. G IZ A -8 6 + M E C H -1 50 :5 0 40 s 22 0. 04 20 1. 06 31 45 25 17 .5 5 24 09 .6 2 20 .0 0 23 .4 0 12 .6 0 15 .8 4 \u22122 5. 69 8. G IZ A -8 6 + M E C H -1 50 :5 0 50 s 22 0. 04 20 1. 06 28 18 23 87 .5 5 22 79 .6 2 15 .3 0 19 .1 0 13 .7 3 17 .6 5 \u22122 8. 52 9. G IZ A -8 6 + M E C H -1 75 :2 5 40 s 24 5. 51 23 0. 85 32 07 26 55 .3 6 25 76 .9 4 17 .2 0 19 .6 0 12 .0 5 15 .1 9 \u22122 6. 07 10 . G IZ A -8 6 + M E C H -1 75 :2 5 50 s 24 5. 51 23 0. 85 31 42 25 35 .3 6 24 46 .9 4 19 .3 0 22 .1 0 13 .1 7 16 .9 1 \u22122 8. 43 11 . G IZ A -8 6 + M E C H -1 25 :7 5 40 s 23 4. 04 22 6. 86 29 52 25 94 .2 3 25 55 .2 1 12 .1 0 13 .4 0 12 .6 6 14 .2 7 \u22121 2. 74 12 . G IZ A -8 6 + M E C H -1 25 :7 5 50 s 23 4. 04 22 6. 86 27 54 24 64 .2 3 24 25 .2 1 10 .5 0 11 .9 0 13 .9 0 15 .8 7 \u22121 4. 18 13 . M E C H -1 + S -6 50 :5 0 40 s 21 7. 88 21 2. 78 25 22 25 05 .5 4 24 76 .8 6 0. 65 1. 80 13 .9 0 14 .0 1 \u22120 .7 6 14 . M E C H -1 + S -6 50 :5 0 50 s 21 7. 88 21 2. 78 24 29 23 75 .5 4 23 46 .8 6 2. 20 3. 40 14 .6 6 15 .5 7 \u22126 .2 2 15 . M E C H -1 +S -6 75 :2 5 40 s 18 6. 95 18 6. 00 26 05 23 26 .0 3 23 20 .2 7 10 .7 0 10 .9 0 13 .6 8 15 .0 6 \u22121 0. 11 16 . M E C H -1 + S -6 75 :2 5 50 s 18 6. 95 18 6. 00 25 09 21 96 .0 3 21 90 .2 7 12 .5 0 12 .7 0 14 .8 9 16 .7 7 \u22121 2. 63 17 . M E C H -1 + S -6 25 :7 5 40 s 18 1. 29 17 7. 33 25 59 22 91 .6 1 22 67 .2 5 10 .4 0 11 .4 0 14 .0 1 15 .4 5 \u22121 0. 31 18 . M E C H -1 + S -6 25 :7 5 50 s 18 1. 29 17 7. 33 23 32 21 61 .6 1 21 37 .2 5 7. 30 8. 35 14 .9 9 17 .2 1 \u22121 4. 84 F m A c S tr ,f ro m ac tu al bu nd le st re ng th ;F m P r S tr ,f ro m pr ed ic te d fi br e bu nd le st re ng th ;C S P, co un ts tr en gt h pr od uc t; *n = 20 . D ow nl oa de d by [ A st on U ni ve rs ity ] at 0 6: 30 3 0 Ja nu ar y 20 14 Table 10. Actual versus predicted single-yarn tenacity (gm/tex). Mixing Tenacity Fm Tenacity Fm Actual Difference Difference S. no. Sample proportion Count Ac Str* Pr Str* tenacity for Ac Str (%) for Pr Str (%) 1. DCH-32 + MCU-5 50:50 60s 14.40 14.39 19.61 26.55 26.62 2. DCH-32 + MCU-5 50:50 80s 12.67 12.66 16.61 23.73 23.80 3. DCH-32 + MCU-5 75:25 60s 15.67 15.64 18.95 17.32 17.49 4. DCH-32 + MCU-5 75:25 80s 13.94 13.90 16.49 15.49 15.69 5. DCH-32 + MCU-5 25:75 60s 14.23 14.09 15.41 7.66 8.59 6. DCH-32 + MCU-5 25:75 80s 12.50 12.35 14.01 10.80 11.82 7. GIZA-86 + MECH-1 50:50 40s 16.85 16.13 19.55 13.81 17.49 8. GIZA-86 + MECH-1 50:50 50s 15.98 15.26 18.14 11.89 15.85 9. GIZA-86 + MECH-1 75:25 40s 17.77 17.25 21.37 16.85 19.30 10. GIZA-86 + MECH-1 75:25 50s 16.90 16.38 20.45 17.34 19.91 11. GIZA-86 + MECH-1 25:75 40s 17.36 17.10 19.53 11.10 12.43 12. GIZA-86 + MECH-1 25:75 50s 16.50 16.24 17.99 8.31 9.76 13. MECH-1 + S-6 50:50 40s 16.77 16.58 16.03 \u22124.62 \u22123.42 14. MECH-1 + S-6 50:50 50s 15.90 15.71 15.77 \u22120.85 0.36 15. MECH-1 + S-6 75:25 40s 15.57 15.54 16.64 6.41 6.63 16. MECH-1 + S-6 75:25 50s 14.71 14.67 16.38 10.21 10.45 17. MECH-1 + S-6 25:75 40s 15.34 15.18 16.41 6.50 7.48 18. MECH-1 + S-6 25:75 50s 14.48 14.32 15.86 8.72 9.74 Note: Tenacity indicates rupture per kilometre; Fm Ac Str, from actual bundle strength; Fm Pr Str, from predicted fibre bundle strength; *Tenacity (g/tex) = 1.1 \u221a FQI + 4.0\u2212(13C/150)." + ] + }, + { + "image_filename": "designv11_69_0003187_mesa.2010.5552012-Figure8-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003187_mesa.2010.5552012-Figure8-1.png", + "caption": "Fig. 8. Step I, find the tangent", + "texts": [ + " If the mobile robot cannot directly arnve at the target point, the robot may choose some sub-targets from the current map which has been drawn. According to the shortest path principle, the robot can determine its motion direction at the current situation. The method for searching the sub-target is as follows. Step 1: In the updated map, the mobile robot explores along the target direction, if there are obstacles, the robot determines the tangent line and the tangent point. If not, then the sub-target is the globe target. As shown in Fig. 8, the large circle is the detection range of the mobile robot. The shaded parts are the obstacles which have been detected. The dotted lines are the obstacles that have not been detected. Points A, B, C, D and E are the tangent points. Step 2: The point which is apart from the tangent point at a distance of the robot size is chosen as the starting point, then repeat Step 1 until all the starting points can be connected directly with the target point. As shown in Fig. 9, points A, B and E need to search the starting points to connect to the target point" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003450_mace.2010.5536045-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003450_mace.2010.5536045-Figure1-1.png", + "caption": "Figure 1. The waviness model of roller bearing", + "texts": [ + " Using the Newmark- method, the differential equations of motion can be solved. The dynamic properties of roller bearing are impacted by vibration of the bearing. If nonlinear vibration displacements of roller bearing are obtained, the dynamic properties of bearing considering nonlinear vibration can be computed. An important source of vibration in roller bearings is waviness. If it is assumed that the waviness of periodic lobes is a sinusoidal function, the radial waviness of the outer race, inner race and rollers, as shown in Fig. 1, can be expressed as ( )1 1 1 1 1 2 cosj n c n n j P A n t N \u03c0\u03c9 \u03c9 \u03be \u221e = = \u2212 + + (3) ( )2 2 2 2 1 2 cosj n c n n j P A n t N \u03c0\u03c9 \u03c9 \u03be \u221e = = \u2212 + + (4) Where, A is amplitude of race waviness, n is waviness order, \u03be is initial phase angle of waviness, 1\u03c9 , 2\u03c9 , c\u03c9 are rotating speed of outer ring, inner ring and cage, N is number of rollers. The roller waviness can be given by: [ ]1 1 cosj nj b bj n W C n t\u03c9 \u03be \u221e = = + (5) [ ]2 1 cos ( )j nj b bj n W C n t\u03c9 \u03c0 \u03be \u221e = = + + (6) Where, C is amplitude of race waviness, b\u03c9 is rotating speed of roller" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002979_978-3-540-70534-5_7-Figure14.5-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002979_978-3-540-70534-5_7-Figure14.5-1.png", + "caption": "Figure 14.5: 3-dof manipulator example", + "texts": [], + "surrounding_texts": [ + "We can now fill in the Denavit\u2013Hartenberg table for this sample manipulator by placing individual local coordinate systems in each joint and finding the four parameters each for the transition from joint to joint (see Figure 14.6). Note that right-handed coordinate systems are used at all times. The overall manipulator transformation (going from base 0 until after joint" + ] + }, + { + "image_filename": "designv11_69_0003086_2009-01-2121-Figure4-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003086_2009-01-2121-Figure4-1.png", + "caption": "Figure 4. The transfer paths of powertrain vibration", + "texts": [ + " As Figure 3 shows, the summed inertia force can be located on the center of the gravity of an engine which has some distance from the center of the gravity of a powertrain, and it causes the pitching motion of a powertrain. In order to cancel out these 2nd-order inertia forces, a BSM is usually applied under a crank shaft. A BSM is composed of two unbalanced rotors which rotate at the twice angular speed of a crank shaft, and compensates the inertia force of the 2nd-order frequency. We have studied the effect of removing a BSM from an SUV with a 2 \u2113 diesel engine by measuring interior cabin noise as well as the source-and-body-side accelerations of each path; as Figure 4 shows, among a lot of structural transfer paths of powertrain vibration such as 4 powertrain mounts, drive shafts, muffler, hoses, and so on, we have focused on the 4 powertrain mounts which are usually main transfer paths connecting powertrain and vehicle body. At first, with the source-side acceleration data, we observed the motion of powertrain at high engine rpm. Figure 5 shows and compares the powertrain motion of the 2nd-order frequency at 3500 rpm before and after removing the BSM. As explained earlier, it was found that the pitching motion of powertrain became much larger due to inertia forces when the BSM was removed, and the source-side motion of the engine mount increased predominantly compared with other mount paths" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002291_j.mechmachtheory.2007.07.002-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002291_j.mechmachtheory.2007.07.002-Figure2-1.png", + "caption": "Fig. 2. Limit line and parameter lines on the generating surface R1.", + "texts": [ + " When Ut = 0, there is a curve L(2) on R1 which is defined by the equations r1 \u00bc r1\u00f0u; v\u00de U1 \u00bc U1\u00f0u; v;u1;u2; \u00de \u00bc 0 U2 \u00bc U2\u00f0u; v;u1;u2; \u00de \u00bc 0 Ut \u00bc Ut\u00f0u; v;u1;u2; \u00de \u00bc 0 8>>< >>: \u00f020\u00de It is worth mentioning that here Ut is zero only at the points on L(2) not all the points on R1. On R1, from Eq. (18), the tangent vector to u1-line is given by r1u1 and the tangent vector to L(2) may be given by r1u2 according to Eq. (20). Because Ut = 0 on L(2), from Eq. (19), we have r1u1 r1u2 \u00bc 0. Therefore, u1-lines are tangential to L(2). Similarly, u2-lines are also tangential to L(2). Furthermore, L(2) is the common enveloping curve of u1-lines and u2-lines on R1 as shown in Fig. 2. The intersections of u1-lines and u2-lines are the meshing points between R1 and R2. R1 is divided into two sections by L(2): useless zone and meshing zone. Any point in the useless zone will never be a contact point in the action. However, any point in the meshing zone will be a contact point sooner or later. Consequently, we define L(2) as the meshing limit line on R1, Ut as the meshing limit function and Ujuk as the partial meshing limit function. Sometimes, L(2) is also called the limit line of the second kind" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000106_vesd.36.6.445.3545-Figure4-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000106_vesd.36.6.445.3545-Figure4-1.png", + "caption": "Fig. 4. Tyre structural model.", + "texts": [ + " Its chassis is a stiffened variant of the Ertz chassis [9], namely the torsional stiffness of each structural element is increased by the factor four. The total mass of the truck model, located at 111 nodes, is equal to10384 kg, and the number of degrees of freedom (NDOF) is equal to 1874. In the suspension sets leaf springs are applied, while the nonlinear dampers are positioned above the axles. The data of the truck structure have been prepared on the basis of the industrial experience collected at the department of the writer. In each example the velocity of the truck is equal to 20 m/s. The tyre structural model (Fig. 4, NDOF 306) is an improvement of the three dimensional rigid ring tyre model [14] by the consideration of the large D ow nl oa de d by [ T he U ni ve rs ity o f M an ch es te r L ib ra ry ] at 0 5: 54 1 2 O ct ob er 2 01 4 carcass deformations near the contact patch in vertical, lateral and yaw directions [8]. There can be seen in Figures 5\u00b16 the applied tyre force characteristics, computed by the cosine version of the magic formula [11] in correspondence with dry and icy roads, respectively. Self-aligning moments are also taken into consideration" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002470_03091900600926898-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002470_03091900600926898-Figure2-1.png", + "caption": "Figure 2. Schematic of the enzyme electrode with enzyme layer and diffusion layer.", + "texts": [ + " If K1 and K2 are two extreme points on either side of the enzyme membrane then the rate of reaction at point K1 in the enzyme membrane (Vk1) as shown in figure 1 is: Vk1 \u00bc DSe @Se @x DA \u00f04\u00de Similarly, the rate of reaction at point K2 (Vk2) is: Vk2 \u00bc DSe @Se @x DA; \u00f05\u00de where D is the diffusion coefficient, DSe is the diffusion coefficient of substrate, @Se @x is the rate of change of diffusion at point x (any point between K1 and K2), and A is the effective area. Therefore the rate of change of mass transfer of substrate can be given by: @Se @x \u00bc DSe @2Se @x2 : \u00f06\u00de For modelling of the amperometric biosensor the following parameters as shown in figure 2, were considered: . l is the thickness of the enzyme layer; . L is the boundary of the diffusion layer; . d\u00bcL\u2013 l is the thickness of the diffusion layer; . Sen and Sbu are the concentrations of substrate in the enzyme and the bulk solution, respectively; . Pen are Pbu are the concentrations of reactive product in the enzyme and in the bulk solution, respectively; . DSen and DPen are diffusion coefficients of substrate and reactive products, respectively, in the enzyme layer; . DSbu and DPbu are diffusion coefficients of substrate and reactive products, respectively, in the bulk solution; and " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001203_esda2006-95425-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001203_esda2006-95425-Figure2-1.png", + "caption": "Fig. 2 \u2013 Definition of wheel load in the no-longitudinal-acceleration case", + "texts": [], + "surrounding_texts": [ + "State-of-the-art in research concerning the control of active dampers is the knowledge that driving comfort on the one hand and driving safety on the other hand can be improved by the control of active dampers. This holds true for passenger cars as well as for other land-based vehicles (such as trucks or tanks) and has been proved in simulations and in real test drives. For an exceptional example refer to Choi [1]. Not talking about driving safety in general, but rather about the RMS on wheel load in particular, this quantity is known as reducible by making use of active shock absorbers. This field has widely been researched mainly in simulation models (refer to Redlich [2], Alberti [3] or El-Demerdash [4]). Several experiments have been undertaken with quarter or half car models (refer to Yi [5]). Val\u00e1\u0161ek [6] was able to reduce the RMS on wheel load in the case of a heavy truck. His goal was to reduce the road damage. Even though there are all those results of the former, until now the possibility to influence the wheel load by active dampers has not been used during ABS-braking situations. In those cases the dampers are switched to a passive setting \u2013 most of the time it is the setting with the highest (hardest) possible damping that is chosen, because a hard damping is supposed to be better for full-braking than a soft damping. Making the assumption that a well directed influence on wheel load can enhance braking performance and having the former research activities (coupling between control of active Copyright \u00a9 2006 by ASME rms of Use: http://www.asme.org/about-asme/terms-of-use dampers and wheel load) in mind, the hypothesis is made that it is possible to enhance braking performance by control of active dampers." + ] + }, + { + "image_filename": "designv11_69_0000132_978-94-017-0371-0_21-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000132_978-94-017-0371-0_21-Figure2-1.png", + "caption": "Figure 2. Setup for tension tests, (a) regions cut tingout specimens, (b) gauge section of a specimen and (c) experimental setup.", + "texts": [ + " Five fan-shaped petal units are jointed and make a morning glory flower. At the lower half of the bud body (Fig.l(b \u00bb, the petal ribs join together and construct a hollow tube. From the whole view of a flower, it appears that the petal ribs support the whole body of a flower. In order to obtain the mechanical properties of the petal ribs and petals, static tensile tests were carried out. For the tests, several rectangular tensile specimens whose gauge section is also a rectangle of 16 x 20 mm were prepared from these two parts, see Fig.2. The average measured thickness of these two parts are shown later in Table 1. After setting the specimen clamped by paper clips on the bed of a profile projector as shown in Fig.2(c), an initial load of 0.0049 N (0.5 gf) was applied by a dead weight. The longitudinal and trans verse elongations of the specimen were measured using four painted markers every 0.0245 N (2.5 gf). Fig.3 shows the nominal stress-strain curves of petals and petal ribs. Although these data were off-set due to an initial load, each data almost cluster on a line. The slope of these lines are 1.03 MPa and 3.0 MPa which correspond to Young's moduli of petals and petal ribs of morning glory flowers, respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001327_1.2080521-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001327_1.2080521-Figure2-1.png", + "caption": "FIGURE 2 \u2014 Schematic drawing of a layered stack of different materials.", + "texts": [ + " The samples were annealed at 100\u00b0C for 2 minutes and then exposed to polarized UV light. The authors are with Fuji Photo Film Co., Ltd., 210 Nakanuma, Minamiashigara, Kanagawa 250-0193, Japan; e-mail: ichirou_amimori@fujifilm.co.jp \u00a9 Copyright 2005 Society for Information Display 1071-0922/05/1309-0799$1.00 Journal of the SID 13/9, 2005 799 The peak irradiance and the energy density were 1.1 W/cm2 and 1.8 J/cm2, respectively. XRD measurements were performed by a grazing incidence in-plane diffractometer (ATX-G, CuK\u03b1 radiation of 0.154 nm, 50 kV, 300 mA, Rigaku Corp.) as shown in Fig. 2. Soller slits of 0.48\u00b0 and 0.45\u00b0 vertical divergences were used for the source and detector, respectively. A divergent slit (1 \u00d7 10 mm) was used to control the exposure area of incident X-ray, while another divergent slit (0.1 \u00d7 10 mm) was used to limit the output signal from a sample. Retardation of samples at \u03bb = 589 nm was measured by using the parallel-Nicole rotation method (KOBRA-WR, Oji Scientific Instruments). The dependence of retardation on incident angles was investigated. Figure 3 illustrates the configuration of the sample and incident light in the retardation measurement" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002553_1.2900720-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002553_1.2900720-Figure1-1.png", + "caption": "Fig. 1 The basic concept of the virtual cam method", + "texts": [ + " We will also how how to combine this method along with Pennock\u2019s method o solve problems that cannot be solved by either Pennock\u2019s or our ethod alone. Contributed by the Mechanisms and Robotics Committee of ASME for publicaion in the JOURNAL OF MECHANICAL DESIGN. Manuscript received June 15, 2007; final anuscript received October 18, 2007; published online April 15, 2008. Review onducted by Qizheng Laio. ournal of Mechanical Design Copyright \u00a9 20 om: http://mechanicaldesign.asmedigitalcollection.asme.org/pdfaccess.ash Referring to Fig. 1, a camlike appendage is introduced into a given linkage system such that this translational cam has contact with one of the links and consequently that link serves as the follower of the cam at this moment. This modification will not change the total degrees of freedom of the original mechanism. By adjusting the position and orientation of this imaginary cam, we get to achieve a desired configuration in that a virtual four-bar loop is formed. From that virtual four-bar loop, some key instant centers, whose exact locations were previously obscure, can now be derived via Kennedy\u2019s theorem. As a result, all instant centers of the mechanism can be finally located. 2.1 Preliminary Considerations. Figure 1 is the schematic of a five-bar segment within a typical kinematically indeterminate linkage. Such a linkage, although containing loops of more than five bars, possesses only one degree of freedom, so that it has a calculable, definitive pattern of motion. Link 1 is designated as the ground link, as shown in Fig. 1. Links 2\u20136 are consecutively connected via Revolute Joints A\u2013D. Since Links 2\u20136 do not form a four-bar loop, we cannot use Kennedy\u2019s theorem to find additional instant centers other than those at the joints. However, we do know from the theorem that the instant center between Links 2 and 6 I62 and the instant centers with respect to the ground link I21 and I61 will lie on a straight line. For example, let us say that Link 4 in Fig. 1 is our link of interest. This means that we want to know its velocity and therefore we are for its instant Center I41. So, I41 is regarded as the key instant center. The two links within the five-bar segment immediately connecting to Link 4 i.e., Links 3 and 5 are called primary adjacent links for Link 4. Should there be any other links connecting to Link 4, they are then secondary adjacent links for that link. The other two links 2 and 6 in the figure also play a very important role in our method, and they are called dangling links in this paper", + " A number of examples are included in this paper to illustrate he use of the virtual cam method. In our first two examples, the xact location of I41 can be quickly obtained from the instant enters of its secondary adjacent links with respect to the ground ink. In some more complicated kinematically indeterminate linkges, as shown in our last example, Pennock\u2019s method is incorpoated to help locate the exact position of the desired Instant Center 41. 2.2 Orienting the Virtual Cam. We start with looking into he instant centers related to the two dangling links. From Fig. 1, triplet of such Instant Centers I12, I26, and I16 lies on a line. (a) (b) ig. 2 To construct I26I12I2i similar to I26I16I6i by graphical ethod: \u201ea\u2026 when I26 is on Line I12I16 and \u201eb\u2026 when I26 is on the xtension part of Line I12I16 62304-2 / Vol. 130, JUNE 2008 om: http://mechanicaldesign.asmedigitalcollection.asme.org/pdfaccess.ash According to Kennedy\u2019s theorem, Instant Centers I2i, I26, and I6i also lie on a line. Moreover, the line extending from I12 to I2i and the line extending from I16 to I6i should both lead to I1i at infinity. Therefore, Lines I12I2i and I16I6i are parallel, and Triangles I12I2iI26 and I16I6iI26 are similar triangles, as depicted in Fig. 1. We are to use this geometric property to find the orientation of the virtual cam; that is, the direction of Instant Center I1i. Figure 2 a depicts the given as well as the required conditions. L1 and L2 are two lines extending from Links AB and CD in Fig. Transactions of the ASME x?url=/data/journals/jmdedb/27875/ on 04/05/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use 1 C l a s t I i I p t p L T F t J Downloaded Fr , the two primary adjacent links of the link of interest. Instant enters I12, I16, and I26 are known. Now, we want to construct a ine, passing through I26, which connects L1 and L2 at Points I2i nd I6i, respectively, so that Triangles I26I12I2i and I26I16I6i are imilar triangles. Comparing Fig. 2 a with Fig. 1, we can see that he direction of Lines I12I2i and I16I6i signifies the direction of nstant Center I1i, which also is the orientation of the virtual cam . Described as follows are the procedures of drawing the Line 2iI6i. First, we connect I26 and I4i. From I12 and I16, draw lines arallel to I26I4i, intersecting L1 and L2 at Points I and J, respecively. Connect Points I and J. Then, through I26, construct a line arallel to Line IJ, intersecting L1 and L2 at Points I2i and I6i. This ine I2iI6i is then the line we want" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002062_s11465-007-0006-x-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002062_s11465-007-0006-x-Figure2-1.png", + "caption": "Fig. 2 Coordinate of a single pad for calculation", + "texts": [ + " Consider the following Reynolds boundary problem arising in fluid lubrication (dimensionless) W Y( )p = (3) where W( )p d p d p = -\u2212 \u239b \u239d\u239c \u239e \u23a0\u239f \u239b \u239d\u239c \u239e \u23a0\u239f y y y y y y y yw m w j l m l 3 2 3 and Y= - + +\u22123 6[( cos sin ) ( cos sin )]x y y xw w w w (4) where d x y= + +1 sin cosw w\u2014oil film thickness, dimensionless; p \u2014oil film pressure, dimensionless; m \u2014dynamic viscosity of oil, dimensionless; j\u2014diameter-to-width ratio; x \u2014displacement of rotor in x direction, dimensionless; y \u2014displacement of rotor in y direction, dimensionless; x \u2014velocity of rotor in x direction, dimensionless; y \u2014velocity of rotor in y direction, dimensionless. w (radian, dimensionless) is the angle between the negative direction of y axis and the oil film location; h is the eccentric angle (radian, dimensionless); Q is the angle between the connection line of the eccentric angle with the center of bearing and oil film location (radian, dimensionless); G is the weight of rotor (dimensionless); l is the axial coordinate of calculation (dimensionless); and v is the rotating speed of rotor (dimensionless), as shown in Fig. 2. Equation (3) is equivalent to the following discrete form of elliptical variational inequalities [4] W( , ) ( ), , , , p q q p p p q K b i i = y y y = y Q 0 0 0 in inV V \u2200 e (5) When state variables X = =( , , , ) ( , , , )X X X X x y x y1 2 3 4 T T are introduced, the corresponding system equation in state space is written as X = = = = + + = + x X y X x F m e t e t y F m g e x x y y y 3 4 2 2 2 \u2212 \u2212 + / cos sin / v v v v v cos sinv v vt e tx- 2 \u23a7 \u23a8 \u23aa \u23aa \u23a9 \u23aa \u23aa (2) where V\u2014the oil field of the single pad; yV\u2014boundary of V; H0 1 ( )V \u2014Sobolev space; Qb\u2014the angle between the connection line of the eccentric angle with the center of the bearing and intersection curve of cavitation field with no cavitation field of oil film (the intersection curve changes with the change of perturbation of journal displacements and velocities), dimensionless; K = { p He 0 1, pi0 in V} Y Y( )q q= \u22c5\u222b\u222bV Vd \u2014linear function W( , )p q d p q p q = y y y y + y y y y 3 2 m w w j l lV V\u222b\u222b \u239b \u239d\u239c \u239e \u23a0\u239f d \u2014the restricte d, symmetric, and elliptical bilinear functional on H H0 1 0 1( ) ( )V Vx By using the eight-node isoparametric FEM, the distribution of pressures of nodal points pi in oil film field is solved", + " (7), (10) and (13), when Pk k x y x y( , , , )= are calculated, because Eqs. (12) and (7) have the same coefficient matrix Wf . It is evident that Jacobian of the oil film forces is obtained when the oil film forces are calculated. Therefore, accuracy and reliability of the analysis of nonlinear oil film forces are ensured, and the computing works spent on the Jacobians are much less than those spent on the oil film forces. After calculating the oil film forces and their Jacobians of single pad (shown in Fig. 2), oil film forces and their Jacobians of elliptical bearing (shown in Fig. 3) can be estimated according to the characteristic of elliptical bearing. O1, O2 are the center of lower pad and upper pad, respectively, as shown in Fig. 3. solution of Cauchy form of differential equation that can be written as X u F X u X F X u X u X u( ) ( , ) ( , ) ( ) = x = \u2212 \u23a1 \u23a3\u23a2 \u23a4 \u23a6\u23a5 \u23a7 \u23a8 \u23aa \u23a9 \u23aa \u2212y y 1 0 0 (16) Equilibrium positions are calculated by numerical integration of Eq. (16) with an initial iterative value (X0, u0)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001538_20060906-3-it-2910.00121-Figure5-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001538_20060906-3-it-2910.00121-Figure5-1.png", + "caption": "Fig. 5 Side (left figure) and top views of wheel 1", + "texts": [], + "surrounding_texts": [ + "Articulated rovers will be used increasing in diverse applications such as terrestrial and planetary explorations (Schenker et al 2003), (Volpe 2003), forestry (Gonthier 1998), agriculture (Baerveldt 2002) mining industries (Cunningham et al 1998), defense applications and hazardous material handling and de-mining (DeBolt et al 1997). Rovers with active suspension mechanisms are capable of modifying their configurations by adjusting their suspension linkages and joints so as to change their center of mass, thus avoiding tipover while traversing rough and sloppy terrain. Rovers with adjustable suspension system have been considered by Sreenivasan and Waldron (1996) for a specific vehicle. More recently, Iagnemma and Dubowsky (2003) presented stability-based suspension control for a specific rover using an essentially geometric approach and performing a rather complex optimization procedure. In a previous paper (Tarokh and McDermott 2005), we developed full kinematics models of articulated rovers and provided analysis of these rovers. The purpose of the present paper is twofold. First, is to present an alternative method for kinematics modeling of articulated rovers, which is straightforward, appealing, and computationally efficient. Second is to develop an optimization technique and incorporate it within the kinematics formulation to achieve simultaneous motion and balance control for rovers with active suspension mechanisms. The developed methodology is applied to a rover for demonstration purposes. 2. KINEMATICS MODEL DEVELOPMENT We define an articulated rover with active suspension system (ARAS) as a wheeled mobile robot consisting of a main body connected to wheels via a set of linkages and joints that can be adjusted, some actively and some passively for keeping the rover balanced. The active linkages and joints have actuators through which their values can be controlled, whereas passive ones change their values to comply with the terrain topology. The goal of balance control is to determine the actuation quantities for balancing the rover to avoid tipover. We will ignore dynamic effects as the rover moves slowly over the terrain. In order to achieve balance control, we must determine the contributions of each wheel and suspension mechanism to the overall motion of the rover. We attach a number of frames starting from the wheel-terrain contact frame then going through the steering and suspension frames and finally to the rover reference frame. Since we are interested in the motion, we relate the translational and rotational velocities of the next frame in terms of the previous frame. Let T aaaa zyxu ][= and T bbbb zyxu ][= denote the position of the current and next frames, respectively. Similarly, let T aaaa ][ \u03b3\u03b2\u03b1\u03d5 = and T bbbb ][ \u03b3\u03b2\u03b1\u03d5 = be the orientation of the current and next frames, respectively, where \u03b2\u03b1 , and \u03b3 are the rotation around x, y and z axis, or pitch, roll and yaw, respectively . The 13\u00d7 translation velocity vector of the next frame b is dependent on the translational and rotational velocities of the current frame a plus any translational velocity added to the frame b itself. This can be written as (Craig 2005) ( ) obbaaabb upuRu &&&& +\u00d7+= \u03d5, (1) where abR , and bp are, respectively, the rotation matrix and position vector of the frame b relative to the frame a, and obu& is the translational velocity added to the frame b. The latter is zero if the joint associated with the frame b is not prismatic. The rotational velocity of the next frame b is dependent on the rotational velocity of the frame a plus any rotational velocity ob\u03d5& added to the frame b itself, i.e. (Craig 2005) obaabb R \u03d5\u03d5\u03d5 &&& += , (2) We start at wheel i ( ni ,,2,1 L= ) contact frame ic which has the translational and rotational velocities T cicicici ]zyx[u &&&& = and T cicicici ][ \u03b3\u03b2\u03b1\u03d5 &&&& = , and perform the frame to frame velocity propagation until we reach to the rover reference frame to obtain rover velocities T rrrr ]zyx[u &&&& = and T rrrr ][ \u03b3\u03b2\u03b1\u03d5 &&&& = . Let the joint variable vector that includes each wheel-terrain contact angle, steering angle, and various prismatic and revolute joint variables be denoted by the 1\u00d7i\u03bd vector i\u03b7 . Then we will obtain an equation of the general form \u239f \u239f \u239f \u23a0 \u239e \u239c \u239c \u239c \u239d \u239b =\u239f\u239f \u23a0 \u239e \u239c\u239c \u239d \u239b ci ci ci i r r u J u \u03b7 \u03d5 \u03d5 & & & & & ; ni ,,2,1 L= (3) where iJ is the Jacobian matrix of the wheel i. Note that the wheel translational and rotational velocity vectors ciu& and ci\u03d5& include various slips. For example irolliici rx \u2212+= \u03b6\u03b8 &&& where ir is the radius of wheel i, i\u03b8& is the angular velocity of that wheel, and iroll\u2212\u03b6& is the rolling slip rate. Similarly ciy& and ciz& can, respectively, have side slip iside\u2212\u03b6& and bounce ibnce\u2212\u03b6& (up and down off the terrain movement) components. In addition, ci\u03b1& , ci\u03b2& and ci\u03b3& can be associated, respectively, with tilt itilt\u2212\u03b6& , sway isway\u2212\u03b6& and turn iturn\u2212\u03b6& slip rate components. In practice some of these slip components are unnecessary due to terrain topology and surface conditions, the path to be traversed (e.g. straight, serpentine, wavy) and mechanical arrangement of the wheels and suspension system. Equation (3) describes the contribution of individual wheel motion and the connecting joints to the rover body motion. The net body motion is the composite effect of all wheels and can be obtained by combining (3) into a single matrix equation as \u239f \u239f \u239f \u23a0 \u239e \u239c \u239c \u239c \u239d \u239b =\u239f\u239f \u23a0 \u239e \u239c\u239c \u239d \u239b \u239f \u239f \u239f \u23a0 \u239e \u239c \u239c \u239c \u239d \u239b \u03b7 \u03d5 \u03d5 & & & & & M c c r r 6 6 u J u I I (4) where the composite identity matrix on the left is nn\u00d76 , ( )Tcn2c1cc uuuu &L&&& = and ( )Tcn2c1cc \u03d5\u03d5\u03d5\u03d5 &L&&& = are 1n3 \u00d7 vectors of composite wheel velocities at the contact points, and \u03b7& is the 1\u00d7\u03bd vector of the joint variables which has both active (actuated) and passive joints. Note that in general some wheels share common suspension links and joints so that \u2211 = \u03bd\u2264\u03bd n 1i i . The composite Jacobian matrix of the rover J has a dimension of )6(6 \u03bd+\u00d7 nn . 2.1 Example The articulated rover with active suspension (ARAS) to be considered here is similar to the JPL Sample Return Rover shown in Fig.1. The schematic diagram of ARAS to be analyzed is shown in Fig. 2. The rover has four wheels with each independently actuated and rotation angles subscripted with a clockwise direction so that 1\u03b8 , 4\u03b8 are for the left side and 32, \u03b8\u03b8 are for the right side. At either side of the rover, two legs are connected via an adjustable hip joint. In Fig. 2 the hip angles on the left and right sides are denoted as 12\u03c3 and 22\u03c3 , respectively. These joints are actuated and used for balancing the rover. The two hips are connected to the body via a differential which has an angle \u03c1 on the left side and \u03c1\u2212 on the right side. On a flat surface \u03c1 is zero but becomes non-zero when one side moves up or down with respect to the other side. The differential joint \u03c1 is passive (unactuated) and provides for the compliance with the terrain. The wheels are steerable with steering angles denoted by i\u03c8 . The wheel terrain contact angle i\u03b4 is the angle between the z-axes of the i-th wheel axle frame Ai and contact coordinate frame ic as shown in Fig. 3. In order to derive the kinematics equations, we must assign coordinates frames. Fig. 4 illustrates our choice of coordinate frames for the left side of the rover. The right side is assigned similar frames. In Fig. 4, R is the rover reference frame whose origin is located on the center of gravity of the rover, its x-axis along the rover straight line forward motion, its y-axis across the rover body and its zaxis represents the up and down motion. The differential frame D has a vertical (along z-axis) offset denoted by 1k and a horizontal distance of 2k from R. The distance from the differential to the hip, denoted by 3k , is half the width of the rover. We now introduce three more frames, all of which have origin at the wheel axle. The length of the legs from the hip to the wheel axle is 4k . The hip frames 41 H,,H L for the four wheels are obtained from the differential frame by rotation and translation as shown with the Denavit-Hartenberg (D-H) parameters dhdhdh ad ,,\u03b3 and dh\u03b1 in Table 1 and in Fig 3. Similarly the steering frames 41 ,, SS L and axle frames 41 A,,A L are defined in Table 1 and Fig 3. We must now use the basic frame to frame equations (1)-(2) and go through the frames sequentially from wheel i terrain contact ic , wheel axle iA , steering iS , hip iH , differential D, and finally to the rover reference R. Equation (1)-(2) for the contact to the axle becomes ( )( ) ( )Ticici,AiAi T cicici,AiAi 00R r00uRu \u03b4\u03d5\u03d5 \u03d5 &&& &&& \u2212+= \u00d7+= (5) where the rotation matrix is \u239f \u239f \u239f \u23a0 \u239e \u239c \u239c \u239c \u239d \u239b \u2212 = cici cici ci,Ai c0s 010 s0c R \u03b4\u03b4 \u03b4\u03b4 , as evident from Fig. 3. Next we form wheel i axle to steering velocity propagation as ( )( ) ( )TiAiAiSiSi T AiAiAiSiSi R uRu \u03c8\u03d5\u03d5 \u03d5 &&& &&& \u2212+= \u00d7+= 00 000 , , (6) where \u239f \u239f \u239f \u23a0 \u239e \u239c \u239c \u239c \u239d \u239b \u2212 = 100 0 0 , ii ii AiSi cs sc R \u03c8\u03c8 \u03c8\u03c8 . The next in the chain is the hip frame, and we can write ( )( ) ( )TiiSiSiHiHi T SiSiSiHiHi hR uRu \u03c3\u03d5\u03d5 \u03d5 &&& &&& \u2212+= \u00d7+= 00 000 , , (7) with \u239f \u239f \u239f \u23a0 \u239e \u239c \u239c \u239c \u239d \u239b \u2212 \u2212 = 010 )(0)( )(0)( , iiii iiii SiHi hshc hchs R \u03c3\u03c3 \u03c3\u03c3 , 14 \u03c3\u03c3 = , 23 \u03c3=\u03c3 , and 4,3 2,1 1 1 = = \u23a9 \u23a8 \u23a7 \u2212 = i i hi . The differential frame velocities are obtained from (1)-(2) and Table 1 as ( )( ) ( )TiiiiSiHi,DiDi T 3i4HiHiHi,DiDi bh00R kb0kuRu \u03c1\u03c3\u03d5\u03d5 \u03d5 &&&& &&& ++= \u2212\u00d7+= (8) where \u239f \u239f \u239f \u23a0 \u239e \u239c \u239c \u239c \u239d \u239b ++ +\u2212+ = 100 0)()( 0)()( , iiiiiiii iiiiiiii HiDi bhsbhc bhcbhs R \u03c1\u03c3\u03c1\u03c3 \u03c1\u03c3\u03c1\u03c3 , \u03c1\u03c1\u03c1 == 14 , \u03c1\u03c1\u03c1 \u2212== 32 , and 3,2 4,1 1 1 = = \u23a9 \u23a8 \u23a7\u2212 = i i bi . Finally, the rover velocities are obtained as ( )( ) ( )TDiDirr T DiDiDirr R kkuRu 000 01 , 2, += \u2212\u2212\u00d7+= \u03d5\u03d5 \u03d5 && &&& (9) Substituting recursively (5) through (8) into (9) we obtain an equation of the form (3) where ( )Tciiiii \u03b4\u03c8\u03c3\u03c1\u03b7 &&&&& = . Due to space limitation, the Jacobian matrices iJ and their elements are not given here but can be found in our technical report (Mireles, et al 2005). The elements of iJ are trigonometric functions of the joint variables iii \u03c8\u03c3\u03c1 ,, and ci\u03b4 . 3. BALANCE CRITERION AND CONTROL An active suspension system is used to operate the rover to achieve balanced rover configurations such that when the rover traverses on a slope or rough terrain, tipover is prevented. We must now define and quantify more precisely the notion of a balanced configuration and express it in terms of rover orientation angles and adjustable joint angles. To this end, we use wheel-terrain contact position vectors iu , which represent vectors drawn from the rover reference point to the wheel-terrain contact point. Each consecutive pair of these vectors (i.e., iu and 1+iu ) form a plane denoted by i\u03c0 . The unit vector perpendicular to this plane is given by 1 1 + + \u00d7 \u00d7 = ii ii i uu uu s ; ni ,,1L= ; 11 uun =+ (11) Assuming that the rover reference frame R is at the center of rover mass, the rover unit gravity vector g can be expressed in terms of pitch and roll angles as ( )Tyxyxy cccssg \u03c6\u03c6\u03c6\u03c6\u03c6 \u2212\u2212= (12) Now we define the balance measure as the dot product between unit vectors is and g , i.e. i T i sg=\u03bc (13) Higher value of i\u03bc represents a more balanced rover. When the gravity vector g lies in any of the planes i\u03c0 , the vectors g and is become orthogonal, resulting in 0=i\u03bc . Tipover occurs when 0i <\u03bc . We must now define an objective function whose optimization results in a balanced configuration. Consider minimization of an objective function of the form \u220f\u2212+\u2212= = n i iraa aaaf 1 3 2 21 \u02c6 \u03bc\u03b2\u03b7\u03b7 (14) where a\u03b7 is the vector of actuated suspension joints and a\u03b7\u0302 is the nominal or desired values under normal operating conditions (e.g., flat surface). The second term represents the rover roll which must be minimized to keep the rover body level. The product term is the tipover measures which must be maximized, hence the negative sign. Note that without the first term, minimizing f would result in a rover configuration that is maximally flat or spread out even when the rover moves over a flat surface. The weighting factors 21,aa and 3a place relative emphasis between achieving rover balancing and the desire to operate near the nominal configuration. The balance and motion control problem may be stated as follows. Given the desired rover forward speed dx& and heading d\u03b3 , determine the commands to the wheel, and actuated joints, which include the steering, such that the rover maintains the desired forward speed and heading while minimizing the balance criterion (14). The composite equation (4) reflects the contribution of various position and angular rates to the overall motion of the rover. In order to control the rover motion while maintaining the rover balance, we must determine commands to the wheels, steering and joints actuators. For this, we rearrange (4) into an equation of the form qBA && =\u03c7 (15) where \u03c7& is the 1\u00d7xn vector of unknown quantities to be determined, and q& is the 1\u00d7qn vector of known quantities. The unknown vector consists of actuation signals such as active suspension joints, wheel roll rates, and un-measurable quantities such as wheel-terrain contact angles and appropriate slips. The known vector consists of desired quantities such as the desired forward rover velocity dx& and heading d\u03b3& as well as sensed quantities such as pitch and roll rates \u03b1& and \u03b2& , and rocker angle rate \u03c1& . The matrices A and B are obtained from the elements of J and the identity matrices 621 ,,, III L in (4). After partitioning (4) into the form (15), the dimensions of A and B are xnn \u00d76 and qnn\u00d76 . In order to be able to solve (15) while minimizing (14), we must have an underdetermined system of equations, so that the null space of A can be used for optimization. In this case we can solve (15) subject to minimization of (12) as (Nakamura 1991) \u239f\u239f \u23a0 \u239e \u239c\u239c \u239d \u239b \u2202\u2202 \u2212\u2212= 0 / )( ## \u03c3 \u03c7 f AAEkqBA && (16) where #A is the pseudo-inverse of A, k is a scalar, E is xx nn \u00d7 identity matrix, \u03c3\u2202\u2202 /F is the 1\u00d7\u03c3n vector of the gradient of the performance function with respect to the active suspension joints \u03c3 , and the zero vector had dimension 1)( \u00d7\u2212 \u03c3nnx . In the next section we specify the above quantities for our ARAS. The gradient can be computed numerically or analytically from (14)." + ] + }, + { + "image_filename": "designv11_69_0000118_s0043-1648(03)00330-2-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000118_s0043-1648(03)00330-2-Figure2-1.png", + "caption": "Fig. 2. The brush model. P: normal load, Qx: tangential load, R: equivalent contact radius, qx: tangential traction, \u00b5p: limiting tangential traction.", + "texts": [ + " Using the equations of motion, vB,1 and vB,2 can be expressed in terms of s1 and s2: vB,1 = (s\u03071 + RBs\u03072)n1 (8) vB,2 = (s\u03071 \u2212 RBs\u03072)n1 (9) The sliding velocities in the upper and lower contacts are then given by vsl,1 = (vA \u2212 (s\u03071 + RBs\u03072))n1 (10) vsl,2 = (vC \u2212 (s\u03071 \u2212 RBs\u03072))n1 (11) The local displacements between the surfaces in each contact can be calculated from the sliding velocities. The local displacements are denoted by \u03b4x,1 and \u03b4x,2 for the upper and lower contact, respectively. \u03b4x,1 = \u222b vsl,1 dt (12) \u03b4x,2 = \u222b vsl,2 dt (13) where vsl,1 = vsl,1n1 and vsl,2 = vsl,2n1. The brush model is illustrated in Fig. 2 (see [7] for a fuller description). Imagine that one of the two surfaces in contact has bristles on it. The bristles will bend and give rise to traction when the opposite surface moves relative to them. Of course, real surfaces have no bristles, but the imaginary bristles represent the elastic deformation of the surfaces in contact. A characteristic of the brush model is that each bristle deforms independently of the others. According to Eq. (14), the tangential traction qx at a point in the contact is proportional to the deformation ux of the bristle at that point, where Kt represents the tangential stiffness of the surface" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002301_0020-7403(67)90046-x-Figure5-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002301_0020-7403(67)90046-x-Figure5-1.png", + "caption": "FIG. 5. A n inadmiss ib le cross-sect ion.", + "texts": [ + " Th i s is a possible n e u t r a l axis , a n d t he cen t ro id GT of t he a r ea on t h e tens i le side of i t is u n i q u e l y def ined as t he p o i n t [~(0), ~(O)]. The c o n t i n u i t y of t he cu rve t r a c e d b y GT as 0 var ies is a s su red b y (5) a n d (6), a n d i ts non - r eve r s ing c h a r a c t e r follows because a t e v e r y pos i t ion d~/dO = 0 a n d d~./dO > 0. The p roo f of (5) a n d (6) is ba sed on Fig. 3, a n d on ly requi res t h a t t h e sectors dO shal l n o t be e m p t y of ma te r i a l . T h u s cross-sect ions such as t h a t s h o w n in Fig. 5 m u s t be e l imina ted , a n d t he t h e o r y r e s t r i c t ed to sec t ions of f ini te w i d t h in all d i rec t ions . As 8 var ies t h r o u g h 360 \u00b0 GT will t he re fo re t r ace a c o n t i n u o u s cu rve a n d r e t u r n to i ts or ig inal pos i t ion . I n do ing so i t c an m a k e on ly one c i rcu i t a r o u n d G, for i t m u s t a t all t imes lie on t he fa r side of G f rom Go. T he rad ius vec to r r is t h u s a s ing le -va lued func t i on of O. The p r o p e r t y of c o n v e x i t y p r o v e d above , a n d t he d e r i v a t i o n f rom t h e locus of t he d i rec t ion of t h e n e u t r a l axis , is obv ious ly a k i n to t h e p rope r t i e s of y ie ld surfaces" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000552_robot.1995.526030-Figure6-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000552_robot.1995.526030-Figure6-1.png", + "caption": "Figure 6: Examples of the second type of singularity for the 3RRR manipulator in which the three vectors ri intersect a t a point", + "texts": [ + " The first occurs when the three vectors ri are parallel. Therefore, the second and third columns of K are linearly dependent, aiid the nullspace of K represents the set of pure translations of M along a direction normal to ri, indicated by vector U of Fig. 5. The platform M can move along the direction of 11 even if the actuators are locked; likewise, a force applied to M in that direction cannot be balanced by the actuators. The second case in which K is singular occurs when the three vectors ri intersect a t a common point D , as shown in Fig. 6. Then, the nullspace of K represents the set of pure rotations of M about the common int,ersection point D . The platform M can rotate about that point even if the actuators are locked; likewise, a moment applied to JW cannot be balanced by the actuators. The third type of singularity occurs when the de- terminants of J and I( both vanish. We have this type of singularity whenever the three vectors r; are either parallel or concurrent a t a commoii point and a t least one leg is fully extended or fully folded, such that none of the rows of K vanishes" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000286_20.877817-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000286_20.877817-Figure1-1.png", + "caption": "Fig. 1. Spindle motor configuration with 8 poles and 12 slots.", + "texts": [ + " Kwon are with Department of Electrical Engineering, Hanyang University, Sa-1 Dong, Ansan 425-791, Republic of Korea (e-mail: freeyoon@elecma.hanyang.ac.kr; bikwon@email.hanyang.ac.kr). H.-S. Yoon and S.-H. Won are with the Research & Development Center, Samsung Electromechanics Co., LTD, Maetan-3 Dong, Suwon 442-743, Republic of Korea, (e-mail: {m0hsyoon; shwon}@samsung.co.kr). Publisher Item Identifier S 0018-9464(00)07184-3. Based on the measurements, it appears that the permanent magnet pole magnetization levels in the spindle motor illustrated in Fig. 1 are different from one pole to another. To examine the spindle motor characteristics according to the variation of the magnetic parameter, situations as shown in Fig. 2 are defined. The symmetric magnetization distribution in Fig. 2, which is expressed as the dotted line, is obtained from the finite element analysis for magnetizing process [4]. At this point we consider the unsymmetric magnetization distribution represented by the solid line in Fig. 2 supposing that the 0018\u20139464/00$10.00 \u00a9 2000 IEEE Table I shows an example of the peak magnetization levels in each pole of the permanent magnet when the ring-type magnet is magnetized in the eccentric state. The permanent magnet material is a bonded Nd\u2013Fe\u2013B. The approach shown in Table I corresponds to a variation of the magnetization level, maximum for the symmetric magnetization distribution. The 2-dimensional governing equation for the permanent magnet BLDC motor with the exterior rotor steel shell shown in Fig. 1 is expressed in a magnetic vector potential as (1) where stands for permeability, for coil turns, for the cross section area of the winding with a phase current and for magnetization of the permanent magnet. is an electrical scalar potential defined at the rotor steel shell. The subscript , and represent each phase. In deriving (1), the uniqueness of the vector potential is forced by the Coulomb gauge condition, . Since we assume that all currents exist in the axial direction only for two dimensional analysis, the conductivity of the rotor steel shell, in (1), is corrected to take account of the transverse edge effect due to a finite length of the steel shell [5]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000002_0954406011520742-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000002_0954406011520742-Figure1-1.png", + "caption": "Fig. 1 Experimental set-up", + "texts": [ + " Since the geometrical features of the molten pool can be used as an eYcient feedback to estimate the molten depth, a model is built to explore the correlation between molten depth, molten pool geometry and processing parameters (laser power and travel speed). Finally, residual stresses predicted by the model were compared with those obtained experimentally by X-ray analysis of the nitrided surfaces. The issues related to the reduction of residual stresses are addressed. The experimental set-up which consists of an Nd:YAG laser, a two-axis computer numerically controlled (CNC) positioning system, a laser strobe vision system, shielding gas and the workpiece is shown in Fig. 1. Fibre optics conduct a laser light of 337 nm wavelength to illuminate the Nd:YAG laser treatment area. The illuminating laser is a nitrogen pulse laser of 5 ns pulse duration, synchronizing with the high-speed shutter of the camera. The camera of the laser strobe vision system is equipped with a UV \u00ae lter that only allows light near a 337nm wavelength to pass. During the illumination period, the intensity of the illuminating laser can suppress the spatter and plasma light. In addition, owing to the re\u00af ection of the mirror-like molten pool, a wellcontrasted image of the molten pool can be obtained" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002200_09544054jem699-Figure7-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002200_09544054jem699-Figure7-1.png", + "caption": "Fig. 7 The integrated pressing and Y-axis scheme for the SSM system", + "texts": [ + " Therefore TP2 and TP5 do not interfere with each other in work scheduling. It is not necessary to set up in the same Y direction two motion elements (TPs) which do not conflict in scheduling; this shows that there is a redundant motion element in the Y direction and these two elements should be combined. Thus it is possible to combine these two motions together and to integrate two DPs into one, which can perform the dual functions for both the pressing motion and the Y motion. The integrated scheme is shown in Fig. 7. In the integrated scheme, TP2 and TP5 are combined into one parameter TP0 2 as follows FR2 FR5 TP0 2 Y-axis and pressing motion: distance, 700mm; accuracy, 0.01mm The integrated control desired value M0 is M0 \u00bc M ITP2 ITP5 \u00fe ITP2 0 \u00bc 71.73. The integrated control efficiency h0 is h0 \u00bc IR/M 0 \u00bc 1.565, which is larger than that of the first design scheme. This indicates that the SSM system is simpler and of lower cost in the integrated design scheme than in the first design scheme without affecting the system precision and working efficiency" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001285_cacsd-cca-isic.2006.4776982-Figure3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001285_cacsd-cca-isic.2006.4776982-Figure3-1.png", + "caption": "Fig. 3. Diagram of the Convertible Birotor", + "texts": [ + " In vertical mode the vehicle depends as much the rotors as the control surfaces (aileron, elevator). The vehicle\u2019s up-down motion is controlled by increasing or decreasing the propeller thrust. The rolling motion is controlled via the difference in the rotors\u2019 angular velocity. Since the control surfaces are submerged in the propeller slipstream (prop-wash), we take advantage of this to generate aerodynamic forces with 0-7803-9796-7/06/$20.00 \u00a92006 IEEE 2202 the elevator and ailerons deflection to provide the pitch and yaw motion respectively [see figure 3]. Let I={iIx , jIy , kI z } denotes the right hand inertial frame stationary. Let B={iBx , jBy , kB z } denotes the rigid-body frame which has as origin the center of mass. Let the vector q = (\u03be, \u03b7)T denotes the generalized coordinates where \u03be = (x, y, z)T \u2208 3 denotes translation coordinates relative to the frame I, and \u03b7 = (\u03c8, \u03b8, \u03c6)T \u2208 3 describes the vehicle orientation by the classical yaw, pitch and roll angles (Euler angles) commonly used in aerodynamics applications [1]. The orientation of the convertible birotor is given by the rotation matrix1 RB\u2192I " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001652_icar.2005.1507413-Figure11-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001652_icar.2005.1507413-Figure11-1.png", + "caption": "Fig. 11. Method to transfigure the mechanism by rotating the end plate", + "texts": [ + " Rotational workspace of normal mechanism is found to be larger than most of the transfigured mechanisms by the parallel transformation for the same \u0398 and d. This is because when we shift the endplate to transfigure the mechanism by the parallel transformation, both the link length and angle between the link and the endplate varies. This causes interference in the links limiting the rotational workspace of transfigured mechanism by the parallel transformation. Rotating transformation is done by rotating the endplate about one of the principle axis and adjusting the link lengths to hold this position (Fig.11). In this paper, we discuss the rotation \u03b1 about x-axis only. 1) Translational workspace: To calculate the translational workspace of normal mechanism, we first rotate the endplate by certain angle using actuators and then moved the endplate along x, y and z-axes keeping the angle constant. For the transfigured mechanism by the rotating transformation, we rotated the endplate by the same angle as in the normal mechanism by adjusting link lengths, and then moved the endplate along x, y and z-axes to find its translational workspace" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001183_20050703-6-cz-1902.01279-Figure7-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001183_20050703-6-cz-1902.01279-Figure7-1.png", + "caption": "Fig. 7. Collision free path construction using the tangent at the approximated obstacle", + "texts": [ + " This new configuration can be added to the original path segment in order to create a collision free path around the point-like obstacle. In many cases the path constructed in a way described above is already collision free corresponding to the approximation of the original obstacle. But in other cases this path is still colliding with the point-like obstacle. In these cases in addition to the collision free configuration with the scaling value of \u201cone\u201d (see above) a tangent at the curve of the approximated obstacle in the configuration space should be taken into account. Figure 7 gives an example. Here the path consisting of the configurations start, q1 and goal is still colliding with the obstacle. Therefore a tangent at the obstacle\u2019s curve in the configuration qt was determined. The configuration q1 was replaced by q2 which is the point of intersection between the two lines start-qt and goal-q1. The resulting path is no more colliding with the point-like obstacle. To determine the configuration qt which is needed for the tangent described above the second part of equation (5) can be used" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003661_ijmtm.2010.031368-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003661_ijmtm.2010.031368-Figure1-1.png", + "caption": "Figure 1 Mechatronics Design Methodology framework (see online version for colours)", + "texts": [ + " The following work also contributes to presenting a novel way to integrate existing specialised techniques and tools that were not specifically designed to produce a mechatronic product in addition to depicting a relationship between students, university staff, experts and manufacturing companies to carry out a project or such discipline. There is a physical and organisational infrastructure to support the MDM, which will be referred to as the MDM Framework. It enables the development of the project through logistic activities and the MDM in a PLM environment (Guerra et al., 2005). These are shown from outside to inside in Figure 1 and will be explained with further detail in their respective sections. The activities embraced by PLM, Project Logistics and Project Coordination will be explained first as a way to give a general view on how to implement the MDM framework properly. The life cycle of a product encompasses the holistic view of its entire development. During the life cycle of a product, several phases can be identified: product design, process development, production, distribution, use and disposal. These are shown at the outer part of Figure 1. All the phases must be taken into account during the development of products. PLM is a way to integrate the information and knowledge produced during this process to improve and increase innovation for a given product (Turban, 2005). Every phase covered by the PLM approach can be applied when designing a mechatronic product. Product design is the first stage to be addressed by PLM (Aca, 2003). It includes hearing the customer\u2019s voice, process identification and an up-front product definition. Identification of clients\u2019 requirements and innovative ideas generation are the steps included in this stage of the MDM to help the students and experts to propose a better project planning", + " The final PLM stage comes when the product has fulfilled its mission and disposal becomes necessary. Recycling concept becomes very important due to the global concern about sustainable development (Schey, 2000). The physical and organisational infrastructure has been considered as being divided by two counterparts, which will be called university and industry. The first stands for the academia and the second are the manufacturing companies. Figure 2 contains the concepts and relationships in what this methodology are based. The elements of this figure could be placed in the centre of Figure 1, but they have been separated for the sake of good visualisation and understanding. There is a need for a tight relationship between university and industry, which is generated by MDM experts. MDM experts need to make contact with companies to develop a plan for the project. In this first contact, industry provides the mechatronic projects that university is going to develop. Besides, both of them establish an agreement on how they will pursue a common objective. Then, the university participants are granted access to a manufacturing facility and a person from it is assigned as their aid" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002979_978-3-540-70534-5_7-Figure13.11-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002979_978-3-540-70534-5_7-Figure13.11-1.png", + "caption": "Figure 13.11: USAL thrusters and rudder turning maneuver [Drtil 2006]", + "texts": [ + " These include a digital camera, four analog PSD infrared distance sensors, a digital compass, a three-axes solid state accelerometer and a depth pressure sensor. The Bluetooth wireless communication system can only be used when the AUV has surfaced or is diving close to the surface. The energy control subsystem contains voltage regulators and level converters, additional voltage and leakage sensors, as well as motor drivers for the stern main driving motor, the rudder servo, the diving trolling motor, and the bow thruster pump. Figure 13.11 shows the arrangement of the three thrusters and the stern rudder, together with a typical turning movement of the USAL. Autonomous Vessels and Underwater Vehicles 204 13 13.5 References ALFIREVICH, E. Depth and Position Sensing for an Autonomous Underwater Vehicle, B.E. Honours Thesis, The University of Western Australia, Electrical and Computer Eng., supervised by T. Br\u00e4unl, 2005 AUVSI, AUVSI and ONR's 9th International Autonomous Underwater Vehicle Competition, Association for Unmanned Vehicle Systems International, http://www" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001309_6.2006-6599-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001309_6.2006-6599-Figure1-1.png", + "caption": "Figure 1. Localizer and Glideslope Tracking Geometry21", + "texts": [ + " 3 of 29 American Institute of Aeronautics and Astronautics The purpose of this section is to describe the approach and landing problem posed for this research. As noted in Section I, the automatic landing consists of intercepting a lateral and vertical beam and tracking the guidance provided to a specified height above the runway, where a flare maneuver is performed. It is assumed that a guidance system is available to provide lateral and vertical guidance to the start of the flare. The control laws developed are independent of the type of approach system used; however, evaluating different approach types is beyond the scope of this paper. Figure 1 show the geometry used to determine deviations from the lateral and vertical beam, and Figure 2 shows the geometry of the flare maneuver. A Category (CAT) III Instrument Landing System (ILS) is assumed for this paper, although the techniques can easily be extended to any guidance system, which provides precision guidance data. Since ILS is assumed, the lateral beam will be referred to as the localizer, and the vertical beam will be referred to as the glideslope. The localizer consists of a transmitter stationed at the far end of the runway which sends out a signal that is approximately 5 deg wide (beamwidth) and is centered on the runway centerline" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000694_b:tels.0000029042.75697.f0-Figure21-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000694_b:tels.0000029042.75697.f0-Figure21-1.png", + "caption": "Figure 21. The topology of the 4 node DDF network when demonstrating the DDF algorithm in real-time. While the aircraft nodes are each connected to a camera, the ground nodes have no local sensor at all.", + "texts": [ + " This illustrates the target map gener- ated on aircraft 1, and includes a zoomed in section of one part of the map. As a result of the communication of information in the DDF network, the maps on all aircraft are identical to that presented here and are therefore omitted for brevity. Results are now presented of a real-time demonstration of the algorithm. The system consisted of two aircraft equipped with vision payloads and two ground nodes with no sensor attached. The topology of this DDF network is illustrated in figure 21. As the ground nodes have no sensor attached, they do not contribute any information to the network. They simply receive DDF information from the aircraft they are connected to and use it to construct their own estimates. Thus, the ground nodes will replicate the estimates on the aircraft even though they do not have any sensors attached. The system was implemented in this way for two reasons: 1. To make the network more complex than just the two aircraft nodes. 2. As the ground nodes replicate the estimates on each aircraft, they are able to provide target information for a GUI without the need for the aircraft to communicate anything other than DDF information" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000411_6.2001-3480-Figure8-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000411_6.2001-3480-Figure8-1.png", + "caption": "Figure 8. Section through Test Housings", + "texts": [ + " This rapid speed change capability allows the rig to be driven like a gas turbine, the pressure response is as quick so it is only the temperature that tends to lag. The spindle is driven via a flat belt drive as indicated in Figure 7. Both spindles utilise angular contact bearings fitted with silicon nitride balls to enable high speeds to be achieved. The bearings are oil cooled and the nose of the disk spindle is also air cooled. Cross turbo charger sealing rings are used as the primary oil seal on both spindles and they have proved very reliable under all operating conditions. Figure 8 shows the housing arrangement used to simulate an aerospace application where rapid response was required. To further simulate the engine the disk was machined with typically 0.005\" of run-out and the air was bought in between the seals through a swirl inducer. The level of swirl is governed by the pressure ratio across the swirl inducer holes, these are easily changed to allow pressure ratios from fully choked down to low numbers to be achieved. When we are testing for industrial applications we tend to use much heavier test housings so as to give a more realistic thermal response to the application" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002979_978-3-540-70534-5_7-Figure16.4-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002979_978-3-540-70534-5_7-Figure16.4-1.png", + "caption": "Figure 16.4: Dead reckoning", + "texts": [ + " Dead reckoning In many cases, driving robots have to rely on their wheel encoders alone for short-term localization, and can update their position and orientation from time to time, for example when reaching a certain waypoint. So-called \u201cdead reckoning\u201d is the standard localization method under these circumstances. Dead reckoning is a nautical term from the 1700s when ships did not have modern navigation equipment and had to rely on vector-adding their course segments to establish their current position. Dead reckoning can be described as local polar coordinates, or more practically as turtle graphics geometry. As can be seen in Figure 16.4, it is required to know the robot\u2019s starting position and orientation. For all subsequent driving actions (for example straight sections or rotations on the spot or curves), Figure 16.2: Beacon measurements green beacon red beacon 45\u00b0 165\u00b0 Localization and Navigation 244 16 the robot\u2019s current position is updated as per the feedback provided from the wheel encoders. Obviously this method has severe limitations when applied for a longer time. All inaccuracies due to sensor error or wheel slippage will add up over time" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002791_pesc.2007.4342355-Figure11-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002791_pesc.2007.4342355-Figure11-1.png", + "caption": "Figure 11. High Frequency loop in physical structure.", + "texts": [ + " By the same analysis, the HFL occurs between the upside equivalent capacitor Ceql and underside IGBT or between the upside FRD and underside equivalent capacitor Cequal, when the phase-leg output current is negative, as illustrated in figure 9. Because the parasitic inductance has effect on the ringing frequency and over-shoot voltage [5], the inductance value should be decreased unceasingly in the IPEM interconnection process to reduce the switching time and switching losses. The phase-leg packaging structure with 34.8\u00d728.45\u00d76.4 mm3 volume is shown in figure 10. And figure 11 illustrates the HFL in the phase-leg physical structure. Due to the symmetrical packaging structure in the phase-leg, the HFL impedance characteristic shown in figure 11 (a) is the same as that of HFL shown in figure 11 (b). Regardless of the contact resistance affect, the HFL packaging impedance of simulated result shown in figure 12 indicates that the loop parasitic inductance value decreases from 8.2nH/DC to 2.3nH/100MHz. The ringing frequency in HFL exceeds 100 kHz,so the AC inductance value in HFL approaches to constant value 2.3 nH. The loop parasitic resistance increase from 0.13 m /DC to 11.6 m /100MHz shown in figure 12 (a). So, during the steady switching states and switching process, the packaging loss evidently decreases comparing with the wire-bonding technology" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000380_iecon.1989.69715-Figure6-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000380_iecon.1989.69715-Figure6-1.png", + "caption": "Fig. 6 brockdiagram o f equipment We used the one chip high function analog a d t a g a t h e r i n g L.S.I. ADC83,B f o r t h e A/D c o n v e r t e r , and t o s u p l y t h e moter s i g n a l from A D 6 6 7 J E D/A c o n v e r t e r . i c h i n g V e l o c i t y Methpd", + "texts": [ + " Suppose t h e c o t r o l l e r c n s i s t s o f t h r e e r u l e s . i f S i i s ZE and S, i s ZE then U i s u n i . i f S f i s ZB and Sz i s PS t hen U i s dec . if S I i s PB and Sz i s ZE t hen U i s acc . Thus, f o r a g iven i n p u t , t h e c o n t r o l ai- gor i thms w i l l check i f t h e r e e x i s t s a cor r e spond ing r u l e s . 3 4 THE EXPERIMENTAL 2ESULTS We show t h e expa r imen ta l a p p a r a t u s o f t h e c rane model i n F ig .5 . And we i l lustrate t h e brockdiagram of t h e c o n t r o l p a r t i n Fig.6 . We measure t h e a n g l e of t h e pen- dulum by t h e senso r . We used t h e potent i o m e t e r made of t h e magnet ic r e s i s t o r f o r t h e s e n s o r of t h e a n g l e of t h e pendulum. Fuzzy C o n t r o l Method v - . . . . . . 8.8 1.0 2.0 3 8 4 8 5 0 6 0 7'0 8'0 3 .+T time Fig.5 scene o f exper imenta l d e v i c e Fig.1 t h e case of normal o p e r a t i o n d u r i n g t h e t i m e 'T4 , t h e v e h i c l e r u n s a t Vmax,and t h e a n g l e o f t h e pendulum i s zezo" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002313_1.2827448-Figure4\u20136-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002313_1.2827448-Figure4\u20136-1.png", + "caption": "Figures 4\u20136 show the corresponding computed 3D self-", + "texts": [], + "surrounding_texts": [ + "t o e\nb a\n7 M\nm s\n8\nr T m l\nh e o\nt m i l\ns o s a T i o m\nw a = o h f\nl\n0\nDownloaded Fr\nMoments of friction forces may be found by integrating tangenial stress with respect to opposite surfaces of substructures. Thus, n the surface of substructure k =0 moment of friction forces is xpressed by\nMk =\nec Sec\nRp ds 17\nThe power of losses, caused by friction in the lubricant layer s etween substructures k and m during time period T of the process nalysis, is computed by the following formula:\nNs = 1\nT 0 T ec Fs ecTVec + Mk k \u2212 Mm m dt 18\nLinkup of the Finite Element and Finite-Difference eshes of the Rotor System Model The procedure provides the following method of the linkup RS odel of the finite element and finite-difference meshes, having\nignificantly different sizes:\na The values of clearances and velocities in nodes in finitedifference meshes of the layer models are found by interpolating the nodal values of the finite element meshes of the substructure\u2019s opposite surfaces based on the contact finite element form functions. b The integration of stresses in lubricant layers is implemented on finite-difference meshes by the Simpson\u2019s method. The reduction of forces for transposing them to nodes of finite element meshes of substructure models is carried out with the use of the contact finite element form functions.\nIntegration of Rotor System Motion Equations An integration method shall be stable numerically within a time ange sufficient for the identification of RS motion parameters. his requirement for rigid systems of differential equation 8 is et by the absolutely stable methods of a step-by-step integration 10 . The type and parameters of the integration method are seected based on the following conditions.\nFirst, due to the small integration steps in modeling of the RS igh-velocity dynamics, the stiffness component of the system\u2019s ffective matrix is suppressed numerically, causing the distortion f physical properties of a simulated object.\nSecondly, there is a high sensitivity of Reynolds equation soluions for pressures to oscillations of nodal velocities of the RS\nodel finite element meshes. Velocity oscillations deteriorate the terative convergence of the solutions for pressures in lubricant ayers, causing numerical instability of the integration process.\nThe solution of the indicated problems is achieved by the retriction of an integration time step t1 t t2 and by the use f integration methods with smoothing. The values t1 , t2 are elected depending on the values of rotation frequency of the rotor nd the highest frequency of the simulated RS vibration mode. he method of rectangles with smoothing velocities 10 exposed tself as an efficient method of integration in solving the problem f 3D RS dynamics. The basic relations for any RS substructure ay be expressed as\nDUt+ t i+1 = FPt + FHt+ t + FFt+ t 19\nhere i is the number of an iteration for an nonlinear problem on time step, D=auM+av G+K is the effective RS matrix, FPt M auUt+avVt + G avUt+Vt is the prognosticated component f the load vector, FHt+ t=FHt+ t Ut+ t i ,Vt+ t i is the nonlinear ydrodynamic component of the load vector, and FFt+ t is the orced component of the load vector.\nThe algorithm of smoothing velocities is assigned by the fol-\nowing formulas:\n31003-4 / Vol. 130, JUNE 2008\nom: http://vibrationacoustics.asmedigitalcollection.asme.org/ on 01/28/201\nVt+ t = av Ut+ t \u2212 Ut \u2212 Vt sm\nVt sm = Vt+ t i + 2Vt + Vt+ t /4 20 The condition of convergence within a time step is satisfied at the moment of equality of the components of the RS model nodal displacement residuals to relative values within the range of 10\u22123\u201310\u22122.\nIt should be mentioned that the effective system matrix D of Eq. 19 is nonsymmetrical due to the antisymmetry of the gyroscopic matrix G. All sparse matrices and vectors are assembled, stored, and processed in a row-wise format 16 . The assembly and decomposition of matrix D are performed only once, before integrating the motion equation.\n9 Software Implementation of the Procedure Simulation procedure of RS dynamics is oriented toward its application in engineering computations performed on a desktop PC. The software system has been developed on the FORTRAN POWER STATION 4, making it possible to implement the following:\na the selection of data on RS geometry, mechanical properties of materials used in parts, and physical properties of lubricants; b the assignment of loads and parameters of RS operation modes; c the finite element and finite difference discretization of the RS model; d the solution of vector equations on a step of numerical integration of RS motion equations; e the visualization of trajectories of RS parts\u2019 motion; f the registration of nodal displacements in the RS model,\nclearances, reactions, and losses in bearings; and g The construction of AFC of the simulated RS with the\nuse of tools of frequency and modal analyses.\n10 Simulation of the Turbocharger\u2019s Rotor System Dynamics\nThe application of the developed procedure to modeling the industrial turbocharger RS dynamics, aiming at the collation of the computed results with those obtained experimentally, is presented and discussed below.\nRotor System Design. The turbocharger\u2019s rotor is of 8 kg in mass and 300 mm in length, operating wheels are 210 mm in diameter, and diameters of the compressor and turbine bearings are 24 mm and 28 mm, correspondingly. The mentioned bearings of the turbocharger are shown in Fig. 2.\nRadial bearing\u2019s bushings 1 and 2 are fixed against rotation and set in dampers with squeezed film. The internal cylindrical carrying surface of bushings is furnished with four axial grooves for lubricant distribution. Oil supply into damper clearances is implemented out of ring-shaped grooves in bearing\u2019s housings.\nThe main axial bearing is formed by the flange of bushing 1 of the compressor bearing and thrust collar 4 on the rotor shaft. The auxiliary axial bearing is formed by floating disk 3 , thrust disk 5 , and housing 7 of the compressor\u2019s bearing. Gasket packet 6 in the oil clearance between the flange of bushing 1 and housing 7 of the compressor\u2019s bearing serves as an axial damper. The boundaries of coupled lubricant layers of the axial and radial dampers of the compressor\u2019s bearing come out into the ring-shaped internal cavity formed by housing 7 and bearing\u2018s bushing 1 . The summarized axial clearance of the rotor in bearings is equal to 0.2 mm.\nExperiment. Experiments were carried out within the range of rotor rotation frequencies 250\u2013670 Hz. Signals of inductive sensors of displacements, set in the vicinity of the compressor shaft end, were transmitted to an oscillograph, a spectrum analyzer, and\nTransactions of the ASME\n6 Terms of Use: http://www.asme.org/about-asme/terms-of-use", + "a a m w v c f b s i r t s p f t p\ns r b T l c b c\n= l o a s p c o\nw c t r s\nt\nJ\nDownloaded Fr\nmagnetograph. Temperatures of bushings, bearing\u2019s housings, nd oil temperature at the inlet and outlet of the bearings were easured by thermocouples. The obtained data on temperatures ere used in computations of thermal dilatation of parts and oil iscosity alteration. The results obtained for the values of relative learances absolute values related to diameters of carrying suraces equal to 0.006 in bearings and 0.005 in dampers are given elow. The value of relative clearance in bearings, 0.006, correponds to the upper limit of the permissible abrasion of bushings n the course of exploration. The RS instability for the mentioned elation of clearances is exposed in intensive self-sustained vibraions on oil film. The computation of the dynamics of such a ystem makes it possible from one side to reveal most fully the otential of the procedure of the numerical RS simulation, and rom the other side to formulate practical recommendations for he inspection of abrasion and substitution of outworn bearing arts.\nRotor System Model. The analyzed RS model comprises ten ubstructures, among them are a shaft with a turbine, mounted otor parts, bearing\u2019s bushings, and a floating disk. The total numer of the nodal degrees of freedom of the RS model was 7212. he number of nodes in model meshes of five radial and four axial ubricant layers was approximately 8 103. The summarized ophasal unbalance of the rotor was equal to 25 g mm, axial force eing within the range of 100\u2013500 N. The value of lubricant visosity at inlets of the bearings is 0.02\u20130.03 N s /m2.\nAverage step of the process discretization was equal to t 0.75 10\u22125 s, corresponding to about 300 steps per rotor revoution. The theoretical number of discretization steps within an scillation period, which guarantees a sufficient accuracy of the pplied integration method, is not less than 100 10 . As it was hown in computations, the assumed discretization parameters rovided necessary accuracy and quick four to eight iterations onvergence of integration process with respect to time equations f nonlinear RS motion.\nThe number of simulated rotor revolution at each point of AFC as within the range of 200\u2013500. Computation time required for\nalculating one rotor revolution using a personal computer Penium IV 1.7 MHz comes to about 30 s. The computational time equired for simulation of a stable synchronous motion for the ame mesh size is two to three times less.\nResults. The results of the processing of experimental data and\nhe results obtained by PC numerical simulation are demonstrated\nournal of Vibration and Acoustics\nom: http://vibrationacoustics.asmedigitalcollection.asme.org/ on 01/28/201\nin Figs. 3\u20136. The amplitude and frequency shown in Fig. 3 were obtained by the frequency analysis of the experimental and computed vertical displacements of the compressor end of the shaft. Trajectories of self-sustained rotor modes are also schematically shown in Fig. 3. Graphs a and b in Fig. 3 comply with the first and the second self-sustained rotor vibration modes, correspondingly. Graph c conforms to synchronous unbalanced oscillations of the rotor on oil film.\nFig. 5 Second self-sustained mode\nFig. 6 Synchronous mode\nJUNE 2008, Vol. 130 / 031003-5\n6 Terms of Use: http://www.asme.org/about-asme/terms-of-use", + "s f i a 1\nc e T 4 m t\nd s a f r f o\no i w e r n a\ns c w F r s m\nfi a F t fl i w\n2 b b\ns t t b r 9 t\n1\nv m p d a w t s\n0\nDownloaded Fr\nustained and synchronous RS modes. All modes comply with orward RS precessions. The mode normalization by amplitude is ndividual; therefore, the relation of displacement scales in Figs. 4 nd 5 to the displacement scale in Fig. 6 is approximately equal to :10.\nInitial clearances in bearings not shown in these figures are omparable by their value with summarized displacements and xceed synchronized RS displacements by the order of magnitude. his is explained by a fictitious overlapping of RS parts in Figs. \u20136, especially in Fig. 6. The applied graphic representation of the odal analysis results visually reveals amplitude and phase relaions of motion modes for individual RS parts. As it is seen from these figures, the analyzed RS presents a ouble-frequency, self-oscillating system. System properties subtantially alter, depending on rotation frequency. The computed nd experimental values of relative frequencies related to rotation requency fR coincide with accuracy of 5%, Fig 3. The values of elative frequencies are within the range of 0.23\u20130.42. Relative requencies of self-sustained vibrations diminish with the increase f rotor revolution velocity.\nThe first, lowest, self-sustained mode exists in the whole range f rotor revolutions. Within the frequency range of 250\u2013330 Hz, t manifests itself in the shape of a straight conical precession, ith the amplitude at the shaft ends 170\u2013180 m and the presnce of a node between bearings, Figs. 3 and 4. With increasing otation frequency, there takes place a displacement of the modal ode toward the compressor\u2019s bearing and an increase of the shaft xis curvature Fig. 4 .\nAt the rotor rotation frequency 396 Hz, the second selfustained mode of the rotor gets excited jumpwise, and beatings, haracteristic for a double-frequency oscillatory process, are atched. Initially, the shape of this mode is close to cylindrical, igs. 3 and 5. At the end of the investigated rotation frequency ange, at the frequency value 650\u2013670 Hz, the second mode hape becomes deflected with the node on the turbine shaft end Fig. 5 . The computed values of amplitudes of self-sustained odes comply satisfactorily with those obtained by experiment. Synchronous unbalanced vibrations appear in the shape of the rst bending mode of the shaft Fig. 6 and are damped down at ll rotor revolution frequencies. As it could be concluded from ig. 3, its amplitude does not exceed 10\u201320 m that is 8\u201315 imes less than the self-oscillation amplitude. The position of the oating disk outside the shaft shown in Fig. 6 is dummy, connectng with depicting only dynamic RS displacements in Fig. 6 as ell as the difference of disk\u2019s and shaft\u2019s vibration phases. The total levels of shaft end displacements are about 00\u2013220 m. Relative eccentricities of middle cross sections of ushings in dampers are equal to 0.75, of those of shaft journals in ushings, being equal to 0.65.\nThe highest bending stress\u201470 MPa\u2014takes place in the tepped transition to the compressor\u2019s end of the rotor shaft. Roation frequency of the floating thrust disk is approximately equal o half the rotor rotation frequency. The computed losses caused y friction at the rotation frequency 667 Hz are equal to 2.8 kW in adial bearings and 6.3 kW in axial ones, total losses coming to .1 kW. The value of total losses obtained experimentally, with he use of the calorimetric method, is equal to 8.4 kW.\n1 Conclusions Optimal design of the rotor systems requires reliable data on ibration frequencies, displacement, and stress amplitudes in the aterial of RS parts, clearances, and losses in bearings. The proosed simulation procedure solves the problem of obtaining such ata, using three-dimensional finite element and finite-difference nalyses of the RS, combined with the step-by-step integration ith respect to time of the motion equation, and the processing of he results obtained by the methods of frequency and modal analyes.\n31003-6 / Vol. 130, JUNE 2008\nom: http://vibrationacoustics.asmedigitalcollection.asme.org/ on 01/28/201\nThe distinguishing features of the proposed procedure are as follows:\na The use of the rotation elastic medium model for taking into consideration main dynamic effects of the RS. The equation obtained with the use of the developed model differs from traditional dynamic equations of the 3D FEM by the introduction of a gyroscopic matrix and a vector of centrifugal forces. b The application of the contact finite element technique for linking up the meshes of elastic and hydrodynamic parts of the RS model. c The application of finite-difference models for coupled lubricant layers of RS bearings and dampers. d Universality with regard to geometry, load, and operation modes of simulated RSs. e Completeness of the obtained information of frequencies and 3D modes of RS oscillations, stress-strain state of RS parts, reactions, clearances, and losses in bearings.\nNomenclature ared coefficient for reducing density and\nYoung\u2019s modulus of a blade material au, av coefficients of the method of inte-\ngrating with respect to time, correspondingly, 1 /s2, 1/s\nC matrix of directing cosines of the hydrodynamic contact surface\nCs contour, m D effective matrix of the system, N/m\nDx matrix differential operator, 1/m E= eEe, Ecf= eEcf e assembly of vectors Ee and Ecf\ne for a substructure, kg m\nE ,Ered initial and reduced Young\u2019s modulus of a blade material, correspondingly, N /m2\nER vector of coordinates of a medium point, m\nEcf e = meNT 0 y z Tdm vector connected with centrifugal\nforces applied to a finite element, kg m\nEe= meNTERdm vector, kg m F vector of nodal forces of a substruc-\nture, N FFt+ t forced component of vector of loads on a time step t+ t of integration, N FHt+ t hydrodynamic component of vector\nof loads on a time step t+ t of integration, N\nFPt prognosticated component of vector of loads on a time step t+ t of integration, N\nFs ec ,Fs hydrodynamic components of nodal\nforces\u2019 vector for the contact finite element and a substructure, correspondingly, N\nFg e ,Fun e vectors of gravitation and unbalance correspondingly of a finite element, N\nfs vector of stresses in a lubricant layer, N /m2\nfR rotor rotation frequency, Hz G= e Ge\u2212GeT\ngyroscopic matrix of a substructure, kg\nGR matrix of coordinates of a medium point, m\nGe= meNTGRDxNdm gyroscopic matrix of a finite element,\nkg\nTransactions of the ASME\n6 Terms of Use: http://www.asme.org/about-asme/terms-of-use" + ] + }, + { + "image_filename": "designv11_69_0002819_6.2007-3813-Figure28-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002819_6.2007-3813-Figure28-1.png", + "caption": "Figure 28. Comparison of mechanical pins and normal jet.", + "texts": [ + " As anticipated when the pins were located in Page 14 the cavity region of the projectile model, the results were similar to that of the normal jets which included an upstream increase in pressure and downstream regions of decreasing pressure immediately behind the pin followed further downstream by increases in pressure. These pressure changes were found to increase in both magnitude and in spatial separation with increasing tunnel Mach number as shown in Figure 27. Upon comparing the round and the rectangular pin and the normal jet at the Tap 14 location, Figure 28, it is apparent that the rectangular pin most strongly influences the pressure distribution on the projectile body. For the round pin this is due to the fact that the flow is disturbed more by the rectangular pin as the round pin provides a 3-D relieving effect. Both of the pins provide a stronger pressure change than the jet as they provide more flow blockage as the pin diameter is roughly three times that of the jet diameter. It should be noted that while the magnitude of the pressure changes are greatest for the rectangular pin, the sum total of the pressure force is greatest for the normal jet and moment on the projectile is greatest for the round pin" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001205_pime_proc_1972_186_090_02-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001205_pime_proc_1972_186_090_02-Figure2-1.png", + "caption": "Fig. 2. Hooke\u2019s coupling mechanism", + "texts": [ + " The case where = -24 leads to applications involving one cycle of volume variation for each rotor revolution; the case where 6 = -3$/2 leads to applications such as the four cycle engine in which two cycles of volume variation occur during each rotor revolution. Many alternative arrangements of gears involving different bevel shaft and layshaft rotation rates are possible. More compact arrangements can be made using sun and planet gears instead of a layshaft. The straight-through mainshaft and large bearing areas give a robust mechanism suitable for application in high pressure machines such as internal combustion engines. 2.2 Hooke\u2019s coupling Fig. 2 is the familiar Hooke\u2019s coupling. In this case the central member (rotor) is outside the coupling and the shafts are constrained by fixed bearings with intersecting but angled centre lines. It has the disadvantage compared with the precessing mechanism that the shafts and trunnion bearings are overhung and the bearings are Proc lnstn Mech Engrs 1972 Vol 186 62/72 at UNIV OF CINCINNATI on June 4, 2016pme.sagepub.comDownloaded from ROTARY PISTON MACHINE SUITABLE FOR COMPRESSORS, PUMPS AND I.C. ENGINES 745 Fig", + " 3, causes the rotor to have exactly the same motion as in the Hooke\u2019s coupling but the constraint of the second shaft is replaced by sliding contact between the flat circular face, F, and the cylindrical apex, C, of the rotor. This contact is practical in some cases because the inertial loads transmitted through the contact can be made small and fluid pressures in the two chambers exert no extra loads at that point. The avoidance of a second shaft and hub sphere clearly simplifies the construction enormously when compared with that of Fig. 2. Both the Hooke\u2019s coupling and the sliding apex mechanisms can be incorporated into variable displace- 3 SEAL SURFACE SHAPES It is shown in Appendix 1 that each of the above mechanisms constrains the rotor to move in a manner which satisfies the condition for making a positive displacement device with ports in a static casing. The shape of the casing is defined by fixing a seal line in the rotor and then moving the mechanism. As the rotor moves this line sweeps a surface which is continuous. In the case of the Wankel engine the casing surface is an epitrochoidal cylinder" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002763_icwapr.2007.4421627-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002763_icwapr.2007.4421627-Figure2-1.png", + "caption": "Figure 2 Fault experiment table of rolling bearing", + "texts": [ + "(13) directly adopts principle of FS(Fourier Series), the calculating amount is much more great. So the signal ( ){ }x n is transformed firstly in the frequency band (0 sf f\u2264 \u2264 )2 by FFT, then the demanded frequency band is zoomed in order to improve the frequency resolution by Eq.(13). This method can easily be carried out and can diagnose fault effectually. The fault of the rolling bearing is simulated on the special experiment table. And the construction of the experiment table is showed in figure 2. The experimental bearing is installed on the output axis of AC machine, and the AC machine drives the bearing directly. (a) Parameter of AC machine: Type: ADBE-56N4, Rated power: P=0.09kW, Rated speed: n=1350r/min. (b) Parameter of rolling bearing: Type: 6305, Pitch diameter: E=43.6mm, Rolling element diameter: d=11.8mm, Rolling element number: Z=7, Angle of contact: . 0=\u03b1 Installing 4371 type accelerated velocity sensors (made by Denmark B&K Company) on four positions, the No.4 and No.2 are installed on the bearing x direction and y direction respectively\uff0cand the No.3 and No.1 is on the x direction and y direction of bearing box respectively (as figure 2). And selecting the U60116C Data Acquisition Instrument, and setting the acquisition frequency as 5760Hz and sampling points as 1024, so analysis frequency is 2880Hz. Figure 3 shows the time-domain signal when the bearing outer ring fault happens. The spectrum of original signal is showed in figure 4. Due to the interferences, the low frequency is completely submerged, and the defect can\u2019t be diagnosed. Here \u2018db1\u2019 wavelet can be adopted to decompose the original signal and wavelet packet transform of three decomposition layers is adaptive" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000282_tmag.2003.810347-Figure8-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000282_tmag.2003.810347-Figure8-1.png", + "caption": "Fig. 8. Contours of magnetic flux density (mechanical angle 7.5 ).", + "texts": [ + " 7, the model used in the annealed B-H curve to the whole stator is Type 0, the model not used for an annealed B-H curve for a part in the root of the claw pole is Type 1, the model not used for an annealed B-H curve to the whole root of claw pole is Type 2, the model not used for the annealed B-H curve to the whole claw pole is Type 3, and the model not used for an annealed B-H curve to the whole stator is Type 4, respectively. Here, all models are used the same mesh and change only the B-H curve of stator. The contours of the magnetic flux density in the root of the claw pole of mechanical angle 7.5 are shown in Fig. 8. From Fig. 8, compared with Type 0, Types 1\u20134 have a thin color at the inside of root of the claw pole. That is, the magnetic flux density in the root of the claw pole becomes small. The cogging torque characteristic in Type 0 is shown in Fig. 9. This figure shows that one cycle is 15 . Although not illustrated, the cogging torque characteristics of Types 1\u20134 also have the same waveform. A comparison of the amplitude of the cogging torque is shown in Fig. 10. The difference of the maximum and the minimum of cogging torque is defined as the amplitude" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002778_tmag.2007.893864-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002778_tmag.2007.893864-Figure1-1.png", + "caption": "Fig. 1. Scheme of the system for detecting cracks.", + "texts": [ + " Furthermore, these parameters\u2019 changes when a crack is detected were compared. By laboratory experiments using the experimental machine for detecting cracks, the possibility of the detecting system was confirmed. The presence of cracks will result in different current density distribution within the line. At the same time, it will result in different magnetic density distribution around the line. This difference has been proposed to detecting cracks within the distribution line. For the purpose to measure the strength of the magnetic field, as in Fig. 1, it is proposed to use a ring-shaped detecting device with magnetic sensors placed alongside its rim at regular intervals. By running this ring along the distribution line, the measurement of magnetic field is possible. Digital Object Identifier 10.1109/TMAG.2007.893864 We can estimate the center point of distribution line by using the magnetic field. As shown in Fig. 2, a magnetic filed is distributed in a concentric pattern around the distribution line without a crack. Therefore, all lines that are perpendicular to the magnetic field direction should intersect exactly at one point, i" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002175_detc2007-34025-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002175_detc2007-34025-Figure1-1.png", + "caption": "Fig. 1 - Schematic of the proposed gear dynamic model.", + "texts": [ + " Recently, there has been some progress in modeling and analyzing hypoid and bevel geared rotor system dynamics [4, 5]. However, all of these hypoid gear dynamic simulation studies are based on pure vibration models that assume the system to oscillate about its mean position without considering any large displacement motion. This does not match physical reality because actual vibration of a geared rotor system may be strongly coupled to the large displacement rotational motion of the driveline system. As illustrated in Fig. 1, the classical hypoid gear vibration model assumes the pinion and gear angular displacements, py and gy are small perturbations and their average values are nearly zero, but in actuality both py and gy are large rotational displacements with superimposed high frequency perturbations, which are affected by the time-varying driving and load torques, and the dynamic characteristics of the entire driveline. Furthermore, gear mesh characteristics, such as mesh stiffness, kinematic transmission error, equivalent mesh point position and line-of-action vary with the actual torque load and gear rotational displacement", + " Due to the inclusion of the multibody dynamic and vibration coupling in the proposed model, certain types of simulation including operating transient response that cannot be obtained from previous models are now possible. A 14 degrees-of-freedom, nonlinear, time-varying lumped parameter model is proposed for use to represent a multicoordinate, coupled multi-body dynamic and vibration model of a hypoid geared rotor system. The formulation consists of a hypoid gear pair (i.e. pinion and gear), engine/driver and load elements as shown in Fig. 1. Pinion and gear are modeled as rigid conical body. Shaft and bearing are simulated by a set of stiffness and damping elements [6-8]. Two local coordinate systems Sl (Xl, Yl, Zl, lx, ly, lz) (l = p, g for pinion and gear respectively) are defined whose origins are at the centroids assuming uniform mass. It is noted that py and gy are large displacement coordinates, which were previously small displacement perturbations in the classical pure vibration case. The corresponding equation of motion can be expressed in matrix form as: }{}]{[}]{[}]{[ FqKqCqM =++ (1) Downloaded From: http://proceedings" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003410_2013.26820-Figure5-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003410_2013.26820-Figure5-1.png", + "caption": "Figure 5. Backlash characteristics.", + "texts": [ + " That is: \u23a5 \u23a6 \u23a4 \u23aa \u23a3 \u23a1 \u0394+\u03c9 \u22c5+\u0394+\u03c9\u22c5 =\u23a5 \u23a6 \u23a4 \u23aa \u23a3 \u23a1\u03c9 \u22c5+\u03c9\u22c5 N tt JNttJ N t JNtJ g ggg g ggg )( )()( )( )()( 12 2 2 111 12 2 2 111 (6) The coefficient of restitution between the two gears can be expressed as the ratio of the relative angular velocity between two gears after and before the impact, as shown in equation 7: N t t tt N tt e g g g g )( )( )( )( 12 11 11 12 \u03c9 \u2212\u03c9 \u0394+\u03c9\u2212 \u0394+\u03c9 = (7) The angular velocities of the two gears after the impact can then be obtained from equations 6 and 7 as follows: 1 22 12 1 22 11 12 1 22 12 1 2 11 1 22 11 1 )()()1( )( 1 )()1()(1 )( g g g g g g g g g g g g g g g g J J N te J J NteN tt J J N t J J Net J J Ne tt \u22c5+ \u03c9\u22c5 \u23a5 \u23a5 \u23a6 \u23a4 \u23aa \u23aa \u23a3 \u23a1 \u2212\u22c5+\u03c9\u22c5+\u22c5 =\u0394+\u03c9 \u22c5+ \u03c9\u22c5\u22c5\u22c5++\u03c9\u22c5 \u23a5 \u23a5 \u23a6 \u23a4 \u23aa \u23aa \u23a3 \u23a1 \u22c5\u22c5\u2212 =\u0394+\u03c9 (8) Applying the principle of angular impulse and momentum, the impulsive torque can now be expressed as: N T T t ttt JT gim gim gg ggim 1, 2, 1111 11, )()( \u2212= \u0394 \u03c9\u2212\u0394+\u03c9 \u22c5= (9) If there is no external load acting on the driven gear, the torque during the impact is given as: N T DT TDTT gim gimpactg gimggimpactg 1, 2,2 1,11,1 +\u2212= +\u2212= (10) Under the drive condition, the pitch line velocity is the same for the two gears, and the torque can be expressed in terms of their mass moment of inertia and angular accelera\u2010 tion. However, the two gears rotate independently in the free condition. Equations 11 and 12 give the torques acting on the two gears under the drive and free conditions: ][ ][ 211 2 2 1 2 ,2 211 2 2 1 1 ,1 ggg gg g driveg ggg gg g driveg DNDT JNJ JN T DNDT JNJ J T \u22c5\u2212\u2212 \u22c5+ \u22c5 = \u22c5\u2212\u2212 \u22c5+ = (11) 2,2 11,1 gfreeg ggfreeg DT DTT \u2212= \u2212= (12) The backlash was characterized by two parameters, back\u2010 lash (2 ) and mesh stiffness (kg), as shown in figure 5 (Wang et al., 2001). The mesh stiffness is defined as a non\u2010linear spring constant along the line of action of two mating gears, as shown in figure 6. Let xGP be the relative space between two mating gears of pitch circle radii of Rg1 and Rg2. The tooth deflection by the impact force between the two gears may be expressed as: \u239f \u23a9 \u239f \u23a8 \u23a7 \u03b5\u2212<\u0394\u03b5+\u0394 \u03b5\u2264\u0394\u2264\u03b5\u2212 \u0394<\u03b5\u03b5\u2212\u0394 =\u03b5\u0394 GPGP GP GPGP GP xifx xif xifx xg , ,0 , ),( (13) The mesh stiffness is then given by: ),( \u03b5\u0394 = GP GP g xg F k (14) The impact force was calculated from equation 2 by divid\u2010 ing the torque transmitted from the engine by the pitch circle radius of the driving gear" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001405_j.topol.2006.07.004-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001405_j.topol.2006.07.004-Figure1-1.png", + "caption": "Fig. 1. A robot with 6 arms.", + "texts": [ + " The homeomorphism types of it when it is singular is also given. \u00a9 2006 Elsevier B.V. All rights reserved. MSC: primary 57M50; secondary 58E05, 57M20 Keywords: Configuration space; Planar linkage; Morse function 1. Introduction We study the configuration space of the linkage of a robot which can move only in a plane. We consider a robot which has n arms such that each arm is of length 1 + 1 and has a rotational joint in the middle, and that the endpoints of the arms are fixed to n equally located points in a circle of radius R (Fig. 1). We assume that its arms and joints can intersect each other. Let us call this robot a \u201cspider\u201d and denote the configuration space of the spiders with n arms of radius R by Mn(R). Let x be a point in Mn(R) that corresponds to a spider such that none of the arms is stretched-out nor folded. All the angles at the joints of the arms belong to (0,\u03c0). The configuration of a spider, if it is close to the above mentioned, is determined by the position of the body. Therefore, the neighbourhood of a generic point in Mn(R) is of dimension 2" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001961_s1387-2656(06)12004-9-Figure11-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001961_s1387-2656(06)12004-9-Figure11-1.png", + "caption": "Fig. 11. Schematic representation of the commercial Biacore system and the change in incidence angle during a binding process. (a) When there is no analyte bound to the sensor surface the reflective intensity plot shows the default incidence angle indicated by a sharp dip. (b) When the analyte binds to the sensor surface the change in the absorbed mass on the surface causes the dip representing the incident angle to", + "texts": [ + " SPR spectroscopy relies on the SPR phenomenon, which allows the realtime measurement of biomolecules binding to biomimetic surfaces without the application of a specific label because the SPR method is dependent on the change in adsorbed mass at the sensor surface [146]. SPR is a surfacesensitive technique where the ligand is immobilised onto a solid support and the solute is in solution and the binding event can be readily detected and analysed. A typical SPR system consists of an SPR detector, light source, flow channel and sensor surface, comprising a conducting surface such as gold or silver. P-polarised light is emitted by the light source and reflected on the gold-coated sensor surface and detected by the diode array detector (Fig. 11). The SPR phenomenon causes a change in the intensity of reflected light at a specific angle and the SPR detector detects these changes in optical properties at the sensor surface following adsorption and desorption of a solute bound to the sensor surface. The change in optical properties depends on a number of factors including the thickness of the gold surface, the wavelength of the light and most importantly the adsorbed mass on the sensor surface. The SPR technique can also be fully automated using the commercially available instruments and large numbers of samples can be rapidly and conveniently analysed", + " Typically, ligands are immobilised on the surface of a sensor chip, which is covered by a thin gold layer. When the analyte is injected over the surface in a continuous flow, it adsorbs onto the immobilised ligand and so changes the incidence angle by modifying the refractive index at the surface of the sensor chip. The resulting sensorgram is a plot of the change in SPR incidence angle against time, which allows the binding event between the analyte and the ligand to be visualised and can be used to gain information on the binding kinetics of the interaction (Fig. 11). There are two commercially available chips that are suitable for studying membrane-based systems. The hydrophobic association (HPA) sensor chip consists of self-assembled alkanethiol molecules covalently attached to the gold surface of the chip which can be used to prepare hybrid BLMs by the fusion of liposomes onto the hydrophobic surface [158]. The availability of the HPA sensor chip significantly improved the preparation of solid supported lipid membranes to investigate membrane-mediated interactions", + " The L1 sensor chip is composed of a thin dextran matrix modified by lipophilic compounds on a gold surface, where the lipid bilayer system can be prepared through the shift also shown in (b) as angle 2. The resulting sensorgram monitors the change in angle as shown in (c). (c) A schematic representation of the real-time sensorgram of a binding event. During the association phase, the analyte is present in the buffer flow and binds to the sensor surface. This is followed by the dissociation phase after removal of the analyte from the buffer flow. Analysis of these sections of the curves provides response units (RU). capture of liposomes by the lipophilic compounds as shown in Fig. 11 [150]. The immobilisation of the biomimetic lipid surface onto the sensor chips is generally a fast and reproducible process. Both the HPA and L1 sensor chips can be conveniently applied to the study of membrane-based biomolecular interactions and to measure the binding affinity related to these interactions and an increasing number of examples of the use of this membrane surface in a wide range of biological applications has been reported including analysis of cytolytic peptide action, membrane-mediated cell signalling and neurodegeneration [161\u2013163]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001384_j.ijmachtools.2006.02.011-Figure9-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001384_j.ijmachtools.2006.02.011-Figure9-1.png", + "caption": "Fig. 9. Process flow and CAD model of a hemispherical shape: (a) process flow to ablate a physical hemispherical shape; (b) CAD model and dimensions of a hemispherical shape.", + "texts": [ + " After the origin of workpiece and tool should be coincided, a rough ablation and a fine ablation process are performed to create the part. A fundamental feature, a hemisphere, is selected to show the applicability of the proposed process. The size of the considered part is 90mm 90mm 45mm. Before the toolpath is generated, the offsetting value and layer interval are calculated using Eq. (9) derived as experimental results. The layer interval is 6.1mm at 110V (input voltage), 40mm/s (ablating speed). The whole process of the RHA process from CAD model to three-dimensional prototypes is illustrated in Fig. 9(a) in the case of a hemispherical shape. Ablating the hemispherical part takes approximately 4min, and the remaining material is almost entirely removed by melting and thermal decomposition. The characteristics of the proposed process, such as rapid shaping and no remaining material, are clearly demonstrated. The measured dimensions of the part are 89.89mm 89.92mm 44.94mm, as shown in Fig. 9(b). The dimensional accuracy of the part is within 0.12%. Considering that a typical commercial rapid manufacturing process has a dimensional accuracy within 1% [11], the proposed RHA process has good dimensional accuracy. The standard test part used to verify the machining process is taken through the RHA process. Usually, the standard test part including fundamental convex and concave shapes, free surface shapes, among others, has been applied to the rapid manufacturing system in order to ARTICLE IN PRESS H" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002628_cdc.2007.4434082-Figure6-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002628_cdc.2007.4434082-Figure6-1.png", + "caption": "Fig. 6. A FPS game model. Note that s1, s2 are always expressed in a camera fixed coordinate system.", + "texts": [ + " Similarly to the UGV case above, the game is controlled using a two joystick pad. The difference is that here, the camera joystick controlles the camera rotation relative to a world fixed coordinate system, while the other joystick controlles the translation of the character, relative to a camera fixed coordinate system. That is, pushing the joystick forwards corresponds to the character moving in the direction of the camera, and pushing the joystick leftwards means moving perpendicular to the view direction. This is illustrated in Figure 6, s1 corresponding to \u201cjoystick forward\u201d and s2 to \u201cjoystick left\u201d. Writing down the equations of motion we get ( x\u03071 x\u03072 ) = R(\u03c8) ( s1 s2 ) (3) \u03c8\u0307 = s\u03c9 or, using assuming the operator controls accelerations( x\u03081 x\u03082 ) = R(\u03c8) ( a1 a2 ) (4) \u03c8\u0307 = s\u03c9 Above we have used the notation R(\u03b8) to denote a rotation matrix, i.e. R(\u03b8) = ( cos \u03b8 \u2212 sin \u03b8 sin \u03b8 cos \u03b8 ) . Note also that R(\u03b8)R(\u03c8) = R(\u03c8)R(\u03b8) = R(\u03b8 + \u03c8). Now consider the search mission described in Example 1 above, carried out with a FPS game interface" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001550_iros.2005.1545550-Figure6-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001550_iros.2005.1545550-Figure6-1.png", + "caption": "Fig. 6. New prototype. Weight is 1.7 [kg].", + "texts": [ + " This motor has the special controller unit that simplifies the building up of closed feedback loop. The motor for pitch motion was located at some distance from the incision hole, and linkage mechanism was added for transmission (circled number 3 in Fig.3). We intended to keep sterilization around incision hole by separating sterilized and nonsterilized part via linkage mechanism. Linkage mechanism also works as a mechanical stopper to limit the working range of pitch for safety. The new prototype is shown in Fig. 6. Weight is 1.7 [kg]. FWM was 62\u00d752\u00d7150[mm3], 0.6[kg], and gimbals mechanism was 135\u00d7165\u00d7300[mm3], 1.1[kg]. III. EVALUATION EXPERIMENTS We conducted mechanical performance evaluation of our new prototype. Torque of pitch, roll, and rotation, and force of translation were measured by 6-DOF strain gauge force/torque sensor (MINI sensor 8/40, BL Auto Tech, Japan). We repeated the measurements for twenty times, and calculated the average and standard deviation. We also measured the working range of each axis, and maximum driving speed without load" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003040_iccasm.2010.5622573-Figure5-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003040_iccasm.2010.5622573-Figure5-1.png", + "caption": "Figure 5. Bearing force and oil film pressure diagram", + "texts": [], + "surrounding_texts": [ + "The elastic deformation under load of the joumal and bearing mainly were obtained through the respective kinetic equation, shown as equation (2), (3). As the kinetic equation contains items of film pressure, and the Reynolds equation of lubrication analysis constitutes another cycle. Under the pressure of oil film, the main bearing kinetic equation is shown as follows: (2) M: bearing mass matrix; X B : bearing displacement matrix; K: bearing stiffuess matrix; D: bearing damping matrix, the linear combination of stiffuess matrix and mass matrix;f: oil film force from the bearing shell. The kinetic equation of the crankshaft main journal, under the oil film pressure and external load, can be shown as follows: (3) m: crankshaft main journal mass matrix; X J : crankshaft journal displacement matrix; fJ : oil film force from the crankshaft journal; fA : external load of the crankshaft main journal. The relation among bearing oil film pressure, oil film thickness, surface roughness, and elastic deformation, is coupling, which is calculated synchronously. Those mathematical models establish the numerical analysis model of flexible multi-body dynamics crankshaft system for iterative solution to obtain film pressure, orbital path and the minimum oil film thickness (MOFT). Connector's force and torque is considered as constraint torque of elastic body, while the elastic body force and deformation as the boundary conditions of connect joint EHD analysis, iterative solution, in order to solve the connective relationship between connect joint and elastic body. IV. RESULTS ANALYSIS This model is a two-stroke six-cylinder marine diesel engine, which is supported by eight main bearings. Under full load, rated speed, bearing clearance for the O.3mm case, taking into account the deformation of elastic bodies, and the system coupling effects among the various components, we established lubrication performances of multi-body dynamic model to work out bearing force, oil film thickness and orbital path in a working cycle. The main bearing force is very important in the structural design of the bearing, plays a leading role on the lubrication condition, and the focus must be considered. As we know from Figure6, the vertical forces of main bearings in a working cycle are larger when neighboring cylinder fires. VI5-181 Because of different crank angle interval, forces are not the same. The force of No.2 main bearing was the largest, and reached to 1518kN at the 249\u00b0CA (crank angle). The forces of No.2, No.3, No.5 and No.6 appeared three peaks in one cycle. For example, the No.3 main bearings appeared peak forces when the adjacent No.2 cylinder fire at about 250\u00b0CA and No.3 cylinder fire at about 130\u00b0CA, and the peak force occurs also at about 10\u00b0CA. At about 10\u00b0CA, the reciprocating inertia force of No.2 and No.3 at the same time will be large downward (the graph of crank-end was shown in Figure7), so the force of No.3 bearing appeared peak force at that time too. The force of No.4 main bearing is small for the gas force and the inertia force to offset. The force of No.8 main bearing is the smallest among them, because it is at the final end of crankshaft, and the work forces of all cylinders have a little impact on it, its working condition is smoother. Figure8 shows the orbital path of main bearings. Analyzing the orbital path, we can know that overload areas and light load areas, direct observation of a bearing and the journal will appear the cavitations and other damages, and judge the bearing lubrication conditions. We can see from the Figure8 that none of the bearings orbital paths has break phenomenon, bearing a smooth work, lubrication performance is well. Main bearings of No.1 and No.7 are on upper shell part of the district-free orbital lines, the minimum oil pressure in this area, and oil film thickness is the largest and more suitable for opening groove. The eccentricity of No.8 main bearing is always small because of their relatively light load. Orbital paths of No.2 to No.6 main bearings on the middle of the lower shell appears alternating back and forth in varying degrees, and the shell is easy to wear in those regions. The rapid centrifugal and centripetal movement of orbital path of No.3 and No.6 main bearings will prone to capitation to cause damage of the bearing shell. The mInImum oil film thickness of the main bearing under each crank angle was shown in Figure9. Table 2 shows the minimum oil film thickness and its time of occurrence (in crank angle indicated). We can know that minimum oil film thickness of No.1 and No.2 main bearings are smaller respectively, 4.89f..lm and 4.73f..lm. Minimum oil film thicknesses of other bearings are larger than 5f..lm. The load of No.8 main bearing is smaller; oil film thickness is relatively greater. Compared with the other main bearing, lubrication condition of No.2 main bearing is slightly worse. V15-182 FigurelO is the average oil film pressure distribution of No.2 main bearing. From the FigurelO we can know visually that oil film pressure distribution is concentrated on the middle of the lower shell. The bearing alloy under larger oil film pressure is easier to fall off, where bearing damage may occur. mo-.-- ----------, ___________ ---, 90 ---0--- O.3rrm -----0-- O.2rrm 0.1 60 120 180 240 300 360 CrankAngle(deg) Figure I I. MOFT under different clearance Bearing lubrication is closely related to bearing clearance. Figurell shows minimum oil film thickness of No.2 main bearing changes with the diffident bearing clearances. The bearing clearance reduces from O.3mm to O.lmm, and the minimum film thickness increases from 4.73\ufffdm to 5.31\ufffd, but the corresponding friction power losses increases and the peak friction power increases from 14.59kW to 17.99kW. Reduction of the bearing clearance is conducive for increasing the minimum oil film thickness. If the clearance is too small, oil flow lacks, friction, wear increases, and temperature oil viscosity decreases go against well film. V. CONCLUSION a) The tribology properties of the bearing and the crankshaft-bearing system dynamic behavior were unified into a computing model, and the lubrication was included in the calculation of the elastic deformation of the bearing for carrying out coupled simulation. b) Because the neighboring crank angle and the adjacent cylinder firing interval angle varied, the main bearing force varied, peak vertical forces of No.2, No.3, No.5 and No.6 reached 1400kN around, which are larger than others. Because of the impact of large inertial force, the main bearings force appeared three peaks in an one working cycle. c) The force of No.2 main bearing is the largest among them, reached 1518kN at the 249\u00b0CA. Minimum oil film thickness is 4.73\ufffdm at the 266\u00b0CA. The lubrication of No.2 main bearing is slightly worse. d) The clearance of No.2 main bearing changes from O.3mm to O.lmm, and the minimum oil film thickness increases from 4.72\ufffdm to 5.31\ufffdm, so does the friction power accordingly. For ensuring adequate oil flowing to bearings in the most suitable working temperature, reducing bearing clearance appropriately will help to improve bearing lubrication. ACKNOWLEDGMENT I would like to extend my sincere gratitude to my supervisor, Shulin Duan, for his instructive advice and useful suggestions on my thesis. Last my thanks would be for those who helped me during the writing of this thesis." + ] + }, + { + "image_filename": "designv11_69_0003808_jrc2009-63011-Figure8-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003808_jrc2009-63011-Figure8-1.png", + "caption": "Figure 8. Service-induced residual hoop stress for braking duration of 60 min (36\u201d freight car wheel - rim thickness = 1.5 in).", + "texts": [ + " The temperature developed on the tread surface during the on-tread braking for 20 min is not high enough to reduce the yield strength of the material at the tread surface; therefore, the material has not yielded in compression nor developed any plastic deformation, thereby developing no additional residual stresses in the wheel rim. Copyright \u00a9 2009 by ASME ms of Use: http://www.asme.org/about-asme/terms-of-use Down Figure 7 shows the residual stress distribution in a 36\u201d freight car wheel with rim thickness 38.1 mm (1.5 in), developed during both the manufacturing process and thermal brake loading for 40 min. The results show that the residual hoop stress on the taping line is -100 MPa (-14.5 ksi). Figure 8 shows the residual stress distribution in a 36\u201d freight car wheel with rim thickness of 38.1 mm (1.5 in), developed during both the manufacturing process and thermal brake loading for 60 min. The results show that the thermal brake loading for 60 min reverses the residual hoop stresses on the taping line from compression to tension in the wheels of considered rim thicknesses. The maximum tensile residual hoop stress developed in the wheel rim is 320 MPa (46.41 ksi). 7 loaded From: http://proceedings" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000803_1.1809635-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000803_1.1809635-Figure2-1.png", + "caption": "FIG. 2. tw parameter position related to the weld joint.", + "texts": [ + " A device was used to ensure minimal distortion both longitudinally and transversally and adequate assist gas flow. Since hard metals are quite reactive, argon was used as the protective gas. During welding the parameters were changed around an optimized value achieved in preliminary trials in order to analyze the influence of the heat input, interaction area and horizontal focal point position (tw). This was seen to play a major role on the weld quality.7 \u201ctw\u201d is considered zero when the laser spot is positioned exactly into the steel/hard metal interface and is positive to the steel side (Fig. 2). Tests were conducted with tw ranging from 0 to 0.3 mm. Welded joints were visually inspected and samples were cut by electroerosion, mechanically polished and etched. Murakami\u2019s reagent was used to evidence the hard metal and the welded zone and Nital (5% in volume) for revealing the steel structure. Optical and electron scanning microscopes were used to study the microstructures in the fusion and heat affected zones of the welds. Auger electron spectroscopy was used to identify existing elements in different structural features in the weld" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000460_ias.2001.955529-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000460_ias.2001.955529-Figure1-1.png", + "caption": "Fig. 1: Steady state equivalent circuit of the induction motor for negative sequence components", + "texts": [ + " THEORETICAL DEVELOPMENT When both positive and negative sequence voltages spV and snV are applied to a symmetrical induction motor, the two sequence currents do not react upon each other. Thus they can be treated independently, using separate equivalent circuits for positive and negative sequence currents. The sequence impedances of a symmetrical motor under unbalanced supply conditions are simply the ratios of the sequence voltage to the sequence current, i.e sp sp p I V Z = , sn sn n I V Z = , and 0 0 0 s s I V Z = . (1) Fig. 1 shows the corresponding equivalent circuit for negative sequence components. From the negative sequence equivalent circuit (Fig. 1), the negative sequence impedance nZ can be expressed approximately by 0-7803-7116-X/01/$10.00 (C) 2001 IEEE mr mr sn ZZ ZZ ZZ + += (2) where sss jXR +=Z , r r r Xj s R \u2032+ \u2212 \u2032 = 2 Z and mm jX=Z s is motor slip, and Rs , rR\u2032 , Xs , and rX \u2032 are stator and rotor resistances and leakage inductances respectively. From Eqn 2 we can deduce that the negative sequence impedance nZ may be expressed approximately by rsn ZZZ += nnrs r sn jXRXXj s R R +=\u2032++ \u2212 \u2032 += )( 2 Z (3) where the magnetising reactance mjX has been omitted, since mrr jXXjsR ))2/(( <<\u2032+\u2212\u2032 " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003428_icicisys.2010.5658813-Figure3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003428_icicisys.2010.5658813-Figure3-1.png", + "caption": "Figure 3. Force analysis of pendulum 1", + "texts": [ + " More specifically, the mathematical model in state space is derived as follows: 0 0 1 0 0 0 0 0 XI 0 0 0 1 0 0 0 0 XI 8 1 0 0 0 0 0 0 0 0 81 Xl AfJ A(mlg/I - fJh) -AfJ AfJh XI 0 0 0 0 81 JI JI JI JI 8 1 x, 0 0 0 0 0 0 1 0 x, 8, 0 0 0 0 0 0 0 1 8, x2 0 0 0 0 0 0 0 0 x, 8, -AfJ AfJh 0 0 AfJ A(m,g/, - fJh) 0 0 8, J, J, J, J, 0 0 0 0 0 A/lml 0 + JI [:: ] 0 0 0 0 0 0 Am212 J, Substitute the value of the parameters mentioned in Tab.1 into matrices A and B, the matrices as follows can be obtained: 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 19.782 8.8879 0 0 -19.782 7.9128 0 0 A= 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 -19.783 7.9128 0 0 19.782 8.8879 0 0 0 0 0 0 0 l.7143 0 B= 0 0 0 0 0 0 l.7143 c. Comparison with Newtonian Mechanics Modeling Choose the two pendulums in Fig.1 as research objects, the force analysis of pendulum 1 is shown in Fig.3. The resultant force acting upon the two pendulums in the vertical direction is: (11) (12) The resultant force acting upon the two pendulums in the horizontal direction can be obtained as: d2 \u2022 N1 + Sh = m1-2 (X1 -/1 smB1 ) (13) dt d2 \u2022 N2 - Sh =m2-2 (x2 -/2 smB2 ) (14) dt The tension generated by the spring along the slide rail is: \ufffdw Lo L S =k ( L-L)-=k\ufffdw ( I--)=k\ufffdw ( l- ) h s O L s L s .J 2 2 d +\ufffd(l) matter which is based on Lagrange Equation or Newtonian Mechanics. Form the above analysis, it can be seen that (15) Newtonian Mechanics modeling based on vector In terms of the rotation law of a rigid body around a fixed axis, the following equations are given: JI\ufffd =MR +MN 1 1 d2 (II cos\ufffd) " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001979_50037-5-Figure7.129-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001979_50037-5-Figure7.129-1.png", + "caption": "Figure 7.129 Constraints and loads applied to the model.", + "texts": [ + " The final action is to apply external loads. In the case considered here a pressure acting on the top surface of the cylinder will be used. From ANSYS Main Menu select Solution \u2192 Define Loads \u2192 Apply \u2192 Structural \u2192 Pressure \u2192 On Areas. The frame shown in Figure 7.127 appears. Top surface of the cylinder should be selected and [A] OK button pressed to pull down another frame, as shown in Figure 7.128. It is seen from Figure 7.128 that the [A] constant pressure of 0.5 MPa was applied to the selected surface. Figure 7.129 shows the model ready for solution with constraints and applied load. Now the modeling stage is completed and the solution can be attempted. From ANSYS Main Menu select Solution \u2192 Solve \u2192 Current LS. A frame showing the review of information pertaining to the planned solution action appears. After checking that everything is correct, select File \u2192 Close to close that frame. Pressing OK button starts the solution. When the solution is completed, press Close button. In order to return to the previous image of the model, select Utility Menu \u2192 Plot \u2192 Replot" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000204_naecon.2000.894996-FigureI-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000204_naecon.2000.894996-FigureI-1.png", + "caption": "Figure I - Landing profile", + "texts": [], + "surrounding_texts": [ + "The automatic landing system being described in this paper was designed with the Bombardier CL-327 in mind. While the craft does look significantly like its predecessor the CL-227, the vehicle is in fact significantly redesigned. One of the more significant changes was the adoption of GPSDGPS technology as one of the primary navigation sources. Though the CL-227 did have an autoland system, the idea for this project was to generate fresh ideas to supplement the existing arrangement. The project itself was a collaborative effort between Concordia University and Bombardier Inc., where the two worked closely during development. Bombardier first laid out the fundamental criteria for the autoland around which a process was devised. The design which followed was to be decidedly novel; a fresh slate approach. This was combined with an overall philosophy of system simplicity which manifest itself in the general avoidance of complicated, time consuming solutions. Beyond this, the each individual component was designed to be easily recodigurable as well as non-specific to this craft. This flexibility allows the ideas to be implemented in other situations or in other research altogether. However, these aforementioned generalised conditions were no excuse for compromising the performance objectives of the system which were lofty by the very nature of the vehicle. After a thorough evaluation, it was evident to us that the autoland process would function through the use of three key components. The first, an algorithm to overcome the navigation source switchover from GPS to DGPS and the inevitable position error between the two sources. The second, a trajectory generator capable of, amongst other things, full adjustability, easy reconfiguration, while being continuously smooth. The third and final component is a system controller to work in conjunction with the trajectory generator to provide complete motion control. However, prior to explaining what these components do and how they function, the terms of the problem are laid out. 2. PROBLEM DESCRIPTION 2.1 Landing Profile The generalised landing profile for the craft consists of 3 points in space: the recovery point (RP), the landing set point (LSP), and the landing point (LP) (figure 1). The line connecting the first two points is the glideslope (GS) and that connecting the latter two is the final descent portion (FD). The FD is to always be a vertical section of a short and constant distance. The GS is characterised by an infinitely variable geometry. Firstly, the angle of descent relative to the horizontal can range from horizontal (OO) to vertical (90\"). The geometry also encompasses the heading or bearing angle and can be anything from 0\" to 360\". Finally the length of the GS requires full adjustability (0-1000 meters). The final component of the landing profile is the safety volume (SV). This is a cylindrically shaped region about the RP which, after switchover from GPS to DGPS, the craft is all but assured to be within this region. Hence the size and shape of the SV is directly dependant on the accuracy of the GPS system. 2.2 System Requirements The autoland system is to begin after switchover from a navigation system which heavily weights GPS information, to one which holds DGPS information in the same regard. This switchover comes after the craft has attempted to reach the RP through GPS based navigational means. Therefore, the fundamental tasks are to intercept the GS prior to exiting the SV and then to follow the landing profile to earth. . It must do the above with strict control accuracy in inherently gusty environments and in the presence of high navigation system signal noise. Furthermore, the solution must abide by and respect all limitations on the craft which include rate-of-tilt, vertical and total velocity caps, and an inherently quicker craft response in the vertical as compared to the horizontal. All this must be done with GPSDGPS as the primary navigation source. 3. INTERCEPT ALGORITHM This algorithm is required due to the navigation system switchover to one of higher accuracy. During both GPS mode and its DGPS replacement, the target is always the RP. The vehicle believes itself to be in the correct spot before switchover, however the more accurate DGPS system reveals some position error. The magnitude of this error varies, but the craft will always be found within the SV. This occurrence follows the definition of the SV. One solution would be to direct the craft to its original target, the RP, and be done with it. This would indeed be the simplest response, but it comes with some flaws. If the craft lies below the RP after switchover, this solution would increase the craft's altitude only to send it down the GS seconds later while carrying out the landing profile. This would be a poor use of time and energy. The intercept algorithm picks a more intelligent spot for the craft to intercept the autoland flight profile , while at the same time assures that it does so before the vehicle exits the SV. The algorithm consists of two distinct parts with the first providing logic for intercept determination while the second devises a simple approach to locating the craft about the RP. 3.1 Target Point Selection The space about the RP must be split into appropriate zones. Considerable deliberations brought about the divisions represented in figure 2a. This particular arrangement depicts the ideal system. From the 6 original zones, only three are unique as the plane containing the landing profile (central plane; CP) is a plane of symmetry. This is apparent when viewed perpendicular to the landing profile (figure 2b). With the three core zones in mind, four courses of action can be taken to chose a GS intercept. Assuming the craft is in zone 1, the appropriate solution is for the vehicle to fly horizontal to the GS. If found within zone 2, the UAV is located above the RP and effectively behind the GS. In this sole instance, the RP is the best target for the craft as it would not encourage the previously mentioned duplication of movement or any other kind for that matter. If the craft is found within zone 3, two solutions are possible depending on the h r\" h LSP Figure 2 - Ideal division of space about RP: (a) 3 - 0 view: ( b ) perpendicular to landing profile (2 -0 ) circumstances. The initial solution uses what is called the cylinder radius approach. It calculates the horizontal distance between where the craft is found and the RP. It then finds the point on the GS which corresponds to this horizontal distance and chooses it as the intercept. When viewed in three dimensions, the horizontal distance is effectively rotated about the RP forming an arc. Seeing as this method does not take vertical position of the craft or the intercept into account, the arc effectively becomes a portion of a cylinder when fully depicted. The one problem with this method, as opposed to the solutions for zones 1 and 2, is that it doesn't assure that the intercept will fall within the SV. If after engaging this procedure the craft falls outside the SV, method four is employed. It selects as the intercept, the intersection of the GS and the SV. This is the last point (lowest altitude) on the GS which can be selected. Combined, the above methods will provide solutions for every possible location about the RP that the craft may find itself. In addition to this, the algorithm may be tailored to fit the accuracy requirements for an assortment of situations, vehicles, or applications. Figure 3 - Replication of ideal system: (a) 3 - 0 view; (b) perpendicular to landing profile (2 -0 ) 3.2 Craft Locating About Rp The object of this portion of the algorithm is to replicate the ideal system depicted in figure 2 simply, yet accurately. As opposed to depicting the planes in space via equations describing them exactly where they lie and then calculate which zone the craft lies within, a simpler and considerably less calculation intensive approach is taken. Imagine 3 planes intersecting one another at right angles to one another similar to the orientation at the origin of the Cartesian coordinate system. With the system's common intersecting point located at the RP, the vertical plane aligned with the landing profile, and another plane aligned horizontally, an effective reproduction of the ideal system is obtained (figure 3) . Though not identical, it is a very reasonable reproduction of the ideal system. Determining which quadrant about the RP the craft lies within is determined via transformations [l]. Imagine the system of planes with the UAV position included, superimposed with the origin of the navigation coordinate system via translations. This would first involve a transformations equivalent to the opposite of the RP\u2019s position. Then, one subsequent rotation of this translated system would be required to align it with the navigation coordinate system. With this, locating the UAV position about the RP would simply involve checking the signs of this transformed position. If it is found that the craft lies within the one quadrant which doesn\u2019t correspond to the ideal system, then a couple of elementary supplemental calculations can be used to quickly resolve the matter. 4. TRAJECTORY GENERATOR This and the following section describes the two components which together make up the craft\u2019s motion control system. This section involves the trajectory generator which provides the ideal locus of points according to specific temporal law. The mandate of the trajectory generator is to provide an ideal path between any two points in space regardless of bearing angle, descent angle, or path length. It must also implement and adhere to the maximum values for total and vertical velocity, as well as respecting a limit on the rate-of-tilt of the craft. The general motion profile involves an acceleration followed by a constant velocity and finally a deceleration to rest or segments A, B, and C respectively. The segments A and C are deemed primary while B is given secondary status as it involves no acceleration. Segment B is by intention the adjustable link to accommodate for various path lengths. Craft stability can be directly related to smooth motion control so therefore acceleration profiles are the primary evaluation tools to assess the various motion profiles. All tolled, four profiles were considered, these being the step, ramp, polynomial, and the cosine profile (figure 4) [21 131 141 [5]. The first three are all burdened with a common weakness. Each contained discontinuities and which in tum reduced the craft performance envelope in so far as it could withstand less rigorous disturbances. The critical points which are required to be free of discontinuities, occur at the end of segment A and the beginning of segment C, both high speed transitions. The solution to this problem is found with the cosine profile. This ultra-smooth solution provides for an extended performance envelope in regards to the craft\u2019s ability to withstand higher winds and greater noise relative to other profiles. The manner in which the base profile is obtained involves first inverting the cosine function and then adding a value equal to the function\u2019s amplitude (figure 5 ) . After manipulation of equations, the acceleration and deceleration profiles can be described by 3 fundamental variables, the peak velocity ( Vp), the duration of acceleration (tA), and the duration of deceleration (tc). How these variables are determined is dependant upon which of three specific cases exists. The cases in question are entirely dependant on geometry between the two points in question which the craft is navigating between. The first case is when the path is considered \u201clong\u201d. This is the initial assumption until proven otherwise. In this case, the fundamental 3 variables are set to their respective maximums. These values are found either by system tuning (i.e. I \\ . r c . ) or ;ire declared according t o safety cot~ccriis (i.e. Vi,). Recull that tlie litiiiting 01' the total or pcrth velocity (V,) is a condition i n the basic premise of the tra,jectory generator. The teriii --long path\" relates directly to the distance t~iveled during acceleration and deceleration seginents. Seeing as in this case fc, and V p are declared to be their respecti\\,e tiiaxiniitiiis. the distance traveled in segments A and C is :iIso a maxiniitm and :i coiist;iiit value and is teinied the critical distance ([I, ,,). If the distance between the two points in qiirstion is longer th:in the critical distance. then h e path is considered long a n c l the constant \\,elocity s e p e n t (5) will cover the additional spnn. The second G I S ~ exists uhen the distance het\\\\'een the two points is less than the critical distance. Thij means that when the luniiniiienial 3 ;ire set ;iccordinF to casc 1. though segmcnt H i \\ effectiwl> negntcti. tlie distance CO\\~L'IIXJ during accelcratiou and clecelei-ation is still too Ion?. The rcniedy for this I s to rednce the fiintlmiental .; \\.nriables fi-om tlieiiiiiaiitiiiiins and rei~io\\c the constant \\ clocit! s c p m \" The reduction of' I , . /(,. arid VI, i \\ cloiie in ;I proportional iiianner so ;IS to prcscrvc the jerk i,rolile :rnd maintain hi$ c\\.stcni eft'iciciic\\. This proportional reduction 01' the :icccIcrntion prolile is accomplished wing a sc:iling I'i~ctor 1 1 . I f /r=0.5. then the \\dues for 1, and the pcak accelcratiori are cach half their inasimums (figure 6). The dccelel.atioii profile receives similar treatnieiit. An ccpation exists which relates the distancc between the 2 points i n space. 1 1 . and the fiind;imental 3 variables. This results in a niociification o f t i i t ' fiindmiental .? wch that after completion o f segments il and c'. the cixt'c will arrive at the second point. Case three occurs whcn the veitical \\:eIocity exceeds its pi-escribed ma.ximiini. This can arise :is :I result ol' V p being tlie only volocity being limited thus far, because of path geometry. or through a combination of the two. Thc remedy is to set the \\wtical velocity cqiial t o its ni;ixiiniini and then 1 ' 1 0 8 5 0 6 06 a8 40 10 C 3 so5 recalculate V, according to what the path geometry will allow. To find t A and tc, scaling factor evaluation is revisited, thereby assuring parameters are as close to optimal as possible. 5. CONTROLLER The unit which actually drives the vehicle is the controller. It takes the results obtained from the trajectory generator and puts them to work. It was a conscious effort, due in no small part to the autoland project\u2019s scope, to keep the controller as uncomplicated as possible. The end result is a simple PID controller supplied with position and velocity errors. Though the control solution was decidedly common, the results achieved are very encouraging. Figure X depicts the desired craft velocity, the desired craft acceleration, the winds faced, and the resulting position error of the craft when the motion control system is mated with Bombardier\u2019s flight dynamics model for the CL-327. Included in the simulation, but not shown, are the navigation signal noise for both the position and velocity signals. The magnitudes for these signals are M.15 m and a . 5 m / s respectively but both are sampled at 25 Hz. The results are shown for the horizontal plane only though the vertical gives comparable results. It is important to note that these disturbances are considered above normal. 6. CONCLUSION The system designed to autoland the craft fulfilled all that was asked of it, but moreover, it does it in a very uncomplicated fashion. Through the use of the intercept algorithm, switchover from GPS to DGPS is a simple maneuver. Motion control is accomplished by the collaborative effort between the trajectory generator and the controller. Excellent accuracies can be realised even in the presence of above normal disturbances. Ongoing flight testing looks to confirm these findings. 7. ACKNOWLEDGEMENTS The authors would like to thanks Bombardier Inc. For their continuous support throughout the project. 8. REFERENCES [ 11 P. A. Egerton and W. S. Hall, Comuuter Graphics: Mathematical First Steus, Prentice Hall, pp. 174-6, 1998. [2] R. L. Anderson,, A Robot Ping-Pong Plaver: Experiment in Real-Time Intelligent Control, MIT Press, pp. 87-94, 1988. [3] M. S. Mujtaba, \u201cDiscussion of Trajectory Calculation Methods,\u201d in Exploratory study of comuuter intemated assembly systems, Binford T.O. et al., Stanford University Artificial Intelligence Laboratory, AIM 285.4, 1977. [4] C. E. Wilson and J. P. Sadler, Kinematics and Dvnamics of Machinery, Harper Collins, pp. 341-4, 1993. [5] M. Brady et al., Robot Motion: Planning and Control, MIT Press, pp. 221-30, 1983. 9. BIOGRAPHIES Michael Bole is a research associate at Concordia University in Montreal and works with Dr. Jaroslav Svoboda. He recently received his Masters degree in Mechanical Engineering from Concordia. His thesis was part of a collaborative enterprise between Concordia and Bombardier Defense division in a bid to design components for a new automatic landing system for Bombardier\u2019s vertical take off and landing, CL-327. He has recently presented work at the Canadian Aeronautics and Space Institute\u2019s (CASI) annual conference and at the Aerospace Industries Association of Canada\u2019s (AIAC)Technology Collaboration Forum. He has also co-authoring of a technical report addressing the yaw activities of Bombardier\u2019s Challenger 604. Jaroslav Svoboda obtained his Dipl. Ing. Degree at the Czechoslovak Technical University in Prague and his Ph.D. at Concordia University in Montreal. His industrial experience includes research and development work in the control systems area in the Czech Republic, Switzerland, and Germany, as well as extensive consulting work in Canada and the United States. He is the author of some 85 conference and journal papers as well as of numerous techtllcal reports. He teaches mechanical engineering at Concordia University where he is also the director of the aerospace program." + ] + }, + { + "image_filename": "designv11_69_0003209_pime_conf_1964_179_275_02-Figure24.1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003209_pime_conf_1964_179_275_02-Figure24.1-1.png", + "caption": "Fig. 24.1. Model traces showing regions of stick and slip for a rolling ball", + "texts": [ + " It will be noted that the non-dimensional creep is negative for specimen E where the contact area is essentially plane and this arises from the resisting torque of the ball support bearings. For specimen A the creep is noticeably positive, the effects of geometric conformity overriding the effects of the transmitted traction. For specimen D is obtained the interesting phenomena that the creep has been virtually eliminated Val I79 Pr 33 at WEST VIRGINA UNIV on June 5, 2016pcp.sagepub.comDownloaded from 142 B. G. BROTHERS AND J. HALLING by the opposing actions of the two effects illustrated in Fig. 24.1. A theoretical treatment of this problem has already been proposed by one of the present authors (Halling (5) ) who studied the combined effect of the two conditions illustrated in Fig. 24.1 by using an approximate solution as follows. The contact zone was divided into elemental strips and each strip was analysed using a two-dimensional elastic solution for the strain pattern arising from the surface tractions. Integration of such effects for all elemental strips in the contact zone yields equilibrium equations which enable a theoretical prediction to be obtained. Using this solution applied to the geometry and loading conditions of these tests yields the results shown in Fig. 24.7, where the experimental points are shown to provide Proc Instn Merh Engrs 196465 reasonable agreement although the theory is less precise when pronounced geometric conformity exists" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002076_2008-01-1044-Figure6-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002076_2008-01-1044-Figure6-1.png", + "caption": "Figure 6: Piston transverse position", + "texts": [ + " The authors recognize that the average Reynolds lubrication equation proposed by Patir and Cheng [19] is traditionally used in such models instead of Eq. (1). However, the scope here is to investigate the effects of piston translation along the wrist-pin and the interactions of an elastic second land with the cylinder bore, thus the standard Reynolds lubrication equation is assumed to be sufficient. The contact pressure is calculated using the Greenwood-Tripp asperity contact model [6, 14]. In turn the piston wear is calculated using the Archard wear model [1]. Figure 6 shows the piston transverse position, in the planes parallel and perpendicular to the crankshaft axis. The piston transverse position is monitored on the piston\u2019s center axis relative to the cylinder bore axis. Here et is the eccentricity at the top of the second land, el is the eccentricity at the bottom of the second land, es is the eccentricity at the top of the skirt, ep is the eccentricity at the wrist-pin level, eb is the eccentricity at the bottom of the skirt, ez is the eccentricity along the wrist-pin, and is the piston tilt; el, es, ep, and can readily be calculated from et and eb" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000824_1.1695555-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000824_1.1695555-Figure1-1.png", + "caption": "FIG. 1. The semiflexible polymer chain modeled as a chain of rigid segments. For each segment, the angle un defines the angle between the segment and the z axis of the laboratory coordinate system, and fn defines the angle between the x axis and the projection of the segment onto the xy-plane of the laboratory coordinate system.", + "texts": [ + " When this field is removed, the molecules will diffuse to a state of random orientation. The decay of the birefringence of this system is closely related to the rotational diffusion properties of the molecules.11,19\u201321 While the simulation results in this work are presented within a TEB context, they are thus relevant for all problems involving rotational diffusion processes in segmented macromolecule models. We model the molecule as a chain of stiff segments connected by ball-socket joints or springs, cf. Fig. 1. The center-of-mass of each segment has the position rn in the laboratory coordinate system. The orientation is given by the components of Vn , which may be given by y-convention Euler angles Vn5(un ,fn ,cn) or\u2014in the case when the angles cn are decoupled from the rest of the coordinates\u2014by the polar angles Vn5(un ,fn). In this case the cn-coordinates can easily be integrated out of the relevant Fokker\u2013Planck equations. As this can be done with the model studied here,11 we will use polar angles to represent segment orientation in the following", + " For later use we define metric tensors associated with the coordinate sets rg , qc , and g, denoted m \u21d2 (rg), m \u21d2 (c) and m \u21d2 (g), respectively, each given by Eq. ~9!. We assume that each segment has a permanent electric dipole moment pc . upcu is the same for all segments, and the orientation of each dipole moment is fixed to the segment orientation. An external electric field can be applied to a solution of these molecules. We assume that this field is parallell to the z axis of the laboratory coordinate system, cf. Fig. 1. The potential energy of segment number n in the presence of this field is given by Vn52pc\u2022E52pcE cos u . ~10! The spring force in the needle-spring chain can be implemented as derived from a Hookean spring potential Vn ,n115 1 2kgn\u2022gn T , ~11! where k is the spring stiffness. The force associated with this potential reads Fn*52 ] ]rn* Vn ,n1152kgn\u2022S ]gn ]rn*D T . ~12! More sophisticated spring potentials, like various FENE ~finitely extensible non-linear elastic! potentials,2,22 can provide a more realistic model, but for the purpose in this paper we will use Hookean springs" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003754_aqtr.2010.5520923-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003754_aqtr.2010.5520923-Figure1-1.png", + "caption": "Figure 1. Inverted Pendulum-Cart System [7]", + "texts": [ + " (2) A general controller HSSC that is not valid only for specific parameters and/or variables of the IPCS, but works well also for different cart and pendulum masses and different lengths of pendulum. It shows almost 150% performance improvements compared to [4]. The rest of the paper is organized as follows: Section II presents the system description and dimensionless model of the IPCS. Section III describes the design of HSSC. Section IV shows the results of our simulations and compares performance with [4]. Finally, Section V summarizes our contributions together with conclusions. The IPCS shown in Fig.1 [7] contains a cart, a pendulum and a horizontal track. The cart is able to move along the horizontal track, and the pendulum can rotate around a pivot fixed on the cart. The state vector of the IPCS is [ ] s x x\u03b8 \u03b8= where\u03b8 is the angle of the pendulum, \u03b8 is the angular velocity of the pendulum, x is the cart\u2019s position and x is the velocity of the cart. The desired position of the cart can be specified at any point on the track. However, we select the origin of the track as the desired position of the cart without losing generality" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000291_ias.1995.530343-Figure8-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000291_ias.1995.530343-Figure8-1.png", + "caption": "Fig. 8: Material mesh of end ring and bar ends and coil meshes of the stator end winding", + "texts": [ + " Figure 7 shows the material mesh when the extrusion is performed in axial direction. From this figure it can be seen that a part of the iron core 11s modelled as well. 51 7 The stator end winding is not incorporated in the material mesh. The end winding is modelled as a set of current driven coils in air. This is feasible since current redistribution due to skin effect is negligible in the stranded stator end winding. In the end ring and the rotor bars, skin effect cannot be neglected. Therefore, they are incorporated in the material mesh. Figure 8 shows part of the material mesh (end ring and bar ends) generated by rotating a base plane around the center line of the shaft together with the coil meshes describing the stator end winding. Referring to figure 8, it is obvious that the generation of the stator end winding coils requires a more complex extrusion compared to the material mesh modelling. Therefore, the building of such complex models using extrusion techniques is only possible if material and coil meshes can be built separately. Because the stator end winding is not modelled in the material mesh, the coil meshes have to represent the actual end winding geometry as accurate as possible. This is not required for the coil meshes used for the excitation of the end ring and bar ends", + " Because the end ring and bar ends are modelled in the material mesh, the coil meshes for exciting them only have to be inside the materials and provide a path for the current to flow. The current occupies the full material available taking skin effect into account. B. Excitations Both stator and rotor coils are defined as current driven. In the 3D model 9 full ring segments and 2 half segments are present (figures 6,7). Therefore, 11 rotor coils are required for the current excitation. The stator winding is represented by 22 current driven coils. Only 2 end winding coils are completely inside the model, the other 20 coils are cut off at the boundaries of the model (figure 8). Figure 9 shows some of the end winding coils used for the end winding excitation Figure 9 shows the two end winding coils that are completely inside the model (coil 1 and coil 2) together with three other coils that are cut off at the boundary of the model. When referring to the cross-section of figure 6, coil 1 occupies the upper half of the first stator slot (the slot in the upper left comer) and the lower half of slot eleven, coil 2 occupies the upper half of the second slot and the lower half of slot twelf" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002791_pesc.2007.4342355-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002791_pesc.2007.4342355-Figure2-1.png", + "caption": "Figure 2. Cell packaging .", + "texts": [ + " In the power electronic system, the packaging processes are divided into two steps. The first is the cell packaging, which is defined as the packaging for original active or passive cell, as illustrated in figure 1. The power circuit part of the system consists of power semiconductor cells, capacitor cells, inductance cells or transformer cells. The driving and control circuit part includes the microelectronic circuit cells, optical cells. For instance, the active cell 3D structure for IGBT die and free-wheeling diode die is shown in figure 2 (a). The fixation space shell is used to accurate chip position arrangement and is composed of the polyimide. After the cell assembly, the shell is filled with the silicon gel to protect the die against dust and aerial discharge around die edges. When the dice operates at high temperature or has large area size, the dice electrode such as gate, emitter and the collector, can contact the Mo plate without soldering and wire-bonding process for the molybdenum TCE, 4.8e-6/K approaching to that of Si or SiC used in semiconductor dice. In contrast, when the operating temperature is lower than 200 or the area size is small, the Cu plates can replace the Mo plate to contact the dice. The metallization Al on IGBT emitter and diode anode can be cleanup, and replaced by the metallization Ti/Ni/Ag to reduce the contact resistance and contact thermal resistance. Moreover, there are about 2\u00b5m thick overlay metallization Ag on the plate surface. The flat plate capacitor cell as shown in figure 2 (b) has no other material except for Ag electrode and Mica insulation medium to form multilayer sandwich. Using this simple and cheap manufacturing process, the power semiconductor dice are packaged to realize Chip Scale Package. Because the metal contact with the dice surface without soldering and the Ag layers are used as stress snubber, the dice can expand along the contacting surface under thermal cycling. As a result, the packaged die operating at high difference in temperature without the solder crack has better reliability" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002347_15397730701404684-Figure3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002347_15397730701404684-Figure3-1.png", + "caption": "Figure 3. Elliptic law and polyhedral form.", + "texts": [ + " So the following equality holds u 1 u 2 = ( 2 1 2 2 )( p 2 p 1 ) (4) Let u and f denote the inclination angles of the tangential relative displacement u and tangential contact force p with respect to the 1-axis, respectively, then Eq. (4) becomes tan u = ( 1 2 )2 tan f (5) which indicates that the direction of tangential contact force can be different from the direction of slip for the orthotropic friction law. The following closed set is then introduced Ce = { p 1 p 2 pn f\u0303 = [( p 1 1 )2 + ( p 2 2 )2] 1 2 + pn \u2264 0 } (6) where Ce describes an elliptic cone, see Fig. 3a. For numerical simulation, a piecewise linearization approximation of the elliptic cone Ce is used in the finite element model. Figure 3b shows the cross-sectional shapes of elliptic cone and polyhedral cone for some negative pn. The linearized approximation of elliptic cone Ce can be written as Cs = { P f\u0303i p pn \u2264 0 i = 1 2 s } (7) f\u0303i = [ cos i 2 sin i i 1 ] p 1 p 2 pn T where i = ( cos2 i + 2 sin2 i )1/2 , = 1 2 , s is the number of the faces of the polyhedral cone. Let the contact relative displacements, be decomposed into two parts, one is elastic part (describing micro-slip) and another is plastic part (describing macro-slip)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003700_tac.1964.1105701-Figure3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003700_tac.1964.1105701-Figure3-1.png", + "caption": "Fig. 3-Equierg.", + "texts": [ + "964 - Short a I.) I +$+J 6 3 I? I - Fig. 3-Synthesis of the optimum control law. If the initial phase point x,] is on any of the six parabolas (solid curves) in Fig. 2 , then the indicated values for u 1 and u2 are utilized and x0 d l proceed to the origin without an>- su-itchings. If x0 is in Region I, the sequence u l = - k l , u2= - k ? is used until the u1= -k1, u 2 = +k? switching curl-e is reached, at which time \u2018 Z A ~ switches to +ki. Region 111 is symmetric with Region I. A U initial points in Region 111 use the sequence u1= + k I , zd", + "/a = k J L I / P ~ I I I u ( I T T - Iuu). This control aw may be simplified if k l / I z z = k 2 / I u d . Therefore, a l = 0 1 2 = m , and 211 = k1 sgn [x2 - x1 - (1/201)x2 I x2 I ] ~2 = kz SF [ - x 2 - x1 - ( 1 / 2 a ) z z XP I 1. Transforming to vehicle terminology, X z = k1 sgn [Q - R - (l/Za)Q I Q ] r ~ ~ ~ u = k 2 s g n [ - Q - R - ( l / 2 0 1 ) Q I Q I ] , where a = k ? I z r / P O I y y ( I,, - Iuz) . If Izz>>Iuu, 01 = k z / P o I v u = k l / P o I , , . X block diagram for this system is shown in Fig. 3 . Given initial pitch and >-ax\\- body rates, the control system will drive these rates to zero in minimum time. ACKSOWLEDGVEST The authors gratefully acknowledge many fruitful discussions with Dr. J. G. Elliott. Minimum-Energy Attitude Control for a Class of Electric Propulsion Devices L. SCHWARTZ ISTRODTCTIOK This paper sketches the derivation of the minimum-energy atti- tude regulator for a satellite about one axis for a class of electric propulsion devices, those devices which require idling power to maintain operability" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001589_iecon.2005.1569312-Figure4-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001589_iecon.2005.1569312-Figure4-1.png", + "caption": "Fig. 4. Magnetic potential due to magnet for 80 degrees position of the rotor", + "texts": [ + " RESULTS AND DISCUSSION The four pole machine dimensions which are kept constant for the computed example are [4]: Rotor outer radius: 58.5 mm Stator outer radius: 95 mm Stator or rotor axial length: 76 mm Stator inner radius: 60.4 mm Magnet depth: 6.3 mm Angular width of magnet: 50 mechanical degrees Number of turns in the slot: 20 Rated current: 15 A The stator has 3-phases in which the coil is to simple layer and 36 slots with 4 mechanical degrees width. The average remanence of NdFeBr permanent magnet is 1T and unity relative recoil permeability. Figure 3 represent the flux density for 3 positions of rotor. Figure 4 shows the magnetic potential distribution due to magnets at no load in the motor at the 80 degrees position of rotor. Figure 5 shows the cogging torque for one slot pitch. Figure 6 shows magnetic potential distribution due to currents at the same position of 80 degrees. The air gap magnetic field density obtained numerically and analytically [5] is in good agreement (Fig. 3). 0 30 60 Postion id degrees 90 -0.4 0.0 0.4 0.8 1.2 Fl ux d en si ty (T ) 5 degrees 20 degrees 0 degrees Fig. 3. Flux density for three rotor positions V" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000250_eeic.1991.162566-Figure12-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000250_eeic.1991.162566-Figure12-1.png", + "caption": "Figure 12. Turbogenerator RMS Component Locations", + "texts": [ + " With regards to generator safety, the most important sensor modules arrayed along the rotor to be mounted in cavities in the back the relatively thick field coil wedges (without compromising the structural integrity of any wedge) with the sensors viewing the stator through small apertures in the wedge faces; Figure 11 illustrates this detail for an indirectly-cooled rotor; (2) ( 3 ) (4) ( 3 ) aspect of turboscanner design, it will be necessary for: ( 1) 44 I , interconnecting wiring to be routed through or along the backs of wedges to the retaining ring region at the nondriven end (NDE) of the rotor (segmented-wedge rotors will require suitable, in-line, frictional connectors); Figure 11 illustrates this detail; support electronics to occupy unused space beneath the NDE retaining ring (by hollowing out blocks and spacers); Figure 12 illustrates this detail for a two-pole rotor; electronics power to be derived from field power through the use of a power conditioner connected to the incoming bus through a mutual protection network; Figure 12 illustrates this detail; and Turboscanner on-rotor subsystems to be spatially distributed on the rotor to preserve rotor dynamic balance. (2-Pole Rotor) Turbogenerators are nearly always used for base-load generation and are expected to run for years at a time. If it is necessary that a unit be taken off line for servicing. the procedure for removing a rotor is a time-consuming task requiring extreme care (most machines have horizontal axes and most rotors are very long compared to the diameter of the stator bores)", + " Accommodating the foregoing environmental and machine safety restrictions se;erely limits t he latitude with which to design a rotor-mounted scanner for turbogenerators. Nevertheless, it will be possible to obey all ground rules and satisfy performance specifications. To illustrate this point, the following describes the turboscanner at the conceptual design level. While it likely that the turboscanners will evolve significantly in reaching commercial status, it is equally likely that the basic approach will be preserved. Figure 12 is a simplified schematic representation of the NDE end of a two-pole turbogenerator rotor showing the generic features of such rotors and the locations of the principal rotormounted subsystems of a turboscanner. As shown in the figure (exaggerated) and in Figure 11, sensors will be mounted behind viewing apertures along four rows of wedges most likely in azimuthally-symmetric orientation with respect to the poles. Four sets of sensors (possibly two) will be provided for reasons of redundancy and mechanical balance" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002791_pesc.2007.4342355-Figure4-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002791_pesc.2007.4342355-Figure4-1.png", + "caption": "Figure 4. The acetabuliform spring stress distribution.", + "texts": [ + " The worst thermal conductivity caoutchouc prevent the heat transfer from the power circuit cells to the driving and control circuit cells which operating at lower temperature than power circuit cells. In additional, the compressed caoutchouc produces pressure on the top DBC to improve the contact performance between the die and the metal plate in the active cells. The spring material dimensions and properties are listed in table I. The acetabuliform spring mechanical simulate result in ABAQUS software, as illustrated in figure 4 shows the stress distribution with different displacement 6.4\u00b5m and 320\u00b5m. The experimental curve as shown in figure 5 indicates that the force increase linearly to 320N with displacement from 0.05mm to 0.2mm and the force is nonlinear to the displacement from 0.2mm to 0.35mm due to the spring plastic deformation focusing on the top edge of the inner round with max stress. And then, the pressure increment fast until the spring becomes a flat plate metal. In the system packaging, the spring operates with the 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002317_imece2007-42929-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002317_imece2007-42929-Figure1-1.png", + "caption": "Figure 1: C-frame ultrasonic metal welder with wedge-reed system and isolation clamp", + "texts": [ + "9 mm thick Novelis (formerly Alcan) AA6111-T4 aluminum alloy sheets were used for the study. Sheets of batch 1 material (coil i.d.1L24-1B615 BL) dated Oct2004 and batch 2 material (coil # H59740) dated Apr-2006 were sheared into 350 mm x 550 mm panels. For tensile testing some of the panels were sheared into 25 mm x 550 mm strips. The materials were used as received without any surface cleaning or preparation. Both batches were coated with oil based MP404 stamping lubricant. A Sonobond Ultrasonic wedge-reed type linear welding system shown in Figure 1(a) was used for all welding. The power controller was a model CLF 2500, S/N 12250010 and is capable of delivering up to 2500 watts of power and is designed to operate at a frequency of 20 kHz. The weld head was a purpose-built welder based on a design by Sonobond Ultrasonics and modified by Ford Motor Company. The welder's pneumatic system is capable of applying zero to 2.2 kN 2 Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 02/05/2016 T of normal clamping force along the reed", + " Clamp pressure applied to the weldment is dependent on the contact surface area of the tip used to interface with the weldment. The welder power controller can operate in one of three different weld modes: energy, time or height; energy mode was used in this study. In energy mode, weld completion is dictated by attainment of a pre-set energy level input into the weld controller by the operator. The wedge-reed components of the welding system are mounted to a steel C-frame, schematically shown in Figure 1(b). A Ford designed Type 10A-FWF (FWF-16) style contact tip geometry made from T1 tool steel was used in this study. The transducer houses the piezoelectric crystal stack that transforms high frequency electrical energy into mechanical energy. The mechanical energy is transmitted through the wedge-reed assembly in the form of vibration to the weld tip. The vibration energy is transmitted from the weld tip to weld samples retained between the tip and an anvil assembly mounted to the bottom leg of the frame" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003932_aim.2010.5695862-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003932_aim.2010.5695862-Figure2-1.png", + "caption": "Fig. 2. Measurement environment in two dimensional space.", + "texts": [ + " The cross-correlation algorithm is used to recognize the coordinates of the sound source with the time difference in the arrival time of the sound to each microphone. The distance from the sound source to the microphone is calculated using the time difference as a result of applying cross-correlation algorithm to the three microphones and the Pappus' centroid theorem. After the three distances are obtained, the coordinates of the faulty insulator are obtained by applying the cosine law, and the coordinate data are sent to the PC through the serial communication channel. Fig. 1 shows the flowchart for the overall algorithm. There are three microphones in Fig. 2, 1M , 2M , and 3M , to receive the sound source, S, which are located in a line with the same interval, 1 2 15l l cm= = . The distance between the sound source and the microphones are represented as 1d , 2d ,and 3d , respectively. The interval between two microphones is selected as 15cm which is heuristically determined to provide high accuracy in measuring the distances between 2-4 m. 1M\u03b8 , 2M\u03b8 , 3M\u03b8 represent the angles from the microphones to the sound source. The distance from the microphone to the sound source, id , can be represented as i id t vel= \u22c5 ", + " (4) where ijn\u0394 is the phase difference when there is the maximum similarity between two signals and cycle represents the sampling time of the analog to digital converter (in the real experiments, cycle =18.8679 s\u03bc ). B. Location estimation algorithm Distance measurement in two dimensional space: The distances between the sound source and the microphones are designated as 1d , 2d ,and 3d . From (1) and (2), they can be represented as 1 1d vel t= \u22c5 , (5.a) 2 1 12d d vel t= + \u22c5 \u0394 , (5.b) 3 1 13d d vel t= + \u22c5 \u0394 . (5.c) When the three microphones are aligned on the line with equal distance (In Fig. 2, 1l = 2l ), the middle line theorem of Pappus provides the following relation: 2 2 2 2 1 3 2 12( )d d d l+ = + . (6) Plugging (6) into (5), and deriving for a variable, 1d result in 2 2 2 1 12 13 1 1 2 2l d d d D \u22c5 + \u22c5 \u0394 \u2212 \u0394 = . (6) where 1D and ijd\u0394 are 1 13 122 ( 2 )D vel t t= \u22c5 \u22c5 \u0394 \u2212 \u22c5 \u0394 , (7.a) ij ijd vel t\u0394 = \u22c5 \u0394 . (7.b) The 1d in (6) represents the distance to the first microphone from the faulty insulator radiating the sound signal. The same procedures are applied for (5.b) and (5.c), and the distances to the second and the third microphones are 2 2 2 2 1 12 13 12 13 2 1 2 2 2l d d vel t t d D \u22c5 \u2212 \u22c5 \u0394 \u2212 \u0394 + \u22c5 \u22c5 \u0394 \u22c5 \u0394 = , (8" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000321_wcica.2002.1020758-Figure3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000321_wcica.2002.1020758-Figure3-1.png", + "caption": "Fig. 3 Systematic Parameters and Kinematic Variables", + "texts": [ + " When the gemel lies on any other point on the\u2019thill of the former body behind its axle, the link is the nonstandard link (see Fig. 2.a). Obviously, when the distance I between the gemel and the midpoint of the axle of the front trailer is 0-7803-7268-9/01/$10.00 0200 I IEEE. shortened to zero, the nonstandard link becomes the standard link. So the standard link is a special case of the nonstandard link substantively, and only the nonstandard link will he taken into account in this paper. 2.2 Parameters and Variables See Fig. 3, systematic parameters and kinematic variables are defined as follow: the angle that the veer wheel ti\" is a ; the distance between a gemel and the midpoint of the axle of former trailer T,., is I<., (i=l,2. . .); the distance between a gemel and the midpoint of the axle of its latter trailer Tj is J ( i = / 2 , . $ ; suppose every body has the same mechanical width d ; the comer created by Ti.] and Ti is 4j ( i= l ,2- . ) (ToistheDR). In this paper, all the discussions will he based on the following assumptions: the system always runs on the horizontal plane; only purely rolling, and spot touching take place between the wheels and the ground; the DR and trailers are all rigid and symmetrical about the longitudinal axis, and all the bodies have the same parameters: When a lTMR runs along a straight path on horizontal grand, all the bodies " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003695_17543371jset49-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003695_17543371jset49-Figure2-1.png", + "caption": "Fig. 2 Forces acting on the shell (at the beginning of the drive phase)", + "texts": [ + " Only motions in the horizontal plane are considered and the resultant forces in the forward direction only are considered in the governing equations. The zero-slip location is assumed to be at the geometrical centre of the face of the oar blade. The oars are considered to be massless and perfectly rigid. Figures 1, 2, and 3 show the major forces acting on each part of the system, with arrows showing the direction of the forces as they would be soon after the start of the drive phase. The system of equations that follows uses the directions of the arrows as a convention for positive forces. From Fig. 2 it can be seen that the equation of motion of the shell is of the form Mut \u00bc N FX Gate FFoot FDrag \u00bc N FGate cos u FFoot\u00f0 \u00de FDrag \u00f03\u00de where M \u00bc Mfixed \u00feN dM \u00f04\u00de FFoot \u00bc F I \u00fe FX Handle \u00fe FAthletes AeroDrag \u00bc F I \u00fe FHandle cos u\u00fe FAthletes AeroDrag \u00f05\u00de with the mass term M a function of time since the motion of the athlete causes different body segments to come to rest, relative to the boat, at different times during the stroke. This is discussed in greater detail below. FDrag is the total effect of hydrodynamic and aerodynamic drag components acting on the shell" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001950_1-84628-179-2_5-Figure5.3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001950_1-84628-179-2_5-Figure5.3-1.png", + "caption": "Fig. 5.3. A tandem helicopter configuration.", + "texts": [ + " Newton\u2019s equations show that the rotational component of motion in a non-inertial frame is given by I\u2126\u0307 = \u2212\u2126 \u00d7 I\u2126 + \u03c4 , where \u2126 is the angular velocity expressed in the noninertial frame; I denote the inertia of the helicopter around its centre of mass with respect to the body fixed frame and \u03c4 is the applied external torque in the body fixed frame. Finally, the full dynamic model, represented in the inertial fixed frame, is given by \u03be\u0307 = \u03c5 (5.26) m\u03c5\u0307 = RG(a, b) \u00b7 |TM | + T 2 T RE2 + mgEz (5.27) R\u0307 = R\u2126\u0302 (5.28) I\u2126\u0307 = \u2212\u2126 \u00d7 I\u2126 + [lM \u00d7 G(a, b)]|TM | + [lT \u00d7 E2]T 2 T + |QM |E3 \u2212 |QT |E2 (5.29) where \u2126 \u2208 R3 and \u2126\u0302 = \u239b\u239d 0 \u2212\u21263 \u21262 \u21263 0 \u2212\u21261 \u2212\u21262 \u21261 0 \u239e\u23a0 (5.30) A tandem rotor helicopter (Figure 5.3) uses two contrarotating rotors of equal size and loading, so there is no net yaw moment on the helicopter because the torques of the rotors are equal and opposing. Typically, the two rotors are overlapped by around 20% to 50% of the radius (r) of the rotor disk, so the shaft separation is thus around 1.8r to 1.5r. To minimize the aerodynamic interference created by the operation of the rear rotor in the wake of the front, the rear rotor is elevated on a pylon (0.3r to 0.5r above the front rotor)", + "5), and the vertical force is achieved by the change of the main rotor collective pitch. For simplicity we will present here the dynamic model of a tandem main rotor helicopter in hovering. We propose a dynamic tandem helicopter model based on Newton\u2019s equations of motion [59] with the assumptions of the standard helicopter with the following changes: 1T The nose rotor blades are assumed to rotate in an anti-clockwise direction when viewed from above and the tail rotor blades rotate in a clockwise direction, see Figure 5.3. 2T The operation of two or more rotors in close proximity will modify the flow field at each, and hence the performance of the rotor system will not be the same as for the isolated rotors. We will not consider this phenomenon to simplify the dynamical model. In order to obtain the final dynamic equations, we have separated the aerodynamic forces into two groups. The first group is composed of translational forces and the second is related to the rotational forces of motion. More details on tandem rotor helicopter dynamics can be found in [78]. Denote by TN and TT the thrust generated by the nose (N) and tail (T ) rotors respectively (Figure 5.3). These forces have no E1 component, so the thrust vectors are defined by TN = T 2 NE2 \u2212 T 3 NE3, TT = T 2 T E2 \u2212 T 3 T E3 (5.31) By simple geometric analysis we obtain expressions in terms of \u03b2, where \u03b2 is the angle between the axis E3 and the actual thrust vector: TN = |TN | sin \u03b2NE2 \u2212 |TN | cos \u03b2NE3 (5.32) TT = |TT | sin \u03b2T E2 \u2212 |TT | cos \u03b2T E3 (5.33) The thrust vector can be represented by the expression Ti = |Ti| \u23a1\u23a3 0 sin \u03b2i \u2212 cos \u03b2i \u23a4\u23a6 = |Ti| \u23a1\u23a3 0 \u03b2i \u22121 \u23a4\u23a6 (5.34) where i = N or T , and considering sufficiently small values of \u03b2i" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001563_095440605x8397-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001563_095440605x8397-Figure2-1.png", + "caption": "Fig. 2 Generating mechanism by using the ring gear cutter", + "texts": [ + " 1d). It can also be noted that vertical sections have different dimensions; those closer to the lateral edges of the special rack cutter are taller. Both lateral sides of the special rack cutter, therefore, exhibit maximum height, while the middle one of plane SQ, in Fig. 1b, also a radial section, exhibits the minimum height. In practice, the special rack cutter can be replaced by a ring gear cutter. The outline of the manufacturingmechanism for generating themodified spur gear is illustrated in Fig. 2, where A\u2013A is the rotational axis of the gear blank and B\u2013B is the rotational axis of the ring gear cutter. Here, rc and vc serve as the pitch Proc. IMechE Vol. 219 Part C: J. Mechanical Engineering Science C11604 # IMechE 2005 at RMIT UNIVERSITY on July 12, 2015pic.sagepub.comDownloaded from radius of the cutter and its angular speed respectively. The profile of the radial section (K\u2013K section) of the cutting tool is identical to the straight-edged rack cutter in Fig. 1c which generates the tooth profile of involute gears" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001755_icma.2005.1626771-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001755_icma.2005.1626771-Figure1-1.png", + "caption": "Figure 1: n-dof manipulator", + "texts": [ + " Section four will instead deals with particular aspects regarding the possibilities for a distributed detection of arm singularity occurrence, and relevant distributed control actions. Section five will then present the overall fully distributed control architecture for modular robots, together with some considerations regarding the most important aspects of its practical implementation. Finally section six will draw some conclusions, together with some indications about both the current research work activities as well as the near future ones. II.CENTRALIZED CONTROL OF A MANIPULATOR Let us consider a generic n-dof manipulator (fig.1) and start by first recalling the corresponding centralized kinematic control scheme (see fig.2), before approaching the related problem of transforming it into a computationally distributed, self organizing, control architecture, as will be later done in section 5. Within such centralized sc me, transformation matrices (1) where the internal transformation matrix is real time kno ave been he nd (respectively representing the externally assigneda position-orientation of the \u201ctool-frame\u201d w.r" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003187_mesa.2010.5552012-Figure9-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003187_mesa.2010.5552012-Figure9-1.png", + "caption": "Fig. 9. Step 2, connect with the target", + "texts": [ + " If not, then the sub-target is the globe target. As shown in Fig. 8, the large circle is the detection range of the mobile robot. The shaded parts are the obstacles which have been detected. The dotted lines are the obstacles that have not been detected. Points A, B, C, D and E are the tangent points. Step 2: The point which is apart from the tangent point at a distance of the robot size is chosen as the starting point, then repeat Step 1 until all the starting points can be connected directly with the target point. As shown in Fig. 9, points A, B and E need to search the starting points to connect to the target point. Ste\ufffd 3: Evaluate the path lengths of all possible paths by EquatIOn (7). The tangent points along the shortest path are the sub-targets. As shown in Fig. 9, point C is the sub-target for the mobile robot. Step 4: The reactive obstacle avoidance module controls the mobile robot moving toward the sub-target, until the robot reaches the global target. V. SIMULA nON EXPERIMENT We use the Player/Stage developed by the USC Robotics Research Laboratory to simulate the motion of the mobile robot using the navigation technology proposed in this paper. The robot is the Pioneer2 OX, it has 16 sonar sensors and the detection range is a circle of radius 5 m. The mobile robot is navigated from the start to the goal without colliding with obstacles in an unknown environment" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001415_ijmr.2006.010703-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001415_ijmr.2006.010703-Figure1-1.png", + "caption": "Figure 1 Schematic illustration of a typical linear motor", + "texts": [ + " Then, a grinding force model based on the relation between the radial infeed velocity and grinding force is employed. In addition to friction and grinding force, force ripple is also modelled. Finally, all of the three factors that have influences on the direct feed drive\u2013positioning and tracking have been taken into account under one framework. A linear motor can be envisioned as a rotary motor cut axially and unrolled flat. It actually consists only of the primary part \u2018stator\u2019 and secondary part \u2018rotor\u2019 as illustrated in Figure 1. The thrust is directly applied to the slide or to the object to be moved. For almost every kind of rotary motor, there is a counterpart in linear motor. The same basic technologies used to produce torque in rotary motors are used to produce force in linear motors. Similar to its rotary counterpart, a linear motor can be classified as either a DC or an AC motor, which can then be further, classified as induction motors, linear synchronous motors, or linear variable reluctance motors (Boldea and Nasar, 1997)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001480_sme-200067068-Figure8-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001480_sme-200067068-Figure8-1.png", + "caption": "Figure 8. Finite element model of the stationary part.", + "texts": [ + " Simulation results and experimental results are plotted in Fig. 7 for the rotation speed up to 7200 rpm. It may be seen that the calculated results agree well with the experimental ones. In the graph, B m n and U m n represent the balanced and unbalanced modes that have m nodal circles and n nodal diameters. Subscripts b and f denote the backward and forward waves. Natural frequencies of the entire spindle/disks assembly-shaft-housing system are then calculated with the modal analysis of the stationary part performed in finite element software ANSYS. Figure 8 is a finite element model of the stationary part. To examine the effect of the flexibility of the stationary part, the comparisons of natural frequencies of the vibration modes without and with the stationary part installed are given in Table 3, when the disks rotate at 5400 rpm. From Table 3 one may see that the flexibility of the stationary part does not affect disk vibration modes with two or more nodal diameters. It also does not affect balanced 0 0 and unbalanced 0 1 modes. The flexibility of the stationary part only affects natural frequencies of unbalanced modes" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000343_cdc.1991.261538-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000343_cdc.1991.261538-Figure2-1.png", + "caption": "Figure 2: Aerodynamic Disturbances", + "texts": [], + "surrounding_texts": [ + "tude c o ~ r o l and momentum management nK controllers stabilize the unstable Bravity gradient torques, keep attitudes and momentums due to aem dmubalw toques small, and are rob us^ to uncertainties in the moments of inertia of the space station. Struchued singular values of the closed loop system are computed to verify the robusmess of the controllers.\nIntroduction Space Station Freedom is a large manned orbiting platfom scheduled for operation in the mid 1990's. The station will operate in a nominal earth pointing orientation. gravity gradient unstable, and at altitudes where signiEcant aerodynamic disturbances cause large momentum buildup m its contml moment gyros. This momentum will be maintained within tolerable levels by balancing aero torques with gravity gradient toques. This gives rise to some fundamental control p e r h \" tradeoffs between stabilization. attihide regulation, and momentum magnitudes. This paper explores these tradmffs using H--optbim 'on. It also describes simplified controllers which achieve near-optimal performance with good robusmess properties and with significantly reduced complexity.\nThe paper is organized into four major sections. Section 2 pmvides backgmund about the station, its dynamics, and its dishubance envir~lrment. Section 3 describes control system designs and performance tdeof\u20acs. and section 4 describes mbusmess analysis with respeu to major mass property variations, using the structured singular value analysis m e ~ o l o g y . Background NomiaalOricncationS:\nAs the space station orbits the earth, the attitude controller attempts to keep the x axis pointed along the flight path. the y axis aligned with the normal to the orbit plane. and the z axis pointed towards the surface of the earth (see Figure 1). Ihe mtating reference frame centered at the space station and oriented in these directions is referred to as Verti_cal Local Horizontal (LVLH). This orientation facilitates the operatim of antemw. star senson, and other fixed expehentat packages. Howcvcr, the solar panels must be kept pointed towards the sun for maximal efficiency, so they m mounted on mtating joints to keep tMU attitude inertiauy 6x4.\nCH3076-7/91/0000-2206$01 .OO 0 1991 IEEE 2206\nDynamics: given by: The space station rigid body attitude dynamics and kinematics are\nJ + i, X J \"= 3%'% X J 5 + TE + (1) & + @ - a x ~ = O i = x , y , z\nwhere the parameters and variables are: J Body momentaf-inextia matrix E = k5e.J Rotation matrix from LVLH to body (D Body angular rate w.r.t. inertid 3 LVLH angular rate w.r.t. inertial q, = IIgdI Oxbital rate ( approx .0011 rad/sec)\nControl and disturbance torques 3 s Ta\n-\nTfie 36b2 & x J & term in the above equations is the gravity gradient toque. It is due to the fact that the gravitational pull of the earth is stronger on jwts of the station that are closer to the eanh. This toque is destabilizing when the low boom of the station is normal to the orbit plane: if one end tips down towards the earth, the gravity gradient toque will pull it down further. Note that rotation of the station about the LVLH z axis does not alter the distance of any part from the earth, so there is no gravity gradient torque about the z axis.\nDishlrbanms: The large cross sectional area of the solar panels causes large drag forces on the station. Roughly speaking, atmospheric particles experience inelastic collisions with the panels, transferring all of their momentum to the station. The resulting aero force is given by:\nE m u = P(ST9!! (2)\nwhere the parameters and variables an?: P Atmospheric density S v Panel m a times unit normal vector Station velocity (LVLH x direction) -", + "each half orbit cycle an opposite face of the solar panel is struck by the atmo~phere, so 2'1 is actually reaified as shown on figun: 2. TIE dominant terms in the Fourier series of this rectified sine wave are the dc and twice orbit rate terms. The atmospheric density also varies during each orbit since the heat from the sun expands the earth's aanosphere, bringing more of it up to the altitude of the space station Q l i s expansion of the atmosphere is called the diumal bulge). The dominant terms in the Fourier series of the density are the dc and orbital rate terms. Let & be the vector from the center of gravity to the enter of pressure on the station. The aero toque, &-,.+ x Lm, can then be represented as a periodic external disturbance with dominant terms at dc. orbit rate and twice orbit rate.. Conaol Tor~ues:\nControl toques for the attitude control system come f\" reaction control system (RCS) thrusters, and f\" control moment gyros (CMG). A CMG is a large spinning flywheel whose momentum E has constant magnitude. The direction of the momentum vector can be tumed by applying a toque x-. to the CMG. The reaction torque on the space station has opposite sign.\nBecause this mechanism of toque generation expends no fuel, the CMGs provide the baseline actuation mode for long-term station operations.\nCMG Control Requirements The control laws which command CMG toques are called the attitude and momentum management system. This system is required to stabilize the station against unstable gravity gradient torques, to maintain attitude errors and CMG momentum accumulation within acceptable bounds, and to reject strong periodic disturbances caused by the rotation of the solar panels and by variation in atmospheric density associated with the diumal bulge.\nThe system must also be robust with respect to uncertainties in the Station's moments of inertia and its flexible dynamics. The primary moment of inertia variations come from changes during the buildup sequence. from movable payloads, and uncertainties in the moments of hmia of the station itself. These combined effects can be quite large. Flexible dynamics on the other hand, do not pose significant robustness pmblems. The lowest flex modes of the station are approximately 0.6 rad/sec, which is approximately 60 times greater than the needed control bandwidth (-10 times orbit rate) and leaves ample room for gain stabilization. Hots of flex dynamics can be seen in [YLB].\nL\"d Equations of Motion Linearized equations of motion are derived in [WBWGLS]. The angles and rates e, 4 are between the LVLH reference frame and a body reference frame nominally slim with LVLH. In this new set of axes. S = e at equilibrium, J is not diagonal, and the linearized attitude equations are given by:\n. .\n(4)\n+Jyy - J,.J 3Jxy UJ) = [ 4Jxy -30, - J3 yf ] (6) +Jxz -3Jyz J,- J,\nIt is convenient to combine the 3 x 3 matrices into a 3 x 6 matrix as follows:\nUJ) = Pe(J). bo1 (7)\nScaling the Inputs, Outputs, States, and Time Because of large numerical ranges in this model, it is often convenient to normalize time by o, and inertia by J, In this new set of units. 00 = 1. J, = .95, J, .14, and J, = 1.\nThis section describes two control design approaches which are being used to synthesize control laws for the competing requirements outlined abve. The first approach is based on the formal H\" optimization theory developed over the last several years[DGKFJ. This approach was used to explore fundamental tradeoffs between the various control objectives and to rapidly produce controllers for higher-level station configuration tradeoffs and operating procedures. The second appmch is more classical in nature, consisting of separate inner-loop attitude control and outer-loop momentum managment designs. This approach produces simpler controllers, easier to schedule with mass property variations. yet possessing performance very close to the formal optimal designs, with good robusmess properties.\nThe overall design tradeoff between stabilization, attitude regulation, and momentum management is illustrated in Figure 3 below:\nWe have three attitude emrs, 8, in roll, pitch and yaw, three momentum components, fl, accumulated in the CMGs, and three control toque components, x, produced by the CMGs. The objective is to design control compensators, K(s), which keep these variables reasonably small in the face of specified disturbance torques, a. Expressed in recently popularized jargon, we want the map\nWd-lm l--f w a e , w h a w c m\nto be small, and we choose its H\" norm to be the criterion of smallness. As ususal, the weighting functions, (Wa ,wh .Wc ,wd), can be freely adjusted and serve to explore the tradeoffs between the various signals in the design problem.\nFigure 4 illustrates a typical design trial using the H\" formalism. The figure shows singular value plots of closed loop frequency response matrices from to H and to a for a case with Uniform (constant)\n2207" + ] + }, + { + "image_filename": "designv11_69_0000749_pime_proc_1973_187_029_02-Figure3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000749_pime_proc_1973_187_029_02-Figure3-1.png", + "caption": "Fig. 3. Front view of engine", + "texts": [ + " OLDERSHAW 3 RECORDS OF CLUTCH JUDDER A vehicle was set up with a special clutch cover assembly prone to judder and tested with two different driven plates in turn, one having a facing known to have poor judder performance (A) and the other having a good performance (B). The vehicle was of conventional configuration having an engine of 2.3 litres capacity at the front of the vehicle, propeller shaft and rear-wheel drive. During the tests, measurements were made of the following quantities : (1) displacement transducer fitted horizontally between the inner wing and the engine cylinder head to monitor the engine rocking on its mountings, as shown in Fig. 3, (2) displacement transducer fitted between the rear spring shackle and the axle casing to monitor the axle wind-up of the spring, and (3) accelerometer to monitor the vehicle acceleration along the longitudinal axis of the vehicle. Take-ups were carried out on various gradients, mostly in first and reverse gears as these produced the most pronounced judder. Figs 4 and 5 show typical results for facing A. Fig. 6 shows a corresponding result for facing B. It can be seen that tests using the good clutch which did not judder showed little or no engine or axle wind-up" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001205_pime_proc_1972_186_090_02-Figure17-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001205_pime_proc_1972_186_090_02-Figure17-1.png", + "caption": "Fig. 17. Comparison of combustion chamber shapes", + "texts": [ + " Side ports are less suited to high speeds because they involve sharp turns in the flow. The corresponding precessing rotor machine can use a port arrangement like that shown in Fig. 7. It features large port areas with no flow turning and no overlap. * The original Wankel engine used a moving trochoid and avoided inertial loads but the complications inherent in moving ports led to the adoption of the present configuration. (6) In the four cycle engines there is another considerable difference in the combustion chamber airflow near t.d.c. Fig. 17 shows successive combustion chamber shapes for both engines from which it can be seen that the rapid transfer of gas across the waisted portion of the Wankel engine chamber is absent from the precessing version and so is the persistence near the trailing apex seal of a pocket of gas which is remote from the combustion process until late in the cycle. These differences may help reduce heat loss and exhaust emissions when compared with the Wankel engine. (7) In the two-cycle casing machines the surfaces can be made without special generating or cam-following machinery and where flat and spherical surfaces are used a high degree of precision and good finish is obtainable by lapping" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003021_jae-2010-1280-Figure7-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003021_jae-2010-1280-Figure7-1.png", + "caption": "Fig. 7. Telescopically-folded tube model.", + "texts": [ + " When the actually normal or shear force, fn or fs, satisfies the condition of Eq. (9), and the failure of the Velcro tie element occurs, and the constrained pair of nodes belonging to Velcro tie element on the interfaces will break and the Velcro tie element will be deleted. The initial tube of telescopically-folded configuration is divided into two connected finite volumes, exterior airbag and inside airbag, by a dummy partitioned membrane allowing for a full continuity of the gas flow. The telescopically folded tube model is shown in Fig. 7. The spacing of tube wall between exterior airbag and inside airbag is lower than 3 mm. The boundary condition is similar as the models of above two tubes. The volume of folded cavity in initial stage of the inflation is 0.0012 m3, nearly one thirds of the whole volume of the tube. The deployment of an inflatable tube with different folded configurations: Z-folded tube, coiled tube and telescopically-folded tube are simulated by the CV method. Moreover, for Z-folded tube, the effects of inflation rate and membrane thickness are also discussed" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001788_j.compstruct.2006.02.029-Figure7-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001788_j.compstruct.2006.02.029-Figure7-1.png", + "caption": "Fig. 7. Energy density distributions along critical \u2018\u2018boundaries\u2019\u2019 and resulting symmetry reactions on part of the pin. Also displacements (1:1) of contact boundary is shown together with the actual contact force distribution. Four cases: (a) FE calculated solution, (b) elliptic distribution, (e) radial elliptic, and (f) radial + 10% tangential elliptic.", + "texts": [ + " The radius of the pin is for all models R = 2 mm and we assume (arbitrarily) the central part 0 6 r 6 0.2 mm to be rigidly supported. This support creates large stress concentrations close to the center, which are without interest for the present study where we concentrate on the boundary in contact with the orthotropic disc. In general with the chosen parameters the displacement of the pin is about a quarter of the orthotropic disc displacements, and these displacements are not neglected in the analysis. Fig. 7 shows the results for the pin corresponding to some of the load cases for which the results are shown in Figs. 5 and 6. In all the different cases discussed we have only varied the clearance DR and the contact size a (the external force F). The cylindrical length (thickness) t = 3.5 mm, the pin radius R = 2 mm, the constitutive matrix components for the pin as specified in (2), and the constitutive matrix components for the orthotropic disc as specified in (3) were unchanged. We now rotate the orthotropic disc 90 getting the flexible orthotropic disc direction in the direction of the external force F, and thus change (3) to Cxxxx \u00bc 7:0 1010 Pa Cyyyy \u00bc 0:27 Cxxxx Cxxyy \u00bc Cyyxx \u00bc 0:38 Cxxxx Cxyxy \u00bc 0:40 Cxxxx \u00f014\u00de The specific result shown in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001949_iembs.2006.260557-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001949_iembs.2006.260557-Figure1-1.png", + "caption": "Fig. 1 (a) Surface model created from CT images. (b) Solid model created in ANSYS. (c) Muscle and bone models. (d) Meshed model in ANSYS.", + "texts": [ + " METHODOLOGY The geometrical models were obtained from 3D reconstruction from CT images of the left leg of just died normal subject (male, age 30, height 170cm and weight 65kg), which was donated by Harbin Medical University the 2nd Affiliated Hospital . The CT images were taken with intervals of 1mm in the neutral position. The images were manual segmented with the help of MIMICS v8.10 (Materialise, Leuven, Belgiu). Then the boundaries of skeleton and skin surface were acquired using 3D reconstruction algorithm, and outputted as STL format, which can be seen in Fig.1a. Additionally, the geometry model of an assistant device, Kirschner nail, was built by CAD software according to its real dimension. 1-4244-0033-3/06/$20.00 \u00a92006 IEEE. 1735 The models, as shown in Fig1b, c and d, were created by ANSYS v5.7 (Swanson Analysis System Inc., Houston, TX). The STL files produced by 3D reconstruction contained the boundary information of geometrical models, which were made of a series of triangles. Trough some lexical analysis, we could get to know the point position and point sequence which make up of the model surface. Then in ANSYS, ANSYS Parametric Design Language (APDL), a powerful language for optimizing the FEM workflow, was employed to create the solid models for each bone and tissue. Firstly, key points were created. Secondly, triangles, which make up the surface of models, were formed through known point sequences. Last, solid models were produced from the closed surface, which can be seen in Fig1b. After all the models were finished, they were then assembled together according to their real position. Additionally, the model of an assistant device, Kirschner nail, has also been built and added to the FE model. Totally, the FE model consisted of part of femur, broken upper tibia, broken lower tibia, talus, calcaneus, Kirschner nail, muscles and other soft tissues. The bones and muscles were all embedded in a volume of soft tissues, which included fat, and other tissues whose geometry were too complicated to be reconstructed" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002428_01495730701738280-Figure8-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002428_01495730701738280-Figure8-1.png", + "caption": "Figure 8 Initial crack growth direction for the case of d = 0 5.", + "texts": [ + " Corresponding to Figures 6a and 6b, Figures 7a and 7b show the variations of energy release rate G as a function of 1 \u2212 0 for 4 different values of 0 150 \u2264 0 \u2264 180 and another 4 different values of 0 340 \u2264 0 \u2264 370 correspond to 2 peak values in Figure 2 for the case of Sr = 0 1, f = 0 1 and d = 0 5. From Figure 7a, maximum value of G(320.3Pa \u00b7m) occurs at 1 \u2212 0 = 0 for 0 = 170 and from Figrue 7b, maximum value of G (318.3Pa \u00b7m) occurs at 1 \u2212 0 = 0 for 0 = 360 . These results are shown schematically in Figure 8a. In much the same way as mentioned before, Figures 8b and 8c show the maximum values of G for a couple of crack direction ( 1 and 0 for Sr = 0 1, f = 0 7 and Sr = 0 5, f = 0 7, respectively for the case of d = 0 5. Similarly, for relatively shallow inclusion (d = 0 3), Figure 9 show schematically the maximum values of G for a couple of crack direction ( 1 and 0 for three cases of (a) Sr = 0 1, f = 0 1, (b) Sr = 0 1, f = 0 7 and (c) Sr = 0 5, f = 0 7, respectively. From Figures 8 and 9, it can be seen that the initial crack growth is caused almost at a radial direction emanating from the inclusion within 1 \u2212 0 \u2264 10 , it shows 2 possible initial crack growth points, which is chosen the first largest G max point and the second largest G max point are located at the angle 0 (130 \u2264 0 \u2264 170 and 330 \u2264 0 \u2264 360 " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000393_2001-01-1076-Figure3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000393_2001-01-1076-Figure3-1.png", + "caption": "Fig. 3 shows the polar load diagram under the", + "texts": [], + "surrounding_texts": [ + "the bearing clearance is constant, and it varies at each position in T E H L , Therefore, the effects. Wmfc the distribution of the oil film temperature exerts on the bearin\u00a7peifermancesa:re4nvestigated> OIL FILM E X T E N T - T h e modified Reynolds equation based on T E H L considers mass conservation algorithm in the boundary condition. The algorithm also is applied . on E H L , in order to confirm tie boundary condition that the mass flow is conserved o^er both the rupture;region m6 tie formation region. Fig, 4 shows the oil film extent of the whole bearing in the cases of crank angle 130 degrees and 360' degrees in T E H L . This figure is exhibited by using the ii ratio, it is found that the ol hole exists in the position of the bearing angle 150 degrees at the crank angle 130 degrees m^6 the bearing angle 20 degrees at the crank angle 360 degrees; and that the oil film extent is formed downstream from the oil hole. Also the ol film extent is formed on the main load side at each crank angle. Moreover, the oil flow into the bearing (inflow) and the oil flow out of the bearing (outflow) which were calculated by TEHL, are shown in Fig. 5. It is clear that mass conservation algorithm is considered with the modified Reynolds equation. OIL FILM THICKNESS - T h e oil film thickness at each crank angle is shown in Fig. 6. During one load cycle, the minimum oil film thickness of EHL and TEHL are 0.77 u m (crank angle 636 degrees) and 0.54 \\x m (crank angle 637 degrees), respectively. The oil film thickness of TEHL becomes approximately 22% thinner than that of EHL over the whole crank angle. Fig. 7 shows the distributions of the oil film thickness at the crank angle" + ] + }, + { + "image_filename": "designv11_69_0002364_detc2007-35166-Figure5-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002364_detc2007-35166-Figure5-1.png", + "caption": "Fig. 5 Regular and sliding mappings for belt speed 0V > .", + "texts": [ + " The switching conditions for the sliding motions are summarized as ( ) ( ) ( ) ( ) 12 12 1 2 0,T T y m y mt t\u2202\u2126 \u2212 \u2202\u2126 \u2212 \u23a1 \u23a4 \u23a1 \u23a4\u00d7 =\u23a3 \u23a6 \u23a3 \u23a6n F n Fi i (32) MAPPINGS The sliding motion is defined by Eq.(19) and Eq.(20) with initial conditions ( , , )k kt y V . For the non-stick motion, select the initial condition on the velocity boundary (i.e., ky V= ). The basic solutions in the Appendix will be used for construction of mappings. In phase plane, the trajectories in i\u2126 starting and ending at the velocity discontinuity (i.e., from \u03ba\u03bb\u2202\u2126 to \u03bb\u03ba\u2202\u2126 , , {1, 2}\u03ba \u03bb \u2208 ) are illustrated in Fig.5. The starting and ending points for mappings iP in i\u2126 are ( ), ,k ky V t and ( )1 1, ,k ky V t+ + , respectively. The stick mapping is 0P . Define the switching planes as ( ){ } ( ){ } ( ){ } 0 1 2 , , , , , ; y i i i y i i i y i i i y t y V y t y V y t y V + \u2212 \u23ab\u039e = \u2126 = \u23aa \u23aa\u039e = \u2126 = \u23ac \u23aa \u039e = \u2126 = \u23aa\u23ad (33a) ( ){ } ( ){ } ( ){ } 0 1 2 , , , , , , , , ; x i i i x i i i x i i i x x t y V x x t y V x x t y V + \u2212 \u23ab\u039e = \u2126 = \u23aa \u23aa \u039e = \u2126 = \u23ac \u23aa \u23aa\u039e = \u2126 = \u23ad (33b) where ( ) 0 limV V \u03b4 \u03b4\u2212 \u2192 = \u2212 and ( ) 0 limV V \u03b4 \u03b4+ \u2192 = + for arbitrarily small 0\u03b4 > " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001509_detc2005-84462-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001509_detc2005-84462-Figure1-1.png", + "caption": "Figure 1: Sketch of mechanism 3-RCRR", + "texts": [ + " In this paper, virtual mechanism principle (VMP) [18] is used in the kinematic analyses to build a 6-DOF virtual PM as 3-PvRCRR (Pv denotes the virtual prismatic pair added to every limb) to make the forward/reverse velocity and acceleration analyses possible. The rates of virtual pairs (Pv) must be set to be zero in order to guarantee that the kinematic solutions of the virtual mechanism (3-PvRCRR) are equivalent to those of the initial mechanism (3-RCRR). In the rear part of the paper, kinematic curves are presented based on a numeric example. As shown in Fig. 1(a), the base and moving platforms of the mechanism are connected by three limbs, each with three revolute joints and one cylindrical pair. Both platforms are equilateral triangles. Each RCRR limb can be represented with five single-DOF pairs as R1(P2R3)R4R5, as shown in Fig. 1(b), for that a cylindrical pair can be replaced by one revolute joint and a coaxial prismatic pair. The first joints (R1) in three limbs are perpendicular to the base platform. All other pairs\u2019 axes intersect at one point called rotation center Or. So the mechanism is a spherical PM if the R1 and P2 in every limb are removed. The R3 of all three limbs and the R1 of the first and second limbs are actuated. Obviously, the five input motions are not symmetrical. So the 3RCRR is not fully symmetrical even if it has symmetrical structure. As shown in Fig. 1(a), the origin, O, of the fixed coordinate frame O-xyz locates at the center of the base platform. The z-axis is perpendicular to the base platform and upward, y-axis is along the direction of the 2nd pair in the 1st limb. The coordinates frame fixed on each limb O(i)x(i)y(i)z(i) (i=1,2,3) is shown in Fig. 1(b). O(i) coincides with the rotation center Or. z(i)-axis is parallel to the z-axis, y(i)-axis directs from the A(i) to the O(i). Generally, traditional Kutzbach-Gr\u00fcbler formula is used to calculate the mobility of a mechanism. But for some PMs, especially the lower-mobility PMs (including the 3RCRR), the traditional Kutzbach-Gr\u00fcbler formula is not valid. To overcome this problem, Huang presented the modified Kutzbach-Gr\u00fcbler formula [14,15] with the screw theory. This formula is simple and efficient for the mobility analysis of PM" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001809_2005-01-0384-Figure10-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001809_2005-01-0384-Figure10-1.png", + "caption": "Fig. 10 FEM Results for Test C", + "texts": [ + " Those fluctuations were caused by the operation of the anti-lock braking system, and they indicate that tire slip occurred intermittently. Figure 9 shows an extreme braking situation. In this simulation, the vehicle was braked without cornering. The rack of the vehicle model was secured to the rack housing during the simulation. Simulation of vehicle cornering behavior The three test conditions A, B and C described in the previous section were simulated using a cluster of eight HP ES40 Alpha 833-MHz CPUs and the distributed memory processing module of LS-DYNA. Figure 10 shows an image of the simulated vehicle cornering behavior under test condition C. The vehicle body rolled to the right at a large angle. The cross-sectional shape of the right front tire at that time is shown in Fig.11. The figure shows the simulated shape of the tire that deformed in the vertical and lateral directions due to body roll and the cornering force of the tire near the contact patch. The simulated results for Tests A, B and C were compared with the driving test data. Figure 12 shows time histories of the acceleration at the vehicle\u2019s center of gravity in the case of Test A" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000107_10402000208982591-Figure10-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000107_10402000208982591-Figure10-1.png", + "caption": "Fig. 10--Sketches of a driver and follower.", + "texts": [ + " On the other hand, in case 2) it takes a relatively long time for a sul,surf'ncc nucleated crock to propagate to the surface. As a subsurf;~ce nucleated crack reaches the surface and becomes a surface onc i t may (depending on its orientation a , the magnitude and 0 tlircction of'thc frictional stress t(x ), and the direction of the load ~iiotion) Icad to n pit formation after just a relatively small numbcr of lootling cycles (due to the \"lubricant wedge effect\") or it m:iy withstand continuing loading for a long time. Based on the prccetling nnnlysis for tlie driver U;0 (see a follower sketch in Fig. 10) the surface cracks, which were nucleated as subsurface ones, open completely while the external load approaches their mouths and allow for high pressure lubricant to penetrate them. Such cracks propagate extremely fast due to large magnitude of the stress intensity factor k,- (see Section \"Numerical Results for Surface Cracks\"). Therefore, the premature pitting experimentally observed in followers is caused by fast propagation of subsurface nucleated cracks that emerge at the surface and experience the \"lubricant wedge effect" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001240_gt2006-90435-Figure4-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001240_gt2006-90435-Figure4-1.png", + "caption": "Fig. 4 The shaft beam element and the coordinate system", + "texts": [ + " The forcing vector { }dF and { }hR are given by { } [ ] [ ] )(000)sin()cos(00000 )(00000000)sin()cos( 2 22222222 2 11111111 dddd dddd Td emtt emtt F \u03b1\u03b1\u03b1 \u03b1\u03b1\u03b1 & & +\u03a9\u22c5+\u03a9+\u03a9 ++\u03a9\u22c5+\u03a9+\u03a9 = { } [ ] sin0000 21 tekrcsrcsR tth T h \u03a9\u22c5\u2212\u2212\u2212\u2212= (7) where the magnitudes of the eccentricity of driving gear and driven gear are denoted by 1e and 2e , in this study, the equal eccentricity are considered i.e. eee == 21 . 2. Shaft A two-noded element is used in the finite element formulation, it is considered for each shaft as shown in Fig. 4. Each nodal point have five degrees of freedom. The kinetic energy of the shaft for lateral-torsional motion can be expressed as { } { }\u222b \u222b \u0393\u2212\u0393+\u03a9+ +\u03a9+\u0393+++= l s l ss dsBBI IBIWVAT 0 sssssP 0 2 sP 2 s 2 sD 2 s 2 s e ))(( 2 1 )()()( 2 1 &&& &&&&& \u03b1 \u03b1\u03c1 (8) The total potential energy of a Timoshenko beam element is given by \u222b \u222b\u222b +\u2032+\u0393\u2212\u2032\u2032+ \u2032+\u0393\u2032+\u2032= l ll dsBWVGA dsGIdsBEIU 0 2 ss 2 ss 0 2 sP 2 s 2 0 s e ])()[( 2 1 )( 2 1])()[( 2 1 \u03ba \u03b1 (9) where \u03c1 , A , I , sDI and sPI are the mass density, crosssection area, area moment of inertia, transverse mass moment of inertia and polar mass moment of inertia of the shaft, respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002236_j.topol.2007.12.004-Figure13-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002236_j.topol.2007.12.004-Figure13-1.png", + "caption": "Fig. 13.", + "texts": [ + ", Sn has two boundary components for all n. Moreover, the lack of parallel stripes is also preserved hence, as in the previous example, Theorem 11 guarantees that Sn,n 1 is an incompressible surface in H2. The technique used to adopt the examples of W. Jaco (Example 29) and R. Qiu (Example 30) in our Main Construction can be used into various setting to produce new examples of incompressible surfaces in H2. We demonstrate one such setting: consider an annulus R0 properly embedded in H2 as shown in Fig. 13. To construct R0 via the Main Construction one needs one disk K1 embedded in the 3-ball C\u03b1,\u03b2 , one stripe of type \u03b11 and one elementary disk \u03b2e. The cardinality of R0 \u2229 \u2202D\u03b1 is 1 and the cardinality of R0 \u2229 \u2202D\u03b2 is 2. Since the cardinality of J0 \u2229 \u2202D\u03b1 is 3 and the cardinality of J0 \u2229 \u2202D\u03b2 is 3 (see Example 29), it is immediate that R0 is not s-isotopic to the annulus J0. Hence, by Theorem 23, R0 and J0 are not properly isotopic in H2. Then proceed as in Examples 29 and 30 to obtain surfaces Rn, n 1 properly embedded in H2, with 1 or 2 boundary components, according to whether n is odd or even" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001989_978-1-84800-239-5_34-Figure3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001989_978-1-84800-239-5_34-Figure3-1.png", + "caption": "Figure 3. Simulation for contact of two pairs of teeth", + "texts": [ + " Because the length of the long axis in the contact ellipse is much bigger than the short axis, we can consider that the pressure load only distributes on the long axis of the ellipse. The orientation of instant contact ellipse can be determined applying the relations between the surface principal curvatures and directions at the instantaneous contact point [5]. The aim of LTCA is to determine the load distribution on the instant contact line and loaded transmission errors under different working conditions. The mathematics model is given by Figure 3. Suppose there possibly are two pairs of teeth in contact ( , ). The contact lines denote the cross section in the principal direction. i is the instantaneous contact point, j is one discrete point on the principal direction. j is the initial clearance of point j; is the elastic deformation; F1 or F2 is the normal force of a pair of teeth. As Figure 3 shown, the tooth pair has an initial clearance i because of the surface modification, and the surfaces of tooth pair is right tangent. The tooth contact can be expressed as the following equations [6]: Minimize 1 1 N j j Z (3) such that -SF + e + IY + IU = (4) eTF + ZN+1=P (5) Subject to the condition that either Fk =0 or Yk =0 (k=1,2,\u2026,N) Fk 0, Yk 0, 0, Zj 0 Here, Zj (j=1,2,\u2026N+1) is artificial variable which is required to be nonnegative, N is the number of all contact pairs; S (S=Sp+Sg) is the integrated flexibility matrix of the instant contact line on the pinion and the face gear, Spij(i,j=1,2,\u2026,N) is the elastic deformation of point j on pinion surface when point i is applied a unit force in the normal direction" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003724_1.4001013-Figure9-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003724_1.4001013-Figure9-1.png", + "caption": "Fig. 9 Flow rates", + "texts": [ + " A6 \u2013 A8 in Appendix A.3 and dentities of Eqs. A16 and A17 in Appendix A.4, one can see uch a relationship between Eqs. A10 \u2013 A12 and Eqs. A13 \u2013 A15 at Y =1 /2. 3.12 Optimization of Load Capacity. Unlike the Rayleigh re summarized in Table 1. ournal of Tribology om: http://tribology.asmedigitalcollection.asme.org/pdfaccess.ashx?url=/d bearing, the finite-width bearing is difficult to optimize due to the summation form of load capacity. This complexity can be attributed to the side leakage of lubricant. Figure 9.2 in Ref. 1 shows trends of the dimensionless load capacity wzs 2 / ubl2 against H0 with four given values of =1 /4, 1/2, 1, or 2 and five values of ns=0.1, 0.3, 0.5, 0.7, and 0.9: the dimensionless load capacity monotonically decreases when H0 increases. Among all cases, a zero minimum film thickness, although vulnerable and impractical, yields the highest dimensionless load capacity. Since the designer is interested in the dimensional load capacity, the dimensional load capacity in Eqs. 44 and A16 should be maximized with a given h0", + " 8 and 9 , respectively. Similarly, the flow velocity in the radial direction can be obtained APRIL 2010, Vol. 132 / 024504-7 ata/journals/jotre9/28773/ on 03/25/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use a v A o t E a p p t h t l m p 0 Downloaded Fr ur ,r,z = uy sin + ux cos = 1 2 p r z2 \u2212 hz A1 nd the flow rate in the radial direction is the integration of the elocity with respect to z Qr ,r = 0 h ur ,r,z dz = \u2212 h3 12 p r A2 pplying the flow conservation inside the area shown in Fig. 9, ne can write Qr + Qr r dr r + dr d \u2212 Qrrd + Q + Q d dr \u2212 Q dr = 0 A3 hus, Qr r r + Qr + Q = 0 A4 quation A4 is identical to the Reynolds\u2019 equation in Eq. 1 fter one substitutes Eq. 9 and Eq. A2 in Eq. A4 . A.2 Identities of Hyperbolic Functions. tanh x/2 = cosh x \u2212 1 /sinh x A5 lim x\u21920 tanh x = x A6 lim x\u21920 coth x = 1/x A7 lim x\u21920 sinh x = x A8 A.3 Parallel-Step Slider Bearings. Analytical solutions to arallel-step slider bearings are briefly listed here in order to comare characteristics among these analytical solutions and the soluions derived in the text" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003926_s00707-009-0150-y-Figure12-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003926_s00707-009-0150-y-Figure12-1.png", + "caption": "Fig. 12 Polhode path simulation with spin polarity control", + "texts": [ + " (54) and utilizing Eq. (53) yields K = 1 + t[(1/T )(\u2202T/\u2202t)]Separatrix 1 \u2212 3 t[(1/T )(\u2202T/\u2202t)]Separatrix . (59) Numerical simulation results are presented in Figs. 10\u201313. Figure 10 shows the simulated time response of the major angular momentum for the completely viscous liquid-filled spacecraft with flexible appendage under the same initial condition as in Fig. 5. Figure 11 shows the simulated controlled time response of the major angular momentum under the same initial condition as in Fig. 5. Figure 12 shows the polhode path with spin polarity control strategy under the same initial condition as in Fig. 5. The trajectory is shown on the surface of the momentum sphere in h1 \u2212 h2 \u2212 h3 space, and the two great circles passing through the h2 axis are the unperturbed heteroclinic orbits. Figure 13 shows the simulated polhode path for a controlled spin transition with the same initial conditions as in Fig. 5. The control system is configured to ensure a positive final spin angular momentum. The polhode first crosses the separatrix at point 1", + " When h1 is zero (point 2), h3 will be less than zero. When h2 crosses zero (point 3), a thruster is fired and the polhode out-crosses the separatrix at point 4. Next, h3 crosses zero (point 5), signaling successful completion of the separatrix crossing. At point 6, h2 crosses zero, the second thruster is fired and the polhode recross the separatrix on the desired side (for comparison between Figs. 12 and 13, all the switch points are marked on the polhode in Fig. 13, while only points 3, point 4 and point 6 are marked on the polhode in Fig. 12). In this paper, the eight-dimensional ordinary differential equations governing the attitude motion of the completely liquid-filled spacecraft with flexible appendage are derived and transformed into the form suitable for the application of Melnikov\u2019s method. By using the Melnikov integral, we obtain the theoretical criteria of the chaotic attitude motion of the spacecraft. In addition, subspace of the full-parameter space are studied analytically and numerically to obtain a qualitative and quantitative understanding of the interaction of the various parameters leading to nonlinear motion" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003664_eacm.2010.163.2.83-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003664_eacm.2010.163.2.83-Figure2-1.png", + "caption": "Figure 2. Contoured joint schematisation", + "texts": [ + " Therefore, the partial loss coefficient depends also on the angular position j of the input link 1 and can be written by i,i\u00fe1 j\u00f0 \u00de \u00bc i,i\u00fe1ri,i\u00fe1 j\u00f0 \u00de i,i\u00fe1 j\u00f0 \u00de5 It is observed that to compute (j), it is necessary first to determine the reactions in the links and the relative velocities. The mechanical efficiency can be obtained as follows \u00bc 1 6 To model the internal losses in the mesh zone, a contoured joint is used. It is assumed that in a contoured joint, sliding and rolling friction occur at the same time. Let us consider a contoured joint between links i and i + 1 presented in Figure 2. According to Equation 5, the global coefficient of the internal power losses in a contoured joint is obtained by summing the partial internal losses owing to rolling and sliding j\u00f0 \u00de \u00bc vi,i\u00fe1 \u00fe \u00f8i,i\u00fe1 \u00bc i,i\u00fe1ri,i\u00fe1 v i,i\u00fe1 \u00fe f i,i\u00fe1ri,i\u00fe1 \u00f8 i,i\u00fe1 7 where Vi,i\u00fe1 is the partial internal loss coefficient owing to 84 Engineering and Computational Mechanics 163 Issue EM2 Mechanical efficiency of a spur gear system Chaari et al. sliding and \u00f8i,i\u00fe1 is the partial internal loss coefficient owing to rolling" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000876_iros.2003.1249272-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000876_iros.2003.1249272-Figure2-1.png", + "caption": "Fig. 2. Object in Graspless Manipulation", + "texts": [ + "ca(Probi)}, n(p r&i )Tf = f c o m i 5 Jmaxi} if robot finger i is force-controlled, where fmaxi is the upper limit of normal force and fcomi is the commanded normal force for robot finger i. Then we define the following matrices: [%iGzT]/- positions of robot fingers are regarded as constant. In real manipulation, however, commanded positions of robot fingers are updated step by step based on the desired object motion. Accordingly, even position-controlled robot fingers could manipulate the object. B. Mechanical Model Consider graspless manipulation of an object as in Fig. 2. We set an object reference frame whose origin coincides with the center of mass of the object. Let pen, . . , p,,,, E be positions of contact p i n t s between the object and the environment. Similarly, let prOb . . ,pmbn E R3 be positions of contact points between the object and the robot finger 1,. . . , n. We denote inward unit normal vectors at contact point p by Let us denote the sets of positions of sliding and nonsliding contacts by Cslide and CStat, respectively. We can identify whether penvi E or penvi E CStat, because the object motion is specified" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002610_arso.2007.4531422-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002610_arso.2007.4531422-Figure2-1.png", + "caption": "Fig. 2 The Developed Knee Surgical Robot Using a Hybrid Cartesian Parallel Manipulator", + "texts": [ + " System Description and Cutting Strategy Traditional milling method of knee surgical robots prepares the knee by milling off the bone in chips using the bottom edge of the mill as shown in Fig. 1 (a). No whole pieces of cut bone are remained. The proposed novel bone resection method is shown in Fig. 1(b). The side edge of the cutter, an end mill, is used to cut the bone by moving the cutter along the cutting planes of the designate shape of bone preparation. Lateral milling of bone was actuated by a hybrid Cartesian parallel manipulator, which is modified from [5] to satisfy sufficient degree of freedoms for desired position and orientation, as shown in Fig. 2. The modified Cartesian parallel mechanism has three limbs with three passive prismatic joints instead of revolute ones. The problems with revolute joints of the CPM in [5] were then avoided of its singularities and limb configuration out of the fixed frame. Furthermore, the robot is kinematically decoupled for alignment purposes and cutting-bone function. As a result, only one motor is needed to be actuated at once time so that the robot mechanism has the features of control simplicity, high positioning accuracy and safety" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002720_paciia.2008.333-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002720_paciia.2008.333-Figure2-1.png", + "caption": "Figure 2 Movement model diagram of the robot and the circular obstruction", + "texts": [ + " Selecting the middle of the front of robot, rather than the center of the front-wheel or rear-wheel of robot as the origin, is based on the following reasons: (1) Laser radar is installed in the middle of the front, the origin of the coordinates accords with the firing point of laser radar. The scanning data of laser radar are also given by polar coordinates form, the location of the obstacle can be very convenient to read out from the coordinates of laser radar. (2) In many cases, the collision between mobile robot and the obstacle occurs in the front of mobile robot possibly. The origin of the coordinates locates in front of mobile robot, which is natural reasonability and convenience in data calculation of the path length and obstacle danger. Figure 2 shows the movement model diagram of a mobile robot and the circular obstruction. suppose: V is the speed of the robot; 'V is the speed of the obstacle; V\u0394 is the relative speed of the robot to the obstacle; \u03b8 is the angle of relative speed V\u0394 and the robot speed V , \u03d5 is the angle of the obstacle speed 'V and the robot speed V ; The region that MON\u2220 covers is the collision region. If in the limited time the direction of the relative speed is adjusted to the outside of the collision region, the robot can achieve the obstacle avoidance", + " The speed and the direction of the robot can be adjusted, but the obstacle can not do any adjustment, so dV as the magnitude of V and \u03d5d as the direction of V can be adjusted, before adjustment, V is unchanged. Robot has two options to avoid the obstacle, one is MON\u2220\u2265\u03b8 where \u03b8 is the angle of relative velocity V\u0394 , the other is NOV\u2264\u03b8 , in the process of practical obstacle avoidance, the robot automatically choose the smallest rotation angle, through continuous measurement and real-time adjustment, the robot can avoid the sudden collision. In practical application, the robot can directly measure the relative speed V\u0394 , adjusting \u03d5d and dV can carry out the collision avoidance control. In figure 2, in the 'OVV\u0394 , let \u03d5\u03b8\u03b1 \u2212= , by sine theory, then: \u03b8\u03b1 sin/sinV'V = (1) For the derivative of V : \u03b8 \u03d5\u03b8\u03b8\u03b8\u03d5\u03b8\u03b1 \u03d5 \u03d5 \u03b1 \u03b1 \u03b8 \u03b8 \u03b1 \u03b1\u03b8 \u03b8 \u03b8 \u03b1 \u03b1 \u03d5 \u03d5 \u03b1 \u03b1 \u03b8 \u03b8 \u03b1 \u03b1 \u03b8 \u03b8 2sin dcossinV'dsinV'dV'sinsin df)dff(dV' V' f dfdfdfdV' V' f dfdfdV' V' fdV \u2212+= \u2202 \u2202 \u2202 \u2202+ \u2202 \u2202 \u2202 \u2202+ \u2202 \u2202+ \u2202 \u2202= \u2202 \u2202 \u2202 \u2202+ \u2202 \u2202 \u2202 \u2202+ \u2202 \u2202+ \u2202 \u2202= \u2202 \u2202+ \u2202 \u2202+ \u2202 \u2202= (2) \u03d5\u03b1\u03b1\u03b8\u03b8 \u03b8\u03d5 dcosV'dV'sindVsinVd VsinsinV' +\u2212=\u0394\u2234 \u0394=\u2235 (3) In the process of the obstacle avoidance of mobile robot, the movement of the obstacle can not be adjusted, so 0'dV = , then: \u03d5\u03b8\u03b1\u03b8\u03b8 dsinVctgdVsinVd +=\u0394 (4) Where, \u03d5Vd can be approximately viewed as the acceleration variable that is vertical to the speed V of the robot, dV is the variable of the magnitude of speed, because the robot can not adjust the magnitude and the direction of 'V , \u03d5d is the variable of the direction of the speed of the mobile robot" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002456_1.25389-Figure11-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002456_1.25389-Figure11-1.png", + "caption": "Fig. 11 Re 4:8 104. Phase portrait onto \u2019\u2013 plane (channel 1\u2013 channel 2).", + "texts": [ + " The presence of the same dominant frequency in all signals indicates the strong coupling between all of them. However, in the spectrum of the second and third harmonics are more visible than in the spectra of \u2019 and . One might notice that the scale of Figs. 7 and 8 differs from that of Fig. 9. This is due to the fact that the peaks in the spectra of \u2019 and are very sharp, and thus reach large values, whereas the peaks in the signal of are spread on larger bands of frequencies and are, therefore, lower. The projection of the attractor from the phase space onto the \u2019\u2013 plane is shown in Fig. 11. Trajectories visit, approximately, a single loop, and the spin motion at Re 4:8 104 is rather regular. It can be noticed that the trajectory appears more as a bundle than as a clear single orbit. This seems to be related to the particular configuration k Sp ec tr um of \u03d5 0.0250.020.0150.010.0050 250000 200000 150000 100000 50000 0 Fig. 7 Re 4:8 104. Power spectrum of \u2019 (channel 1) vs the reduced frequency k fc=U1. D ow nl oa de d by U N IV E R SI D A D D E S E V IL L A o n Fe br ua ry 2 3, 2 01 5 | h ttp :// ar c", + " Because, apparently, the ratio between these two values is not a simple rational number, the motion cannot be identified as periodic and is supposedly quasi-periodic. The presence of the additional higher frequency produces a more complex structure of the attractor, which is shown in Fig. 13. Themain differencewith the case at lower Reynolds number is the presence of smaller orbits within the bigger ones. The general shape of the larger orbits remains unchanged but they become slightly larger with \u2019 now ranging between 40 and 40 deg, as opposed to 30 deg previously observed. The inner region of the attractor is thus more visited than what happened in Fig. 11. It should be noticed that the smaller orbits arising in the central region have almost the same shape as the larger ones, a typical characteristic of fractals. D ow nl oa de d by U N IV E R SI D A D D E S E V IL L A o n Fe br ua ry 2 3, 2 01 5 | h ttp :// ar c. ai aa .o rg | D O I: 1 0. 25 14 /1 .2 53 89 A further increase of the Reynolds number to Re 1:12 105 carries along amore complex structure of the attractor projected onto the plane \u2019\u2013 . In Fig. 14, two big trajectories seems to be associated with three small ones" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001336_iros.2006.281764-Figure22-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001336_iros.2006.281764-Figure22-1.png", + "caption": "Fig. 22 (a)- LongitudinAl tire guni behlavioi; (I)t Longitudinal tire giii-i I)ond graphlE ilio(lel witli (lorivtive caulsalitV", + "texts": [], + "surrounding_texts": [ + "A. System Descript;ion In this part, we limited ourselves to a tire-road system, made up of six parts: the gum. the tube tire, the wave front, the wheel, the environment and the roadway as it is shown in Figure (1-a). In the following modelling development, only lonigitudinal and uormal dyuamics are taken in corisid eration for mechanical aspect due to the tire-road contact. Thermal power exchange is also modelled betwveen the gum aud the tire tube and betwveen the enviroument and the tire tube. Finally, pneumatic and hyldrodynaamnic phenomenon are described successively by the pressure variation inside the tire tuLibe auid hydrodynamics force generated by the waw front. [11: the longitudinal qarter of vehicle velocity 'Land the linear tire velocity r.w, with r the tire radiuis and considered constant due to the tire rigidity and w the angular velocity of tire. X rwx t, (1) Figure (2-a) shows the longitudinal tire gum behavior when the wheel makes an angular and translation motions after contact with the roadway. At the contact level, the generated longitudinal effort Fa is decomposed into three forces: inertial torce due to mass M, elastic force from the spring k and viscous friction force through the resistance RS. The resistance element RS is used in bond graph theory to model the active resistance which generate the entropy flow from mechanical friction ect. This irreversible transformation fomn echanical to theriral power provide the thermal flow Q to the tire tube. B. Wod Bond Grah The word bond graph represents the technological level of the model where global systern is decomiposed into six subsystems (see Figure (1 -b)). Comparing to classical block diagiram, the imrput and outpuLit of each subsystems define a power variables represented by a conjlugated pair of effort-flow labelled by a half arrow. Power variables used for the studied system are: (Force, linear velocity) (F x), (Tempernture, ThermaiFlow) (T, Q), (Pressure, Volume Flow) = (P. V) (Torque. Angnlar velocity) = (r, ). These true and psetudo bond graph variables are associated respectiviely with mechanical translation and hrydrodynamic, thermal, pneumatic, amrd mechanical rotation. C Bond Craph Modes In this subsection, dynamic bond graph models of each subsystern descriibed in Figure (1-b) are presented. C. l Tire Gum C.iLa Lonigitudinal tire gum behavior bond graph. Tire gum is considered as a viscoelastic muaterial, xhich is deforming with a behavior located between a viscous liquid and an elastic solid. Vhen the tire is in coitact with tire road, kinemnatically, one definles a slip velocity Qi, This latter is the difference between two eollinear velocities at the center of tire contact Bond graph model of this subsystem is given in Figure (2-b). Thus, the corresponding bonid graph which reflects the viscoelastic phenomenon of the tire-road contact is developed in Figure (2-b). In our case, the known and measurable inputs are respectively the vehicle velocity xLand the angular wheel velocity w. They are represented by a flow sources (SF: XL) and (SF c). Then, according to the viscoelastic characteristic of the gum, tranmsmitted mechanical power (in the case where the tire is in contact with the roadway) is decomposed in to two parts: The first one is transformed into kinetic through inertia of mass M modelled by I element and which describe the dynamic of all tire points outside of physical contact. The second part and due to the contact conifiguration, the slip velocity xi5is transformed into friction (generating a thermal powem RS element) and into elasticity modelled by a storage element C of value kL The longitudinal effort FI is estimated using bond graph I element in derivative causality. This causality conflict introduces implicit equation in numerical simulation. This is due of presence of imposed flow source (SF : XL). The slip velocity i is calculated by 0 junction and modulated transformer used to transfer anguIla-r velocity c to linear one. The deduced equation is givle in (1) The mechanical equation of the tire longitudinal motion, referring to the c;ontact suiface cotuld be synthesis from the bond graph scheme and given as follows: from 1 junction associated with C arid RS elements, we have: Es-FF S-.kx, + RSt5s friction is represented by a dissipative bond graph element (R: RN) while elasticity is represented by an element C of elasticity kT which store a pote-ntial energy. The qnarter vehicle inertia due to its mass M is represented by conservative element (I: M). (2) from 1 junction and I element in derivative causality, equation (3) is deduced: F-1 ALTL RS.x5 ka, (3) xvlere x, describes thre slip displacemernt, S6L the lorgitudinal acceleration of the vehicle and M the quarter vehicle mass. To avoid a derivative causality which introduce implicit equation in numerical simulation is introduced a pad in between (SF: XL) and (I M) element [3]. This pad is given as a rigid stiff spring-damper combination with a big value of stiffniess and friction coefficient (ft6 1 00, ke = oc). Let's consider that during a dynamnic slip motion of the wxheel, the points of the tire which do not beloing to the contact are nlot very deformed than those of contact, so it can be the physical meaning of the added pad. The bonid graph model of the longitudinal tire gum behavior incluLding the new pad is given iiin Figure (3). This last bond graph is in integral causality. Let's prove that flow f4 of element (I : M) is equal to c\u00a3Lafter adding the new pad. According to afected causality, the thermal power flow Q generated by the actix e RS element is given by equation (4) and detailed in [5]. RS is the resistance value antd a a parameter depending on the slide angle. 2Q RS.W a (4) C.1.b Normal tire guru behavior bond graph. The guLru is a viscoelastic material which can be described by (damper spring and niass) systeo of Figure (4-a). The input variable corresponds to the normal effort Fy deduteed from the pressure variationi in the tire tulbe. The viscous normal In this case, the dynamic element (I: M) is in itAegral causality, because the effort FN is known. That is why the normal mass elocity ktcan be deduced by integration. (R. Rv) is a dissipatiw f element with aniy thermal power generation. The corresponding bond graph model is given in Figure (4-b) The following mechanical equation (5) is deduced from bond graph of Figure (4-b) M.i1 + RN.( 1-2) + K.(xj X2) FN (5) x1 and zi are the vertical guim displacement alnd vertical gum velocity due to the load anid x2 and x2 are the vertical gum displacement and vertical gum velocity due to the roadway profile. C.2 Wave Front and hydrodynamic effect Two kinds of forces interaction can generate a wave front phenomena: internal forces interaction between fluid mnolecules and externld forces interaction between fluid mole cules and those of the tire circumference (see Figure (5-a)). Each fluid molecule does not ruLn out at the same vlocity [71 ard follows a velocity profile of Figure (5-a). Bond graph model of hydrodynamic phenomenon is given in Figure (5-b). When each particle located in a cross-section perpendicular to the overall flow is represented by a velocity vPector, the obtained curve from the vectors extremities represents the velocity profile of Figure (5-a). The movement of the fluid can be regarded as resulting from the slip of the fluid layers the ones on the others. The velocity of each layer is a function of distance, h of this curve in the fixed plan: Xf XLf (h). Let us consider two distant contiguous layers of fluid of dh. The friction force F. which is exerted on separation surface of these two lyers, opposes to the slip of a layer on the other. Thus, the force F. is proportional to the difference in layers velocity dif, on their surfice so and invwrsely proportionlAl to dh: Let's consider that distribution of pressure during the contact tire-road is represented by an ellipsoidal function of longitudinal and lateral positions in 3D [4], as it is shown in Figure (6). From (9), FN could be expressed as follows: 1 fj(A y) PoJf 1 a2 b2 (10) After development of equation (10), the following expression of fry is deduced:h+ dh (a) (b) FNt -. (i (1 12)3) with I describes the contact surfaice lenigth. iwhere describes the coefficient of dynamic viscosity. our case, we take the difference in layers velocity dx equal to the velocity slip x5, (Z.e velocity of the first layer is Xf 0 and for the second irf in the slip su-rfice for water height level of h). So, F1; cau be written as: Figz. 6. Pineuimiatic and tlietinical effedts in tie tube = .80. Xs h This transformation of the slip velocity to the lhdrodyniamlicfeorce FR can be represen;ted in a bond graph model of Figure (5-b), by a modulated gyrator elemient GY with the constant as a modulus. C.3 Tire tube Tire ttibe is a synthetic ruibber sheet located inside the tire and describes an elastic enclosure containing a perfect gas under pressure. The tire tube and the correspondinlg bond graph model are given in Figures (6) and (7). Tire tube exchanges the thermal power with the external envi- ronment dLue to the heat conduLctivity of its wall, anid stores two types of eneirgies, pneumatic and thernal. The thermnal exchange is modelled by the dissipation element R and the storage phenomenon is represented by the two port C element. Durinig the slip stage, the temLperature \\ariation acts on the iinternal pressure inside the tire, tube according to the perfect gas equation (8) FU.V n.,ro.T (8) with: P the gas pressure inside the tire trrbe, V the gas vTolume of the tilre tuibe, n is moles number anid T the temperature iniside the tire tube. Knowing that normal effort Fr is proportional to contact pressure P and con:tact surface A as given by equation (9). = P.A (9) with: (x, y) are the two dimensiunus contact sllrface a11d a, b are constants. Thus the bond graph model of the tire tube is given by Figure (7). Resistance (R : R1) represents the heat con- ductivity of the tire tube walls. It ensures the heat transfer with the external environment. Input model corresponds to the heat flow originally from thi frictioni of gutm Q, and the output one represents the normal effort FN5, obtained by ellipsoidal distribution of pressure due to transformer TF. The two ports C element describes entropy S and xoluume V variations inside the tire tuhoe and is detailed in The thermal flow exchanged between environment and tire tube is given by following equation (see Figuire (7)): 1 Q0- 07R 7rI N Ri (12) where: Qo exchanged heat flow between environment and tire tube, T and are respectively tire tube and en- vironnment temperatures. and the global inlet therrmal flow Q1 to the tire is calcu- lated from 0 junction: (11) Q1 = Q0 + Q (13) thus, for the inilet flow entropy S to the tire, it is calen lated by tIhe transformer TF according to Car not equation as follows: s = T (14) this transformer is added just to transform the thermal flow Q, to entropy flow S. We note that the constituitive equation of the last transformer concerns onlly the relationship between flows (x.e. temperature is not transirmed). D. Global system modelling Global Tire-road bond graph mnodel is given in Figgure (8), where interconnection appears clearly between all subsystems represented in word bond graph. For this global model, one needs two principal measurements: linear velocity of thexwhicle L, arid angular xheel velocity w. The external temperature can be added as pas rameter online or offhine and the roadway profile can be considered as external input. So, tire guLr xertical deformation xr x can be estimated through the normal eflort TFx variation dtre to variation of presstrre P inside the tire ttmbe.IThis pressure variation is caused by viation of teniperature T of the tire tube which is generated front transformation to heat flow Q of the tire-raod longitudinal friction force RS.X. In order to represent more the water presence oni roadway, waxe front subsystem car be optionally added to the global nmodel as shown in Figure (8).- III. SIMUIoATIN TESTS A. SoftwomjeIplementation Simulation step is done on a specific bond graph software SMIBOLS 2000, which is an object oriented hierarchical modelling. It alloxs users to create models using bond graph, block-diagram and equation inodels. Differential causaities and algebraic loops are solved ouLt using its powerful symbolic soltution engine. Nonlinearities and user code can be integrated in single editing IDE (integrated development environment). The iconic modelling facility allows system-morphic model layout. It also has manW post-processing facilities over the simulated result. Thanks to a developed genieric item database which consists of a set of predefined models, and has been incorporated as capsules in the software SYMBOLS 2000, the designer can easily build the dynamic models of sewvral transportation systems from the Process and Instrumentation Diagram (P&ID) just connecting different sub models. The global dynamic in syibolic format is obtained connecting diffrent icons. Behind each submodel the bond graph model is hinted. If parameter values are available, the model can be simulated using own Symbols function or Matlab-Simulink (the bond graph model can generate S -Matlab function). B. Simulation results Simulatiou parameters are: Po 2.iO (Pnscnl), V ,0,3 (iM3) To 45 (oC), 1 0, 1 (m), r = 0,25 (at), KN k -6 (N mt), M 500 (k), RS - 12 (N.s/at), RN = 9 (N.s/rat) so 0, 1 (aMt2), R1 = 0, 23 (Wqattsam.0K), A 0, 3 (mt2) Wave frotA phenomenon is present at the tire-road conAact level (see Figure (9-(a))), where the grip mechanism depends inversely on the water height. One notices a fill of effort value from 0, 63kN to 0, 58kN for a water height of 0,3mm. By increasing the water height on the roadway, the grip is degraded and the effort reached a value of 0, 52kN for 2ma/s of slip velocity and 1, 5mm of water height. In this phase, the micro-indenters are flooded, and only the macro ones continue to operate. On a wet roadway, water viscosity increases when temperature decreases. Thus, the ground becomes more slipping, and the grip potential is degraded. The maximum effort value goes from 0, i68kV to 0, 58kW when water viscosity varies from 0, 5. 10 3Pas to 3.10-3Pals with the same slip speed profile (see Figure (9 (b))). The curves of Figure (9- (c)) are obtained for an initial value of temperature inside the tire tube of 450C, and two values of ambient temperature, 150C arid 30 C. The shift between the two curves shows the influence of the ambient temperature on temperature variation inside the tire tube. For the case of 150C and for slip velocity profile of Figure (9-(c)), a temperature decreasing from 45 C to 44.92 C is noticed after a braking period where the tire ttube releases heat to the external environment because its temperature is higher than outside temperature. The temperature increases gradually in the tire tulbe during the phase of increasing slip velocity. Pressure variation is proportional to the variation in the temperature. by comparing the curve of Figure (9-(c)) with that of Figure (9-(d)). OnLe notices that pressure decreases with temperature, which occurs indeed for values of slip going from lam/s to 0.2an/s, when pressure varies from 2Bar to 1.68Bar and temperaturte goes from 450C to 44.920C). Then a progressive increase of pressure is observd when slip and temperature increase." + ] + }, + { + "image_filename": "designv11_69_0001453_s11223-006-0086-6-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001453_s11223-006-0086-6-Figure1-1.png", + "caption": "Fig. 1. Loading scheme, stress distribution diagram, and specimen shape for fatigue testing (points A and B indicate the range of failure areas along the specimen length).", + "texts": [ + " Flat cantilever specimens with a working cross section of 1 mm in thickness and 5 mm in width were tested in lateral bending using a V\u00c9DS-400A type electrodynamic vibrator under resonant transverse vibration condition. The criterion of the specimen failure was the drop in the frequency by 1% as compared to the initial resonance value, which corresponded to the occurrence of a 0.5 mm deep surface macrocrack in the working section of the specimen under testing. The specimen shape and stress distribution diagram experimentally determined as a function of bending are presented in Fig. 1. Prior to testing, specimen surfaces were polished, sharp edges were rounded off (r \u2248 0.5 mm) to eliminate stress raisers. Three types of specimens were cycled: 1) specimens prepared from the condensate; 2) specimens of the substrate material (conventional sheet of Ti\u20136Al\u20134V); and 3) the so-called composite specimens prepared from the substrate material with a 300 to 700 \u03bcm condensate layer deposited onto one side. Noteworthy is that, in the process of the strain gage calibration of specimens, no differences between the moduli of elasticity of the substrate and condensate were observed", + " Light bands on the micrographs correspond to the condensate layers enriched by aluminum. The high-cycle fatigue limit \u03c3 \u22121 was determined for 10 7 cycles. In order to obtain additional data, the specimens unfailed during the prescribed number of cycles were retested at higher loads. The fatigue test results for 30 composite specimens are given in Fig. 3. The specimens unfailed at 10 7 cycles are presented by two values of the stress amplitude: the maximum stress amplitude and stress at the fracture site during retesting (see Fig. 1). It is seen that the experimental points can be divided into two groups approximated by the curves with their corresponding fatigue limits of about 500 and 300 MPa. To clarify the causes for this scatter and determine the microstructural factors responsible for the formation of cracks of critical sizes in the specimens, fractographic studies were carried out. It was found that in the specimens of the second group, a critical crack was nucleated in the condensate, at the place where relatively large droplets appear on the specimen surface (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003261_mmar.2010.5587270-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003261_mmar.2010.5587270-Figure1-1.png", + "caption": "Fig. 1. The unicycle mobile robot", + "texts": [ + " (11) Additionally, it may be required that the terminal output y(T ) is bounded as ylb \u2264 y(T ) \u2264 yub or takes a prescribed value yd = ylb = yub. After discretization, the optimal control problem (10)\u2013(11) can be re-formulated as a constrained optimization problem, and solved using one of many available algorithms. In this paper we have employed ACADO toolkit [8], [9]. The results will be reported in the next section. For illustration and comparison of the motion planning algorithms introduced in the previous sections, we shall study a number of motion planning problems of the unicycle. The robot is shown in figure 1. Its kinematics are described by the following control system reflecting the exclusion of lateral slip of the wheels{ q\u03071 = u1 cos q3, q\u03072 = u1 sin q3, q\u03073 = u2 y = k(q) = q = (q1, q2, q3), (12) where (q1, q2) denote the robot position in the XY plane, and q3 is the robot orientation. Suppose that bounds are imposed on the 2nd component of the control function, u2(t) \u2208 [\u22125, 5], with intention of not permitting the unicycle to make rapid turns. Having incorporated these bounds, and used a regularizing term r2(u) = arctanu2, we get the extended system (5) q\u03071 = u1 cos q3, q\u03072 = u1 sin q3, q\u03073 = u2, q\u03074 = p(u2 \u2212 5, \u03b1) + p(\u2212u2 \u2212 5, \u03b1) z = (k(q), q4), (13) and regularized system (6) q\u03071 = u1 cos q3, q\u03072 = u1 sin q3, q\u03073 = u2, q\u03074 = p(u2 \u2212 5, \u03b1) + p(\u2212u2 \u2212 5, \u03b1) + arctanu2, z = (k(q), q4)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001210_tac.1972.1100049-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001210_tac.1972.1100049-Figure1-1.png", + "caption": "Fig. 1. The geometrical configuration and the coordinate system.", + "texts": [ + " In Section 11, the proper geometrical configuration, the equation of motion, and the complex transfer-function met.hod wil be formulated and demonst.rated for a free rot,ating gyro. Some control problems for rot.ating bodies will be investigated in Section 111. Suggestions for other applications will be given in the Conclusion. 11. GEONETRICAL CONFIGURATION AKD THE EQEATION OF MOTION The orientation of the rectangular body axes (X, Y b , Z b ) with respect. t o inertial coordinates (Xi, Y;, 2;) is given by t,hree Euler angles el, en, e, (see Fig. 1). The angular velocities and the moments of inert.ia in body axes are p , q, r, and I=, I = I , = Iz, respectively. A useful set. of aeroballistic axes will be defined as follows. The aeroballistic axes (Xa, Ya, 2,) identify with the body axes before the third rot.ation ea is carried out; in other words, the aeroballistic rot.ate around the body axis X6 n4t.h angular velocit,y (- 43) . The forces and moments acting on the body will be defined in the aeroballistic axes. This presentation is very convenient for fast, rotat,ing bodies ( p >> q, p >>T)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001538_20060906-3-it-2910.00121-Figure4-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001538_20060906-3-it-2910.00121-Figure4-1.png", + "caption": "Fig. 4 Reference R, differential D, and hip H coordinate frames", + "texts": [ + " On a flat surface \u03c1 is zero but becomes non-zero when one side moves up or down with respect to the other side. The differential joint \u03c1 is passive (unactuated) and provides for the compliance with the terrain. The wheels are steerable with steering angles denoted by i\u03c8 . The wheel terrain contact angle i\u03b4 is the angle between the z-axes of the i-th wheel axle frame Ai and contact coordinate frame ic as shown in Fig. 3. In order to derive the kinematics equations, we must assign coordinates frames. Fig. 4 illustrates our choice of coordinate frames for the left side of the rover. The right side is assigned similar frames. In Fig. 4, R is the rover reference frame whose origin is located on the center of gravity of the rover, its x-axis along the rover straight line forward motion, its y-axis across the rover body and its zaxis represents the up and down motion. The differential frame D has a vertical (along z-axis) offset denoted by 1k and a horizontal distance of 2k from R. The distance from the differential to the hip, denoted by 3k , is half the width of the rover. We now introduce three more frames, all of which have origin at the wheel axle" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003245_physreve.81.061704-Figure6-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003245_physreve.81.061704-Figure6-1.png", + "caption": "FIG. 6. Left: side view of the 4 layers unit cell of a Sm-CFi2 . at right top view assuming a clockwise rotation, the difference in azimuthal angles between layers 1 and 2 or between 3 and 4 is taken as 23,26 . The in-plane projection of the bissectrix of layers 1 and 2 makes the angle 0 with x.", + "texts": [ + " All these information can be gathered when writing the OOP of the phase Qij = 1 \u2212 3 2 sin2 \u2212 1/3 0 0 0 \u2212 1/3 0 0 0 + 2/3 + J 2 sin2 cos 2 0 sin 2 0 0 sin 2 0 \u2212 cos 2 0 0 0 0 0 \u2212 I sin cos 0 0 cos 0 0 0 sin 0 cos 0 sin 0 0 4 taking the definitions of and 0 given in the Fig. 5, one finds 26 that the polarization PS is proportional to I= 1 +2 cos /3 while the macroscopic quadrupole ij is a function of , I and J= 1+2 cos 2 /3 and its main eigenvector is tilted with respect to the layer normal. E. Sm-CFi2 The unit cell is commensurate to four layers with unequal changes of the azimuthal angle from layer to layer = or \u2212 see, e.g., Fig. 6 with now no net polarization at larger scale the Sm-CFi2 is not ferrielectric and a macroscopic precession of the structure around the layer normal 0 =q2 z, in almost all the studied compounds, q1 and q2 have the same sign while q2 and q1 have the opposite . All these information are gathered in the OOP of the phase, Qij = 1 \u2212 3 2 sin2 \u2212 1/3 0 0 0 \u2212 1/3 0 0 0 + 2/3 + J 2 sin2 cos 2 0 sin 2 0 0 sin 2 0 \u2212 cos 2 0 0 0 0 0 5 with the definitions of and 0 given in the Fig. 6, one finds 26 that the macroscopic quadrupole ij which has the layer normal as one of its eigenaxes is a function of and J= \u2212cos . F. Sm-Cd6 A last commensurate phase with six layers has been predicted by H&T 23,27 and recently evidenced by Shun Wan et al. 28 . We will not develop on it but it has a symmetry close to that of Sm-CFi2 and similar properties. G. Sm-C Last but not least, this phase shows a periodic precession with a short period which is not commensurate to the layer thickness. Its macroscopic OOP is simply uniaxial, Qij = 1 \u2212 3 2 sin2 \u2212 1/3 0 0 0 \u2212 1/3 0 0 0 + 2/3 6 H&T have shown that the Sm-C is fundamental for the obtention of the commensurate subphases" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002603_iita.2008.95-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002603_iita.2008.95-Figure1-1.png", + "caption": "Figure 1 Inverted pendulum system simplified model", + "texts": [ + " In the proposed algorithm, the reachability of sliding surface is completed by making the future value of sliding mode track s(k) = 0. Due to the future information of sliding mode, control signal is able to adjust immediately to prevent system states cross sliding surface, hence chattering can be avoided. 4. Simulation Here, we apply the algorithm proposed in this paper to a inverted pendulum system. In the system, the position and velocity of cart, the angle and angle velocity of pendulum are measurable. The simplified model of the system is shown in Figure 1, and system parameters are shown in Table 1. Suppose the cart moves to right side is positive direction and the pendulum rotates as clockwise has positive angle. By force bearing analysis, the dynamic equation of pendulum is ( ) ( ) ( ) ( ) ( ) ( ) ( ) ( ) ( ) ( ) 2 0 1 22 2 2 22 22 22 0 1 2 1 sin cos 3 1 cos 3 sin cos 1 cos 3 sin sin cos 1 cos 3 cos cos 1 3 p p p p p p p p c p p p p p p p c p p p c p p p p p p p c p p p p p c p p p p p m l Ku m l f x f m l x m l m m m l m l g m l m m m l m m m l g m l m l m m m l f xm l f m m um l m l m \u03b8 \u03b8 \u03b8 \u03b8 \u03b8 \u03b8 \u03b8 \u03b8 \u03b8 \u03b8 \u03b8 \u03b8 \u03b8 \u03b8 \u03b8 \u03b8 \u03b8 + \u2212 + = + \u2212 \u2212 + \u2212 + \u2212 = + \u2212 \u2212 + \u2212 + ( ) ( )2 cosc p p pm m l \u03b8+ \u2212 where x is cart displacement, x is cart velocity, x is cart acceleration,\u03b8 is pendulum angle, \u03b8 is pendulum angle velocity, \u03b8 is pendulum angle acceleration, u is the input voltage, y is the system output" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003447_1.3503875-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003447_1.3503875-Figure1-1.png", + "caption": "FIG. 1. Sketch of the geometry of the torus.", + "texts": [ + " In the coordinate system x ,y ,z attached to the torus, the problem is reduced to the problem of the flow past the torus in a time-independent domain. Assuming the flow remains axisymmetric for all time makes the toroidal coordinate system, x = c sinh cos cosh \u2212 cos , y = c sinh sin cosh \u2212 cos , 1 z = c sin cosh \u2212 cos , where 0,2 , \u2212 , and 0,2 , c 0 is the characteristic length, the natural choice. The surface = 0 defines a torus, z2+ r\u2212c coth 0 2=c2 csch2 0, and the surface = 0 defines a spherical bowl, z\u2212c cot 0 2+r2 =c2 csc2 0, where r2= x2+y2=c sinh /cosh \u2212cos . Figure 1 shows the torus with radius b=c coth and the circular cross-section radius a=c csch 0. If a and b are given, one can find c and 0 as the following: c = b2 \u2212 a2, 0 = ln b a \u2212 b a 2 \u2212 1 . The torus geometry is described by the aspect ratio parameter Ar=b /a, which is the ratio of the torus diameter 2b to the cross-section diameter 2a. In terms of the toroidal coordinates and the assumption of axisymmetry, the governing Navier\u2013Stokes equations in dimensionless form are given by v t + 1 h v v + v v + 1 c v 2 sin \u2212 v v sinh = \u2212 1 h p + 2 Re 1 h2 2v 2 + 2v 2 \u2212 1 ch sin v + 2 sinh v \u2212 2 sin v + coth h2 \u2212 1 ch sinh v + sin c2 sinh 2 \u2212 2 cosh cos + sinh2 v + 1 ch cosh \u2212 2 c2 sin2 + sinh2 + 1 c2 cosh cos \u2212 1 v , 2 v t + 1 h v v + v v + 1 c v 2 sinh \u2212 v v sin \u2212 1 h p + 2 Re 1 h2 2v 2 + 2v 2 \u2212 1 ch sin v \u2212 2 sinh v + 2 sin v + coth h2 \u2212 sinh ch v + cosh ch \u2212 2 c2 sin2 + sinh2 v + 1 c2 sin2 + cosh cos \u2212 1 + 1 \u2212 cosh cos 2 sinh2 v \u2212 sin sinh c2 v , 3 1 h v + v \u2212 2h sin c v + coth \u2212 2h sinh c v = 0, 4 where p is the pressure, v and v are the velocity components in and directions, respectively, and h=c / cosh \u2212cos ", + " The comparisons of our numerical results with the data from the literature allowed us to conclude that the numerical method and computer code are well suited and can be used to simulate the flow over a torus rotating about its centerline. The characteristics of flow past a torus rotating about its centerline at the Reynolds numbers Re=20, 30, and 40 with a rate of rotation of 0.5 2.5 for a variety of aspect ratios were studied. The torus is placed in a vertical stream from down to up of uniform flow velocity U , as shown in Fig. 1. The direction of angular velocity at the torus surface is such that the rotating surface accelerates the uniform stream on the outer ring surface due to the no-slip requirement. On the inner ring surface, the rotational velocity of the wall is opposed to the oncoming flow direction. The main aim of the present research is to find a selfpropelled regime of motion. The self-motion of the torus is caused by the propulsive fluid fluxes produced by the torus on its rotating boundary. It should be noted that the selfmotion of the body has to be considered in whole space" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002781_wcica.2008.4592987-Figure3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002781_wcica.2008.4592987-Figure3-1.png", + "caption": "Fig. 3 Wheel-terrain contact frame Ci", + "texts": [ + " S1 and S3 are the front and rear steering frames on the right, while S4 and S6 are the front and rear steering frames on the left correspondingly. Wheel 1, wheel 3, wheel 4, and wheel 6 are independently steered with the steering angles j, j=1, 3, 4, 6. Ai is defined as the driven wheel frame built on each wheel axle with the driving angle devoted by i, i=1, 2, \u2026, 6. The pose transformation matrixes between each coordinate frame can be easily described using the Devavit-Hartenberg (D-H) parameters. Consider the wheel-terrain contact relationship and the wheel slip model. Figure 3 shows each wheel-terrain contact frame Ci, i=1, 2,\u2026, 6. The z direction of contact frame Ci is perpendicular to the tangent plane of the terrain at the contact point, while the x direction is parallel to the tangent plane of the terrain. The angle of the z direction of the contact frame Ci with respect to the z direction of the driving wheel frame Ai is i, which is defined as the wheel-terrain geometric contact angle. To consider wheel slip model, label the wheel-terrain contact point coordinate frame as Ci(t\u2022 \u2022t) at time (t\u2022 \u2022t), Ci(t) or Ci at time t" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001666_fie.2006.322586-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001666_fie.2006.322586-Figure2-1.png", + "caption": "FIGURE 2. (B) THE EASY-SHUCK CORN HUSKER TO ASSIST THE HARVEST, HUSKING, AND CARRYING OF LOADS OF CORN THROUGH FIELDS BY TEAM A-MAIZE-ING.", + "texts": [ + "00 \u00a9 2006 IEEE October 28 \u2013 31, 2006, San Diego, CA 36th ASEE/IEEE Frontiers in Education Conference M4F-8 Overall, the 17 student teams, experimental and control, selected a wide mix of assistive technology problems for their projects. The common element of satisfying those in need in a third world context of some push-pull problem, tended to direct most student teams to farming, gardening, or basic transport sorts of problems. Of particular concern for many students were people who suffer from back problems. Interestingly, students did a lot of research on life in third world countries and became sympathetic with the need to carry heavy loads over rough terrains for long distances. Figure 2 (a) and (b) show early concepts for two of the projects. Observations and analysis of assessment data indicate there is a discernable greater quality of the projects of interdisciplinary teams of engineering and industrial design students as compared with engineering only teams. Data also indicates that interdisciplinary teams value and are more amenable to projects that are more complex due to being open ended, human centered, and collaborative as compared with engineering only teams. A survey was given to all the students where they were asked to rate presentations from poor to excellent for each of the seventeen (17) teams (questions 1 - 17 on the survey) as well as answering twelve additional questions (questions 18 \u2013 29 on the survey) pertaining to assistive technology projects, team dynamics and collaboration" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001652_icar.2005.1507413-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001652_icar.2005.1507413-Figure2-1.png", + "caption": "Fig. 2. Normal mechanism", + "texts": [ + " Thus, there is also a need to find most beneficial mechanism for the task to be carried out in that workspace. There are some papers about the workspace of parallel mechanism[5][6]. In this paper, we discuss the principle of workspace and manipulability calculation of adjustable parallel mechanism that can adjust the link lengths. We analyze and compare these parameters for the mechanisms set at different positions. Further, we also discuss voids of these mechanisms and method to cover the voids of one mechanism by using the other. Fig.2 shows linearly actuated parallel mechanism with six degrees of freedom. It has six links assembled in parallel configurations. On the upper end of these links is the endplate fixed. The lower ends of the links are fixed on the actuators. To achieve six degrees of freedom each of these links has individual actuators. Actuators are fixed on the base plate. There are the passive linear joints with a lock on each link. When we put the lock ON or OFF the passive linear joints, it can be fixed and released respectively. When the actuators are fixed and the lock is put OFF the passive liner joints, the link length can be adjusted passively by moving the endplate. 2020-7803-9177-2/05/$20.00/\u00a92005 IEEE We divide this mechanism into normal and transfigured types. When the endplate is horizontal at initial position and all the links are equal, we call it the normal mechanism (Figure 2). Link lengths of the normal mechanism remain fixed at all positions. The link manipulations are done using the actuators. The transfigured mechanism can be achieved in two different ways: One is offsetting the endplate by some distance and adjusting the links to hold this position (Fig.3 (a)). This transformation is called \u201dparallel transformation\u201d The other is adjusting the links upon rotating the endplate about one of the principle axes (Fig.3 (b)). This trasnformation is called \u201drotating transformation\u201d Thus, the links of the transfigured mechanisms are not equal and vary depending upon position or angle by which the end plate is transfigured" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002979_978-3-540-70534-5_7-Figure8.4-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002979_978-3-540-70534-5_7-Figure8.4-1.png", + "caption": "Figure 8.4: SoccerBot", + "texts": [ + " The network software uses a Virtual Token Ring structure (see Chapter 7). It is self-organizing and does not require a specific master node. A team of robots participated in both the RoboCup small size league and FIRA RoboSot. However, only RoboSot is a competition for autonomous mobile robots. The RoboCup small size league does allow the use of an overhead camera as a global sensor and remote computing on a central host system. Therefore, this event is more in the area of real-time image processing than robotics. Figure 8.4 shows the current third generation of the SoccerBot design. It carries an EyeBot controller and EyeCam camera for on-board image processing and is powered by a lithium-ion rechargeable battery. This robot is commercially available from InroSoft [InroSoft 2006]. LabBot For our robotics lab course we wanted a simpler and more robust version of the SoccerBot that does not have to comply with any size restrictions. LabBot was designed by going back to the simpler design of Eve, connecting the motors directly to the wheels without the need for gears or additional bearings" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002786_sice.2008.4654962-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002786_sice.2008.4654962-Figure1-1.png", + "caption": "Fig. 1. TORA system.", + "texts": [ + " Consequently, the switched nonlinear controller could be carried out and the fuzzy control rules are represented as follows: Control Rule i: If zi(t) is greater than Z and less than Z , then )(xKu i= (28) where the lower and upper bounds are defined as iZ M F\u03c3= \u2212 and iZ M F\u03c3= + , respectively. In the above, 1 2 n i j j \u03c3 \u03b1 \u03b1 \u2212 = , 1 2 n i j j \u03c3 \u03b1 \u03b1 \u2212 = , and F is the limit of the state zi(t). In this section, we employ the translational oscillator with an eccentric rotational proof mass actuator (TORA) to evidence our control strategy. The TORA is a nonlinear coupling system. This problem provides a benchmark for - 1848 - examining control design techniques within the framework of a familiar spring-mass type system [1]-[4]. Consider the TORA system in Figure 1, the oscillator consist of a cart of mass M with an eccentric rotational proof and is connected to a fixed wall by a linear spring of stiffness k. The cart travels in one-dimension only from right to left. There has a mass m connected to the proof-mass actuator within the cart and the moment of inertia about the center of mass is I. The length from the center of mass of cart to the rotating proof mass is estimated for e. Let x1 and x2 denote the translational position and velocity of the cart with 12 xx = " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002496_ijtc2007-44396-Figure15-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002496_ijtc2007-44396-Figure15-1.png", + "caption": "Fig. 15 shows the load component of the shaft, which acts as a part of load of the foil bearing. It is easy to know that the load of a single foil bearing can be calculated as half of the total weight of load component and the rotor, for two foil bearings are operating. The weight of load component, which finally determines the load of foil bearing, can be changed with the symmetrical accession or remove of the additional small mass unit. In the apparatus, the eccentricity ratio of foil bearing is measured with two displacement sensors on the shaft, which are vertical to each other. A series of experiments are done at different rotational speed with various numbers of mass units on load component. The experiment result is shown in Fig. 16.", + "texts": [ + " Two journal bearings (MWFB) are used to support the shaft system with a spiral thrust bearing in axial direction. The shaft is coated by ceramic, while the sleeve is made of the martensitic stainless steel. At the right end of the shaft, a reverse thread was processed to equip the thrust disk and a through hole was drilled for measuring the rotational speed. The sleeve was assembled under pressure and fixed with the shaft. loaded From: http://proceedings.asmedigitalcollection.asme.org/pdfaccess.ashx? Fig.15 Load component of shaft Tatble 3 Calculated load capacity in comparison with experiment Rotational speed (krpm) Eccentricity ratio Wanalysis (N) Wexp (N) Deviation (%) 50 0.95 4.0042 4.0180 0.34 50 0.9 3.5443 3.3810 4.83 50 0.85 3.1452 2.9204 7.70 50 0.8 2.7961 2.695 3.75 30 0.95 2.3806 2.891 17.6 30 0.9 2.1002 2.45 14.3 30 0.85 1.8598 2.254 17.5 30 0.8 1.6517 1.97 16.2 8 Copyright \u00a9 2007 by ASME url=/data/conferences/ijtc2007/71808/ on 06/03/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use Down According to Ref" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000512_810105-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000512_810105-Figure1-1.png", + "caption": "Fig. 1 - Principle of measurement Fig. 2 - Cutting of gear", + "texts": [ + " Based on analysis of the tooth profiles of hypoid gears, the angular position of the ideal gears and the position of the measuring probe contacting the tooth surface at the measuring location are computed first. measured with a three-coordinate measuring machine. Some experiments have been conducted and have proved the usefulness of the method. PRINCIPLE OF MEASUREMENT The cylindrical coordinate system is used for the measurement. A test gear is set on a table whose axis of rotation is the z axis. A ball point probe travels in the x-z plane as shown in Fig. 1. The measuring force is given normal to the x-z plane. Let us denote Xo as the distance from the z-axis to the center of the ball point probe contacting the reference point of a tooth and Zo as the distance from the surface of the rotational table to the center of the probe. Then Xi and Zi are denoted as those coordinates of the center of the probe in the x-z plane which contacts a point of an ideal tooth at a certain angle <&i around the z axis from the reference position. Computation of the variables Xo , Z\u00ab , Xi , Zi and 3>i is shown in the next section", + " For the convenience of measurement these vectors are transformed from the x\"y'z' coordinate into the xyz coordinate. At first the vectors G*M and n in the x\"y'z' coordinate are transformed into the x'y'z coordinate (in Fig. 4) as follows [GM]x 'y ' z = Gfc[irQ-(0, 0 , M\u00bb) = ( G M X ' , Giiy' , GMZ) [ n J x ' y ' z = n \u2022 TG = ( n x ' , n y ' , nz ) (15) where / s i n T M 0 \u2014 C O S T M \\ TQ = | 0 1 0 (16) XCOSTM 0 s i n T u / MD: the mounting distance of the gear. The position vector GMO of the reference point on the gear tooth and the unit normal vector n0 at the point shown in Fig. 1 (where \u00ab =\u00abo, s=so) are given by [ G W l x ' y ' z = ( G M O X ' , G n o y ' , G M 0 Z ) (17) [ n 0 ] x ' y ' z ~ ( n o x ' , n 0 y ' , noz) (18) t h e n X0 = / G M O X , J + GMoy'2 (19) Zo = Giioz (20) o = a r c s i n (\u2014Gsjoy'/Xo) (21) where 4>o: the angle between the center of the probe contacting the reference point and x1 direction as show in Fig. 1. The unit normal vector no is transformed into the xyz coordinate shown in Fig. 1 as follows [ n 0 ] x y z / c o s < t > 0 sino 0N = [ n 0 ] x ' y ' z I \u2014sin<&0 cos ' l ' o 0 (22) 0 In a similar manner the position vector Gjii \u00b0f t\u0302 e measuring point and the unit normal vector n*i at the point (where <*=\u00ab!, s=si) are given as [GMilx'y'z = (GMIX', Giiiy' GMIZ) (23) [nJx'y'z = (nix', n;y', niZ) (24) then we get the coordinates of the probe at measurement, Xj and Zj, and the angle <&i' for the measuring point at the reference position defined in Fig. 1. Xj = JGIHK'1 + GMiy'\"* Z i = GMiZ (14) 4> i = arc sin (\u2014 GM;y'/Xi) Thus the rotational angle i = i' sini' 0^ [n*i]xyz = [ n i ] x ' y ' z | -s in<&i ' cosi' 0 ](29) 0 0 1, Now all the values necessary for the measurement of the gear are calculated. The values for the pinion can be calculated similarly. As mentioned above the measuring point is represented by the two parameters \u00abL and si " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000815_zamm.200310098-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000815_zamm.200310098-Figure2-1.png", + "caption": "Fig. 2 Extension of Mikhlin\u2019s systems of nodal coordinates for slip-lines on a curved principal surface. The radii of curvature, R and S, are measured in the tangent plane to the principal surface.", + "texts": [ + "1) Since the normal strain-rate, \u03b5n, vanishes, \u03c4n = 0, by the flow-rule, and \u03c4\u03b1\u03b1 = \u03c4\u03b2\u03b2 = 0, by incompressibility and the flow-rule. The yield condition now becomes \u03c4\u03b1\u03b2 = 1, or \u03c4 = lm + ml, (3.2) leading to \u03c4 : \u03c4 = \u03b4, (3.3) where \u03b4 is the biaxial Kronecker tensor. Taking the scalar product of \u03c4 with (2.8) we obtain, by (3.3), \u03c4 \u00b7 \u2207\u03c3 \u2212 \u03b4 \u00b7 \u2207\u2126 = 0 or \u2207\u2126 \u2212 \u03c4 \u00b7 \u2207\u03c3 = 0. (3.4) From (2.8) and (3.2) n \u00b7 \u2207\u03c3 = 0, (3.5) while, from (3.4) and (3.2), n \u00b7 \u2207\u2126 = 0. (3.6) Torsion normal to a principal surface is impossible so that, if \u03c6 is the angle that a slip-line turns through about n (anticlockwise) in moving along that line (Fig. 2), l and m remain independent of n and n \u00b7 \u2207\u03c6 = 0. (3.7) By (B.3), the operator for divergence, \u2207\u00b7 = l \u00b7 \u2202 h\u03b1\u2202\u03b1 + m \u00b7 \u2202 h\u03b2\u2202\u03b2 + n \u00b7 \u2202 hn\u2202n , and (3.2), we obtain \u2207 \u00b7 \u03c4 + 2(\u03c4 \u00d7 n) \u00b7 \u2207\u03c6 = 0 (3.8) and, by (3.3), this becomes \u03c4 \u00b7 (\u2207 \u00b7 \u03c4 ) + 2n \u00d7 \u2207\u03c6 = 0 or, on substituting from (2.9), \u03c4 \u00b7 \u2207\u03c3 \u2212 2n \u00d7 \u2207\u03c6 = 0. (3.9) Noting that, by (3.2), (\u03c4 \u00d7 n)2 = \u03b4, c\u00a9 2004 WILEY-VCH Verlag GmbH & Co. KGaA, Weinheim we obtain, from (3.8) by (2.9), \u2212 (\u03c4 \u00d7 n) \u00b7 \u2207\u03c3 + 2\u2207\u03c6 = 0 and, on taking the scalar product of this equation with \u03c4 , by (3", + "3) c\u00a9 2004 WILEY-VCH Verlag GmbH & Co. KGaA, Weinheim ZAMM \u00b7 Z. Angew. Math. Mech. 84, No. 4 (2004) / www.interscience.wiley.com 271 and we obtain, from (3.4) and (4.1), \u22022\u03c3 \u2202\u03b1\u2202\u03b2 = 0, (4.4) while, from (3.10), (4.3), by (3.1), \u22022\u03c6 \u2202\u03b1\u2202\u03b2 = 0. (4.5) Taking the divergence of (2.8), we get, by (4.3), \u22072\u03c3 \u2212 2 h\u03b1h\u03b2 \u22022\u2126 \u2202\u03b1\u2202\u03b2 = 0. (4.6) Letting R and S equal the radii of curvature of the curves of intersection of the respective \u03b1 and \u03b2 maximum shear surfaces with the tangent plane to the n-principal-surfaces (Fig. 2), we have h\u03b1 = R \u2202\u03c6 \u2202\u03b1 , h\u03b2 = \u2212S \u2202\u03c6 \u2202\u03b2 . (5.1) From (2.9), (3.8) we obtain \u2212\u2207\u03c3 + 2 (\u03c4 \u00d7 n) \u00b7 \u2207\u03c6 = 0 (5.2) and from this equation, using (3.2) and (5.1), we may show that \u2207\u03c6 \u00b7 \u2207\u03c3 = 2 ( 1 R2 \u2212 1 S2 ) and from (5.1) we obtain (\u2207\u03c6)2 = ( 1 R2 + 1 S2 ) , while, by (5.1), (5.2), and (3.2), (\u2207\u03c3)2 = 4 ( 1 R2 + 1 S2 ) . Thus, for the angle A between the lines \u03c3 = constant and \u03c6 = constant in the principal surface, we obtain cos A = ( 1 R2 \u2212 1 S2 )/( 1 R2 + 1 S2 ) . (5.3) On the other hand, referred to slip-line coordinates, (3", + "12) describe two orthogonally intersecting families of spirals located on curved or plane principal surfaces. When R/S is constant, the angle A given by (5.3) becomes constant and eqs. (5.4), (5.5), by (4.5) or otherwise, lead to the result, \u22022\u2126 \u2202\u03b1\u2202\u03b2 = 0. (6.13) It follows from (4.6) and this result that, in the special cases governed by (6.9), \u03c3 is harmonic. In the general case (6.7) leads, by (6.3) and (6.4), to the result \u22022 \u2202\u03b1\u2202\u03b2 ln ( R S ) = 0. (6.14) In three dimensions, Mikhlin\u2019s co-ordinates [12] may be used to define the position vector, r, of a point P (Fig. 2), where the slip-lines are parallel to l and m, by r = sl + tm + \u03c1n. (7.1) If we differentiate this equation with respect to \u03b1 and take the scalar product of the resulting equation with m we obtain, by (B.1), (B.3), s \u2202\u03c6 \u2202\u03b1 + \u2202t \u2202\u03b1 = 0. (7.2) c\u00a9 2004 WILEY-VCH Verlag GmbH & Co. KGaA, Weinheim 274 R. L. Bish: Rotation-rate continuity in bi-axial plastic deformation In the same way, differentiating (7.1) with respect to \u03b2 and taking the scalar product with l, we get \u2212t \u2202\u03c6 \u2202\u03b2 + \u2202s \u2202\u03b2 = 0. (7.3) Moreover, differentiating (7.1) with respect to n and taking scalar products with l and m, we obtain, by (B.1) and (B.2), since the components of \u2207hn both vanish (see derivation of (3.1)), \u2202s \u2202n = 0, \u2202t \u2202n = 0, showing that the n-principal-lines are straight (Fig. 2). Differentiating (7.2) with respect to \u03b2 and substituting from (7.3) we obtain, by (4.5), \u22022t \u2202\u03b1\u2202\u03b2 + ( \u2202\u03c6 \u2202\u03b1 \u2202\u03c6 \u2202\u03b2 ) t = 0, (7.4) while in the same way from (7.3) we get \u22022s \u2202\u03b1\u2202\u03b2 + ( \u2202\u03c6 \u2202\u03b1 \u2202\u03c6 \u2202\u03b2 ) s = 0. (7.5) These equations have solutions similar in form to (C.1), (C.2) with (s, t) replacing (R, S) and s(0, 0) = 0, t(0, 0) = 0. On the plane boundary consisting of two orthogonally intersecting circles of radii r1, r2, shown in Fig. 3 and described by (6.10), we have s = r1 sin \u03b1 + r2(1 \u2212 cos \u03b2), t = r2 sin \u03b2 + r1(1 \u2212 cos \u03b1), wherein the net angle has been set equal, on each boundary arc, to the relevant slip-line coordinate" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001293_rob.20099-Figure3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001293_rob.20099-Figure3-1.png", + "caption": "Figure 3. A 2D proof of concept example.", + "texts": [ + " As such, the offset noted could be used as a feedback for the iterative guidance e.g., corrective actions of the target to its desired pose until such offsets reach the level of random noise. For proof-of-concept, a 2D e.g., x ,y , localization of a triangular platform is addressed in this subsection, while the generic 3D LOS-sensing-based methodology is described in Section 3. The LOS-sensing system for this particular example utilizes three modules, placed symmetrically outside the platform\u2019s workspace Figure 3 . The three detectors are placed centrally on each side of the platform, respectively. After the long-range motion of the platform, the guidance algorithm first relocates the platform so that at least one laser beam hits the center of a corresponding detector as close as random errors would allow . The subsequent corrective actions move the platform in such a way as keeping the first LOS hitting the center of its respective detector, while repositioning the platform such that the remaining LOS also hit their targets" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002285_1.2839012-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002285_1.2839012-Figure2-1.png", + "caption": "Fig. 2 Paths of contact of a crossed-axis helical gearing", + "texts": [ + "org/about-asme/terms-of-use u fl p I a u n m i w a e m fl v l t i p a A w d w f 0 Downloaded Fr p = 2 cos N 1 As soon as the axis of the gear is directed in either way by a nit vector n, a tooth flank is a left-hand flank or a right-hand ank according to whether the following quantity is negative or ositive q P \u2212 O \u00b7 n 2 n Eq. 2 , P\u2212O is the vector from a point O on the gear axis to point P on the tooth flank, whereas q is the outward-pointing nit vector orthogonal to the tooth flank at point P. Two meshing helical involute gears\u2014gear 1 and gear 2\u2014are ow considered. As is known, in order for the gears to mesh, they ust have the same normal base pitch. This condition translates nto the following equation see Eq. 1 N1 N2 = 1u1 2u2 3 here quantities ui i=1,2 are defined by ui = cos i i = 1,2 4 With reference to Fig. 2, the distance between the skew gear xes is denoted by a0. As soon as the axis of gear 1 is directed in ither way by unit vector n1, unit vector n2 is so directed as to ake a left-hand flank of a tooth of gear 1 contact a left-hand ank of a tooth of gear 2 which also implies that if the angular elocity vectors of gear 1 is positive with respect to n1, the anguar velocity of gear 2 is negative with respect to n2 . The common perpendicular to the gear axes intersects the axes hemselves at points A1 and A2" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002301_0020-7403(67)90046-x-Figure4-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002301_0020-7403(67)90046-x-Figure4-1.png", + "caption": "FIG. 4. D i rec t ion of neut ra l axis.", + "texts": [ + " ~ (z a + z'~) (dO) 2 (5) ~2 d~, = ~ (za+ z'q) dO (6) Equa t ions (5) and (6) combine to show t h a t d~ d~ = 0, to f i rs t -order quant i t ies so t h a t a t a n g e n t to the cent ro ida l locus a t GT m u s t be paral lel to A ' - A , the neu t ra l axis. Since G T is the po in t on the locus giving r th is p roves the s t a t e m e n t m a d e above. Thus : The neutral axis is the area bisector parallel to the tangent to the centroidal locus at the po in t corresponding to the M ~ value. This is i l lus t ra ted in Fig. 4. The centroidal locus has been cons t ruc ted , and the value of M~ for an appl ied couple wi th a hor izonta l axis is given by the vert ical radius vec tor GQ : M ~ = a~ A r The neu t ra l axis, as shown, is the area bisector paral lel to the t a n g e n t a t Q. C O N V E X I T Y A N D C O N T I N U I T Y O F T H E C E N T R O I D A L L O C U S I f ?) and 5 are considered as funct ions of the p a r a m e t e r 8, we have , adop t ing the no t a t i on d~/dO = ~), etc., = - - - . , , _ y z - - y z ~3 At the po in t GT we have a l ready seen t h a t ~ = 0 ; thus a t GT d ~ /~ h V = ~ (7) We m a y e x p a n d d", + " The p r e sen t resu l t s could in fac t be r ead i ly de r ived in th i s way , b u t i n s t ead t h e y h a v e b e e n deduced f rom t h e l imi t ing case of t he p l a n e sect ions a s s u m p t i o n (so t h a t all p o i n t s in t h e e rom-see t ion are a t t h e yield s t ress in e i t h e r t ens ion or compress ion) t o g e t h e r w i t h equ i l i b r i um; no f u r t h e r pos tu l a t e s h a v e b e e n necessary . P U R E F L E X U R E F o r c o n t i n u o u s cross-sect ions of f ini te w i d t h in all d i rec t ions i t was e s t ab l i shed a b o v e t h a t t he een t ro ida l locus is a s ing le -va lued c o n t i n u o u s closed curve . I t m u s t enclose t he cen t ro id G, w h i c h is i ts cen t re of s y m m e t r y a n d b isec ts d i a m e t e r s in all d i rect ions . A rad ius vec to r c an the re fo re be d r a w n to tho locus f rom G in a n y d i rec t ion , see Fig. 4, a n d t he co r re spond ing m o m e n t a NAr will be a p p r o p r i a t e to ful ly p las t ic bend ing . The re can be no s h a r p corners on t he locus, for b y (6), dS/dS~O, so t h a t two a r b i t r a r i l y close n e u t r a l axes c a n n o t give t h e same p o i n t on t h e locus a n d hence t h e t a n g e n t a t a g iven p o i n t mt~st be u n i q u e l y defined. T h u s to a n y axis of app l ied m o m e n t t he re cor responds one a n d only one possible n e u t r a l ax i s : a p u r e l y f lexural couple can a lways p roduce pu re f lexure" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002496_ijtc2007-44396-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002496_ijtc2007-44396-Figure2-1.png", + "caption": "Fig. 2. The sketch of the foil deformation and air pressure", + "texts": [ + " The bearing radial clearance between shaft surface and bearing surface is designed to be 20\u00b5m [6]. a. Structure of MWFB b. b. The geometry of Foil SOLUTION OF REYNOLDS\u2019 EQUATION Reynolds\u2019 Equation The air pressure distribution of foil bearing is yielded from the Reynolds\u2019 equation by taking the air as an ideal compressible gas flow. The dimensionless compressible Reynolds` equation under isothermal condition is given as [2] ( )3 3p p ph ph ph z z\u03b8 \u03b8 \u03b8 \u2202 \u2202 \u2202 \u2202 \u2202 + = \u039b \u2202 \u2202 \u2202\u2202 \u2202 (1) where 2 06 , , , a a p h z R p h z p C R p C \u00b5 \u03c9 = = = \u039b = (2) Then, according to Fig. 2, the film thickness can be given as following: 1 2 0 ( , ) ( , ) 1 cos( ) S z S z h C \u03b8 \u03b8 \u03b5 \u03b8 \u03b8 + = + \u2212 + (3) 0\u03b8 is the angular position of the minimum film thickness, while 1( , )S z\u03b8 and 2 ( , )S z\u03b8 are the deformations of foils nloaded From: http://proceedings.asmedigitalcollection.asme.org/pdfaccess.ashx? (shown as the dashed line in Fig. 2), calculated from elastic deformation equation and explained in the next section. The foil bearing essentially does not generate subambient pressures and there is no bearing load at the position of S\u03b8 , so, the air pressure at the top of bearing equals to pa. In circumference direction, the film pressure will increase due to the relative motion of shaft and foil. At an unknown angular position 2\u03b8 , the film pressure will decrease to ambient, called Reynolds\u2019 boundary condition. Here the pressure must fulfill both the ambient pressure and zero pressure gradient boundary condition" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000639_bf02984317-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000639_bf02984317-Figure1-1.png", + "caption": "Fig. 1 A three-joint link robot", + "texts": [ + " This application verifies that the constrained motion of nonholonomic system is efficiently obtained by the proposed method. A three-joint l ink robot The equation of motion for this system is derived by Newtonian or Lagrangian mechanics as [A, ~,C,,~,C,,I[~,] [ o ~,s,,-~s4\u00a2,] / Cl+C, - c , o ]lo,] + -c, c,+c,-C~llO, I o -C3 C3 J[t)3J + - ~ ~+~-~/ /~/+/~c~/=/P~/ o -/G /G JLO3J [G~cos &j [Fq (30) where Al= Ii + ( ~ + mz + ml ) ll 2 A2=I2+(~+m3)l~, A s = I 3 + ~ l~ Consider a three-joint link robot shown in Fig. 1 moving in the XY-plane. Rigid bar 1 has length /1, mass rnl, and moment of inertia It. Bar 2 has x2, m2, and /2, and bar 3 has/3, m3, and /3. Figure 1 shows a rotational spring and a dashpot at each joint, and shows known or unknown disturbances acting on the system. K1, K2, a n d / ~ denote spring stiffnesses, and C1, C2, and Cs are damping coefficients. This system is a highly nonlinear system described by q ( t ) = [01 02 03] r. Bx=( m~+m3)lll2 m3 1113, m3 &13 B~=~- B3=~- G,=( ~-+m2+m3)gll G2=(~+ma)gl2 ' ma G3=~- gl~ C , j = c o s ( 8 , - Os), S ~ = s i n ( O~ - Os) P l = F l l x s i n alt cos O~-F2lx s in flxt sin 0 + Hzlx sin a2t cos Ox-H2l l sin/~2t sin 01 P , = H t l sin a~t cos &-H2/2 sin/32t sin 02, P3=0 Let the Cartesian coordinate of end-effector be (xe, ye)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001701_amc.2006.1631675-Figure5-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001701_amc.2006.1631675-Figure5-1.png", + "caption": "Figure 5. Planar mobile manipulator", + "texts": [ + " Indeed, once a corridor is chosen, it will constitute a search space in which elements of sets SPp and SPa will be randomly selected to produce trial paths q( ). Although, initially this path might present undesirable distortions, they will be attenuated progressively as the process converges. Here, the optimization is performed using a simulated annealing method which is known for its efficiency when exploring large solutions spaces [38], [39]. In this section, the proposed method is applied to plan the minimum-time trajectory of the 2-link planar nonholonomic mobile manipulator (Figure 5). Parameter values are grouped in Table I. Where: 2Lw Lv: dimensions of the rectangular platform (see Figure 5). r is radius of the wheel, L1 and L2 are the length of the links, d is the distance between middle of the wheel axis and first joint of the manipulator, mp is the mass of the platform, mw is the mass of the wheel, m1 and m2 are the mass of the link, Izp is the inertia moment of the platform around z\u2019-axis, Izw and Iyw are inertia moments of the wheel around z\u2019-axis and y\u2019-axis in the coordinate system respectively, Iz1 and Iz2 are inertia moments of the links around z\u2019-axis. Constraints on joints position and driving torques are given as follows: 90,90 21 aa qq )(0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003301_amc.2010.5464114-Figure8-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003301_amc.2010.5464114-Figure8-1.png", + "caption": "Fig. 8. The simplified diagram of the systems. (a) Master system. (b) Slave system.", + "texts": [ + " We assume that person moves only PIP (proximal interphalangeal) joint and DIP (distal interphalangeal) joint, and the other joint is fixed. Fig. 7 shows a finger model used in this paper. This model is a human joint model, so the estimated angle and the estimated torque are calculated based on this model. The middle phalanx and distal phalanx is assumed link 1 and link 2 respectively. In Fig. 7, \u03b8, m, and l stand for angle, mass, and length of link respectively, subscript 1, 2, and fix stand for link 1, link 2, and angle of proximal phalanx, which is fixed, respectively. In this paper, the master system is assumed like Fig. 8 (a) and the slave system is assumed like Fig. 8 (b). In the master system, human puts on the rings which are connected on the linear motors by the universal joint. And the y-axis position and the y-axis applied force are obtained by the linear actuator. In here, it is assumed that the rings are contact with the center of the middle phalanx and the distal phalanx. The master system represents the exoskeleton system. In the slave system, the end effector of x-y table reproduces the trajectory and the applied force in the finger tip point. This system represents the endoskeleton system" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000710_robot.2004.1307395-Figure8-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000710_robot.2004.1307395-Figure8-1.png", + "caption": "Fig. 8. Snakebard", + "texts": [ + " The computational efficiency for our motion planner for the snakeboard model can be further improved by using fast collision detection routines for circular arc paths as described in [ 191. VI. EXPERIMENTAL RESULTS In this section, we describe the experimental setup on which we have tested our motion planning results. In Section VI-A, we describe the snakeboard prototype that we have built in our lab. The experimental results for the motion planner are shown in Section VI-B. A. Descriprion of the Snakeboard We have built a snakeboard model consisting of four wheels that are always in contact with the ground (for further details on the snakeboard, see [24] and [17]). As shown in Figure 8, the two \u201csteering wheels\u201d are coupled and rotate symmetrically. The two remaining wheels, called the \u201ctraining wheels,\u201d are fixed parallel to the snakeboard\u2018s line of motion. Encoders are attached to the two training wheels to track the snakeboard\u2018s position and velocity. Two actuators are located on the top of the snakeboard. One motor controls the angle of rotation of the two coupled steering wheels. The other motor is used to rotate the rotor on the underside of the snakeboard. Both motors are equipped with internal encoders to track displacement and velocity" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001455_s00542-006-0293-x-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001455_s00542-006-0293-x-Figure1-1.png", + "caption": "Fig. 1 Experimental apparatus", + "texts": [ + " Moreover, Hashimoto developed the prediction model for traction coefficient based on the contact mechanics between paper web and steel roller (Hashimoto 2005a, b; Hashimoto and Okajima 2006). However, as far as the authors know, there is no literature treating the method how to estimate the slip onset of velocity between the web and roller (Hashimoto 2005a, b). This paper describes two kinds of theoretical model for predicting the slip onset velocity between the paper web and steel roller with experimental verifications. Figure 1 shows the test rig for observing the onset of slippage. The endless type of web is supported by the five rollers. The web tension is applied by the tension control roller, and the magnitude of web tension is measured by the load cell and displayed on the tension unit. The rotating roller is driven by the driving motor through belt. The web velocity is measured by LDV. The test conditions are listed in Table 1. The slip ratio between the web and roller is defined as: s \u00bc Ur Uwj j Ur \u00f01\u00de Two kinds of paper, uncoated paper (newsprint) and coated paper, are used as test webs" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001652_icar.2005.1507413-Figure5-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001652_icar.2005.1507413-Figure5-1.png", + "caption": "Fig. 5. Method to transfigure the mechanism by parallel trasnformation", + "texts": [ + " To compare the workspace of normal mechanism with the transfigured mechanisms, we define the ratio of workspace volume U as follows: U = Vt Vn (1) Where Vt is the workspace volume of transfigured mechanism and Vn is the workspace volume of normal mechanism. We calculate this ratio for all 4 types of workspaces to find the relative properties of normal and transfigured mechanisms at various positions. Parallel trasformation is done by offsetting the endplate by the angle \u0398 from y-axis and distance d from the initial position of the endplate (Fig.5). We then calculate the translational workspace that is all reachable points of the mechanism in x, y and z axes for various \u0398 and d. Fig.6 shows the translational workspace of the normal mechanism. It was found to be 120[deg] symmetric on x-y plane. Thus the positions for transfiguring the mechanism were limited to \u0398 = 0 to 60[deg]. 1) Translational workspace: We calculated the translational workspace for the different mechanism by adjusting the endplate at various distance d in this range of \u0398" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000597_cbo9780511547126.019-Figure17.6-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000597_cbo9780511547126.019-Figure17.6-1.png", + "caption": "Figure 17.6.2: Determination of disk surface D: (a) and (b) installment of generating disk; (c) line L\u03c3 D of tangency of surfaces \u03c3 and D; (d) illustration of generation of surface \u03c4 by disk surface D.", + "texts": [ + " It follows from the discussion above that design of a gear drive formed by a double-crowned pinion and a profile-crowned gear is the precondition for reduction of noise and vibration and localization of bearing contact. 17.6 LONGITUDINAL CROWNING OF PINION BY A PLUNGING DISK Longitudinal crowning of the pinion tooth surface, in addition to profile crowning, is applied for transformation of the shape of the function of transmission errors and reduction of noise and vibration. We recall that errors of shaft angle and lead angle cause a discontinuous linear function of transmission errors [see Section 17.5 and Fig. 17.6.2(c)], and high acceleration and vibration of the gear drive become inevitable. Generation of the pinion by a plunging disk enables avoidance of this defect. Figure 17.6.1 shows the generating disk and the pinion in the 3D-space. The surface of the disk is a surface of revolution and is conjugated to the profile-crowned surface of the pinion. The profile-crowned surface \u03c3 of the pinion is a helicoid and is determined as the envelope to the parabolic rack-cutter (see Section 17.4). It is assumed that during the process of generation of the pinion, the pinion performs a screw motion about its axis and is plunged with respect to the generating disk that is held at rest", + " It is assumed that the two components of pinion screw motion and the plunging motion are provided to the pinion. However, one or two of these three components of motions may be provided to the generating disk but not to the pinion. Cambridge Books Online \u00a9 Cambridge University Press, 2009available at https://www.cambridge.org/core/t rms. https://doi.org/10.1017/CBO9780511547126.019 Downloaded from https://www.cambridge.org/core. University of Warwick, on 18 Sep 2018 at 01:19:11, subject to the Cambridge Core terms of use, P1: JXR CB672-17 CB672/Litvin CB672/Litvin-v2.cls February 27, 2004 0:58 Figure 17.6.1: Generation of pinion by disk. Details of the developed approach are as follows: (i) The profile-crowned surface \u03c3 of the pinion is considered as given. (ii) A disk-shaped tool D that is conjugated to \u03c3 is determined [Fig. 17.6.2(a)] (see in addition Chapter 24). The axes of the disk and pinion tooth surface \u03c3 are crossed and the crossing angle \u03b3Dp is equal to the lead angle on the pinion pitch cylinder [Fig. 17.6.2(b)]. The nominal center distance EDp [Fig. 17.6.2(a)] is defined as EDp = rd 1 + \u03c1D (17.6.1) where rd 1 is the dedendum radius of the pinion and \u03c1D is the generating disk radius. (iii) Determination of disk surface D is based on the following procedure (see Chapter 24): Step 1: Disk surface D is a surface of revolution. Therefore, there is such a line L\u03c3 D [Fig. 17.6.2(c)] of tangency of \u03c3 and D that the common normal to \u03c3 and D at each point of L\u03c3 D passes through the axis of rotation of the disk (see Chapter 24). Figure 17.6.2(c) shows line L\u03c3 D obtained on surface D. Rotation of L\u03c3 D about the axis of D enables representation of surface D as the family of lines L\u03c3 D. Step 2: It is obvious that screw motion of disk D about the axis of pinion tooth surface \u03c3 provides a surface that coincides with \u03c3 [Fig. 17.6.2(d)]. Cambridge Books Online \u00a9 Cambridge University Press, 2009available at https://www.cambridge.org/core/t rms. https://doi.org/10.1017/CBO9780511547126.019 Downloaded from https://www.cambridge.org/core. University of Warwick, on 18 Sep 2018 at 01:19:11, subject to the Cambridge Core terms of use, P1: JXR CB672-17 CB672/Litvin CB672/Litvin-v2.cls February 27, 2004 0:58 (iv) Our goal is to obtain a double-crowned surface 1 of the pinion, and this is accomplished by providing a combination of screw and plunging motions of the disk and the pinion. The generation of a double-crowned pinion tooth surface is illustrated in Fig. 17.6.3 and is accomplished as follows: (1) Figures 17.6.3(a) and 17.6.3(b) show two positions of the generated double- crowned pinion with respect to the disk. One of the two positions with center distance E (0) Dp is the initial one; the other with EDp(\u03c81) is the current position. The shortest distance E (0) Dp is defined by Eq. (17.6.1). (2) Coordinate system SD is rigidly connected to the generating disk [Fig. 17.6.3(c)] and is considered fixed. (3) Coordinate system S1 of the pinion performs a screw motion and is plunged with respect to the disk. Auxiliary systems Sh and Sq are used for better illustration of these motions in Fig. 17.6.3(c). Such motions are described as Cambridge Books Online \u00a9 Cambridge University Press, 2009available at https://www.cambridge.org/core/t rms. https://doi.org/10.1017/CBO9780511547126.019 Downloaded from https://www.cambridge.org/core. University of Warwick, on 18 Sep 2018 at 01:19:11, subject to the Cambridge Core terms of use, P1: JXR CB672-17 CB672/Litvin CB672/Litvin-v2.cls February 27, 2004 0:58 follows. Screw motion is accomplished by two components: (a) translational displacement l p that is collinear to the axis of the pinion, and (b) rotational motion \u03c81 about the axis of the pinion [Figs. 17.6.3(b) and 17.6.3(c)]. The magnitudes l p and \u03c81 are related through the screw parameter p of the pinion as l p = p\u03c81. (17.6.2) Plunging motion is accomplished by a translational displacement a pl l 2 p along the shortest distance direction [Fig. 17.6.3(c)]. Such motion allows us to define the shortest distance EDp(\u03c81) [Figs. 17.6.3(b) and 17.6.3(c)] as a parabolic function EDp(\u03c81) = E (0) Dp \u2212 a pl l 2 p. (17.6.3) The translational displacements l p and a pl l 2 p are represented as displacement of system Sq with respect to system Sh. The same translational displacements are performed by system S1 which performs rotation of angle \u03c81 with respect to system Sq . (4) The pinion tooth surface 1 is determined as the envelope to the family of disk surfaces D generated in the relative motion between the disk and the pinion" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002905_cl.2008.230-Figure9-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002905_cl.2008.230-Figure9-1.png", + "caption": "Figure 9. A minigrid-attached film-coated GOx-immobilized p-benzoquinone-mixed carbon paste electrode.36", + "texts": [ + " Plots of magnitude of current decrease as a function of glucose gave a calibration line for quantitative analysis of glucose concentration. In this case, since the detection was conducted by electrochemical reduction, influence by the interference substances was avoidable in the similar manner as that for the reductive detection of glucose using GOx and HP, as shown in Figure 6.35 Electrochemical Techniques Aggressive exclusion of the interference substances by electrochemical means has been proposed by fabricating the specifically designed sensor chip, as shown in Figure 9.36 The sensing electrode was composed of carbon paste containing p-benzoquinone and GOx layer, on which a dialysis membrane was covered. Then, a gold grid electrode was put on the dialysis membrane and was fixed by a nylon sheet. Glucose sensing was conducted by polarizing the sensing electrode at +0.2V vs. SCE. When potential of +0.5V vs. SCE was applied to the gold grid electrode, influence by ascorbic acid was excluded by its oxidation before reaching the sensing electrode. This way was confirmed to be effective for the ascorbic acid concentration up to 180mg dm 3" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000698_ie030855v-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000698_ie030855v-Figure2-1.png", + "caption": "Figure 2. Definition of the flow geometry and coordinate system for simple shear flow. (a) The lower plate is at rest, and the upper plate moves in the x direction with a constant velocity V. H is the gap separation. (b) Cartesian coordinate system with x the flow direction, y the velocity gradient direction, and z the vorticity axis. The director n is defined by the tilt angle \u03b8 and the twist angle \u03c6.", + "texts": [ + " The equations are solved using Galerkin finite elements for spatial discretization and a fourth-order Runge-Kutta adaptive method for time integration. Convergence and mesh independence were established in all cases using standard methods. Spatial discretization was judiciously selected by taking into account the length scale of our model. The selected adaptive time integration scheme is able to efficiently take into account the stiffness that rises as a result of the disparity between time scales: \u03c4i , \u03c4e. Figure 2a shows the rectangular coordinate system for simple shear flow used in this paper; in this frame x is the flow direction, y the velocity gradient direction, and z the vorticity axis. The lower plate is at rest, the upper plate moves in the x direction with a constant velocity V, and H is the gap separation. The director n is defined by the tilt angle \u03b8 and the twist angle \u03c6 (see Figure 2b). The Dirichlet boundary conditions for Q are describing fixed director orientation along the vorticity axis, a uniaxial state with the scalar-order parameter equal to its equilibrium value. The initial state is assumed to be uniaxial and at equilibrum. The orientation of the director at t ) 0 is assumed to be parallel to ns, with thermal fluctuations introduced by infinitesimal Gaussian noise. The simulation parameters are 2.75 < U < 5 and \u00e2 ) 1.2, and the selected ranges for the dimensionless parameters are 103 < R < 5 \u00d7 106, 0 < Er < 5 \u00d7 106, and 0 < De < 7" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003622_s0026-0657(09)70201-2-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003622_s0026-0657(09)70201-2-Figure2-1.png", + "caption": "Figure 2. Electric cover. Left: Conventionally drilled cooling channels. Middle: Optimised conformal cooling. Right: Simulated temperature distribution within tool. Image courtesy of PEP, Legrand.", + "texts": [], + "surrounding_texts": [ + "8 MPR September 2009 0026-0657/09 \u00a92009 Elsevier Ltd. All rights reserved.\nin business\nChannelling quality for moulded parts using fast manufacturing Direct metal laser-sintering (DMLS) technology may sound like something out of science fiction. But as its manufacturers and users report, it\u2019s gaining ground among customers as a way to produce complex parts quickly and accurately. Some developments have clear economic advantages, says Liz Nickels...\nDirect metal laser sintering (DMLS) produces solid metal parts by locally melting and solidifying metal powder with a focused laser beam, layer by layer. A 3D computer-aided design model is \u201csliced\u201d into layers, and the lasersintering technology then builds the geometry. The technology\u2019s Munichbased manufacturer, EOS, says that it can create extremely complex metal parts in a relatively short cycle time and is particularly suitable for industries that no longer need to produce a large volume of identical parts.\nPotential applications cited by EOS include tooling, aerospace and automotive industries, designer objects, con-\nsumer goods such as toys, and robotics. New applications are being found in the medical field \u2013 dental prostheses, implants and devices.\nAccording to the EOS website: \u201cLaser-sintering enables a rethinking in product development and production. It is a departure from tool-based, inflexible technologies in favour of generative, flexible methods.\u201d Claudia Jordan, an EOS spokesperson, said: \u201cThe most groundbreaking aspect of the technology, in our opinion, is freedom of design. This makes it possible to produce really attractive products, such as customised implants and lightweight structures for aerospace and professional cycling.\u201d\nBut is the technology \u2013 also called e-manufacturing by the company \u2013 an innovation too far for customers in the current economic climate, where innovation can take second place to the stability of \u201ctraditional\u201d technology?\nOne user of the technology is LBC Laserbearbeitungscenter GmbH, a manufacturer of inserts, parts and prototypes for diecasting and injection moulding. Currently, 20% of its work involves producing prototypes for aircraft, the medical industry or for the automotive sector. It was established in October 2002 and only two years later started using EOS\u2019s DMLS technology.\nIn the injection moulding tooling business, there is hardnosed busi-", + "September 2009 MPR 9metal-powder.net\nness reasoning behind the deployment of innovative technology. The DMLS system allows LBC to build moulds that include internal conformal cooling channels. These are carefully calculated and designed to reduce the time taken to cool the mould in each moulding cycle, making for faster throughput and improving productivity and, therefore, profitability.\n\u201cWe\u2019re very pleased with the technology. It is excellent for mass production moulds,\u201d said Ralph Mayer, managing director of LBC. \u201cWith conformal cooling we can achieve cycle time reduction up to 60%, although the average time reduction is 30% to 40%.\u201d\nConformal cooling is, in fact, a fairly recent commercial reality. \u201cIn theory, EOS has been dealing with the concept of conformal cooling for over 10 years,\u201d said Jordan. \u201cBut it was only when we started to implement MS1 (a maraging steel in fine powder form) at the end of 2006 that EOS could actively offer conformal cooling.\u201d\nDMLS allows for almost any shape in heating/cooling channels and thus can already improve the effectiveness of\ncooling. However, by using conformal cooling with DMLS, routing options for cooling channels are almost infinite.\nEOS goes so far as to claim that certain geometries of products can only be manufactured at required quality standards with conformal cooling. \u201cIt is suitable for making very complicated parts to a very high quality in a much faster time than traditional moulding,\u201d agrees Mayer. \u201cFor example, a gear wheel requires special milling tools to be created, and can take up to four months using existing technology. Using DMLS, it is possible to make a prototype in just one week.\u201d\nBenefits of conformal cooling\nAccording to an EOS white paper, conformal cooling works by creating a suitable cooling channel at a well defined distance to the cavity, which is impossible using a conventional drilled cooling mechanism (Figure 1). Cooling channel cross sections can take almost any shape. Turbulence of the coolant (the desired high Reynolds number)\nwithin the system can thus be controlled by actively choosing different cross sections and by switching between different cross sections. As a consequence, turbulence inside the coolant stream is generated close to the mould cavity along the whole path of the channels.\nChanging the cross sections or forking the cooling channel can be done without splitting up the form. This allows for additional heat/cooling advantages in areas that cannot be reached by conventional methods.\nConformal cooling can also improve quality due to better control of the injection moulding process. Warping and sink marks are minimised by evenly cooling out the plastic melt, thus minimising internal stress. Scrap rates are reduced or eliminated. Avoiding internal stresses helps to produce better parts with the same amount of required material. In fact, certain geometries are only possible to manufacture at required quality standards with conformal cooling, EOS says.\nCombined systems with separated cooling and heating channels are also possible, and the technology can also", + "10 MPR September 2009 metal-powder.net\nperform a split between main systems (for the control of the global temperature) and specific systems (for the handling of close to cavity critical temperatures), opening up the potential for future applications. Heating/cooling at critical parts inside the tool, which often cannot be reached by conventional methods, becomes feasible (eg long and lean cores, areas around hot-runners or small sliders). Using special copper heat conductors or other complex measures becomes obsolete. If necessary, it is also possible to undercool mould cavities, thus reaching optimal cycle times by minimising cool down times in tooling cavities.\nAn evened out temperature level can help to improve tool life time. This becomes relevant especially in die casting tools that are exposed to extreme temperature variations.\nIn the white paper, EOS also sets out the drawbacks of conventional cooling. The distance from cavity to cooling channel differs, as only straight line drilling channels are possible and as a consequence heat dissipation cannot take place uniformly in the material.\nThis can result in uneven temperature levels on the cavity surface, uneven cooling-down processes resulting in internal stresses and thus negative impact on part quality (warpage). The drilling procedure is not without risk: in the case of deep drilling there is always a danger of hitting ejector holes (wandering drill), or the drill can even break. As a consequence, the whole mould insert could become unusable.\nIn conventional cooling, optimising the cooling channels helps condition the tooling temperature, enabling uniformity. This temperature level can be influenced to achieve on the one hand a lower temperature for quicker cooling, or higher temperature for better product surface quality on the other.\nConventional cooling channels are drilled into a tool. This limits design to straight lines, easily accessible by a drill. Tooling cavities can pose limits to position and routing of conventional cooling channels. With DMLS, however, the cooling channels can be positioned freely. The cross sections can be optimised to mould temperature control requirements.\nFuture developments Some of the results already obtained by using DMLS to create injection moulds and mould inserts with conformal cooling include a 20% increase in mould productivity and a 50-hour toolmaking time for a blow-mould (Es-Tec), reduced cycle times from 15 to 9 seconds, enabling a 75% increase in productivity on a four-bottle blow mould with DMLS inserts (SIG Blowtec) and reduced cycle time by two-thirds using a DMLS designed core effectively cooling down a critical hot spot (LaserBearbeitungsCenter).\nEOS says that DMLS can make electric discharge machining (EDM) and milling obsolete in many cases, especially with part geometries where slides, inserts or other tool components with complex characteristics are required. But one issue with the technology is that building rate is still a problem, according to Mayer. \u201cCurrently it is much faster to build components in large numbers using traditional moulding technology. At the moment, using DMLS, we can only produce 8cm3 per hour. However it seems likely that this will improve in the future.\u201d" + ] + }, + { + "image_filename": "designv11_69_0001563_095440605x8397-Figure4-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001563_095440605x8397-Figure4-1.png", + "caption": "Fig. 4 Coordinate systems for meshing simulations with assembly errors", + "texts": [ + " The gears are considered to be rigid bodies, and no deformation occurred under the meshing process. 2. The profile error is very small and can be neglected. 3. The effects due to temperature variation and dynamic loading have also been ignored. 4. The assembly errors can be measured. C11604 # IMechE 2005 Proc. IMechE Vol. 219 Part C: J. Mechanical Engineering Science at RMIT UNIVERSITY on July 12, 2015pic.sagepub.comDownloaded from When assembly errors exist, coordinate systems are set up as shown in Fig. 4. In Fig. 4a, the fixed coordinate system Sm(XmYmZm) is set at the assembly location of the gear, with its rotational axis as the Zm axis. Coordinate system Sg(XgYgZg) is attached to the gear, and ug is the angular displacement of the gear. Three axes of the fixed coordinate system Sf(XfYfZf) parallel those of Sm, and the distance between the two origins, Of and Om, is equal to the ideal centre distance of the gear pair. Coordinate system Se(XeYeZe) is chosen at the erroneous pinion assembly position, with its Ze axis being the rotational axis of the pinion. Coordinate system Sp(XpYpZp) is attached to the pinion, and up is the angular displacement of the pinion. The linear and angular errors of axial misalignment are defined as shown in Fig. 4b, where three components of d are linear errors, and ux and uy are angular errors. Let matrix Lij, which is the rotation matrix for the rotation transformation from coordinate system Sj(XjYjZj) to Si(XiYiZi), be obtained by deleting the fourth row and the fourth column of matrix Mij. Based on gearing theory, both meshing surfaces of the gear and pinion must have a common contact point and a common unit normal vector at every instant. When assembly errors occur, their relations can be represented as MfgR (g) g \u00bc MfpR (p) p \u00bc MfnMneMepR (p) p (14) Lfgn (g) g \u00bc Lfpn (p) p \u00bc LfnLneLepn (p) p (15) Proc" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002741_sice.2008.4654660-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002741_sice.2008.4654660-Figure1-1.png", + "caption": "Fig. 1 Engagement geometry in iL -frame", + "texts": [ + " And we will derive the equations in iL -frame which is defined at every filtering instant of t i as follows: (L1) The origin of iL -frame locates at the target position at t i and it moves with the velocity equivalent to the target velocity. (L2) The X-axis of iL -frame directs along the line-of-sight(LOS) vector at t i . And it holds PR0001/08/0000-0247 \u00a5400 \u00a9 2008 SICE - 248 - until the homing guidance completes. Note that the iL -frame is inertial since it moves without acceleration and rotation. Since iL -frame moves together with the target, the trajectory formed in iL -frame will be relative movement of the missile to the target. Fig. 1 shows the engagement situation of the missile and target in iL -frame at t k ( k i ). In the figure, relV , , and are relative velocity, heading error angle, and LOS angle at the current time t k , respectively. Let\u2019s assume that the missile flies around the collision triangle. The approximated relative motion of the missile to the target will be derived in the similar way of [2]. In this case, the LOS angle can be approximated by y x (1) where c gox V t (2) and, x , y , and cV are X and Y elements of the relative position vector, and closing velocity, respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002301_0020-7403(67)90046-x-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002301_0020-7403(67)90046-x-Figure1-1.png", + "caption": "FIG. 1. T he n e u t r a l axis is a n a rea bisector .", + "texts": [ + " His so lu t i on was based u p o n a n impe r f ec t a c c o u n t of t he g e o m e t r y of d e f o r m a t i o n , b u t m a y well g ive good a p p r o x i m a t i o n s in m a n y cases. T H E C E N T R O I D A L L O C U S No method has yet been found for calculating the value of the fully plastic moment about a given axis directly, nor for finding the direction of the neutral axis corresponding to a given moment axis. Given the direction of the neutral axis, however, it is a simple task to find the corresponding fully plastic moment. Fig. 1 shows a general cross-section of a beam with an arbitrarily chosen direction of neutral axis. In the fully plastic state, assuming ideal plasticity, the whole area A T on 77 78 E . H . BBowN the tensi le side is s t ressed to t h e yie ld s t ress a~, a n d t he r e m a i n i n g a r ea Ac is s t ressed to - a v. F o r axia l equ i l ib r ium, i f t h e appl ied forces reduce to a b e n d i n g couple only, a ~ A T - a ~ A c = 0 o r A AT = Ac = ~ (1) where A deno te s t he comple t e a rea of t he cross-sect ion", + " E q u a t i o n (1) comple t e ly d e t e r m i n e s t he pos i t ion of t he n e u t r a l axis in a g iven d i rec t ion , since t h e r e c a n n o t be two para l le l a rea b isec tors of a c o n t i n u o u s cross-sect ional shape . The possible n e u t r a l axes of a g iven cross-sect ion are t h u s t he f ami ly of b i sec tors ; in genera l t h e y do n o t all m e e t in a po in t . The r e s u l t a n t t ens ion is a force a~ A/2 a c t i n g t h r o u g h Gr , t he cen t ro id of AT, a n d the equa l r e s u l t a n t compress ion s imi lar ly ac t s t h r o u g h Go, Fig. 1. As t h e two areas a re equal , t h e cen t ro id G of t he whole cross-sect ion m u s t b i sec t t h e l ine G r - G e. I f G-G~, = G - G v = r t h e m a g n i t u d e of t h e ful ly p las t i c m o m e n t is g iven b y M~ = a v A r (2) a n d t he d i rec t ion of i ts axis is p e r p e n d i c u l a r to G r - G c . ~ L LOcus FzG. 2. T he cen t ro ida l locus is a lways skew-symmet r i ca l . A p lo t of t h e cu rve t r a c e d ou t b y G r a n d Gv as t he inc l ina t ion of t h e n e u t r a l axis is r o t a t e d t h r o u g h 180 \u00b0 will be t e r m e d \" t h e cen t ro ida l locus\" " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000396_2001-01-1445-Figure12-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000396_2001-01-1445-Figure12-1.png", + "caption": "Fig. 12 Setting of various points for the front suspension system in SAMF", + "texts": [ + " The stiffness sensitivity, which means the FRF change ratio per additional unit stiffness, can be obtained as follows. ab abunitKab k G GG S \u2212 = , \u02c6 (13) Where unitKabG , \u02c6 is calculated by substituting the unit stiffness for K (f) of the second term in Eq. (12). The FRF data used in SAMF is acquired experimentally by an impulse excitation test for the suspension system fixed in the suspension tester. Three translational degrees of freedom are considered for the input force and the responses. The excitation force is provided at specific points as shown in Fig. 12, to constitute full 3x3 matries of FRFs. Each point is equipped with a rigid and small excitation adapter and is indirectly excited due to structural restriction. The response at each point is measured by a tri-axial accelerometer. As the evaluation point and direction of the response, the lateral direction at the center of the front cross member is selected, i.e., the causal vibration of road noise mentioned in previous section. Moreover, the same location, the center of the cross member, is selected as the evaluation point for the riding comfort characteristics, considering the evaluation of the input force to the body" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003424_s12206-009-1173-y-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003424_s12206-009-1173-y-Figure2-1.png", + "caption": "Fig. 2. Positions of the intake and exhaust ports around the top and second rings.", + "texts": [ + " We used a twostroke air-cooled single-cylinder gasoline engine with a bore of 62 mm and a stroke of 58 mm. In this engine, the cylinder has an intake port on the thrust side, an exhaust port on the anti-thrust side, and scavenging ports on both the front and rear sides. In this piston, we installed two rings: a barrel-faced half keystone top ring on top and a taper\u2013faced rectangular second ring below it. Both rings had a width of 2.0 mm, a thickness of 2.8 mm, and a tension of 11N. The ring grooves in our piston use stop pins to prevent rotation of the rings under engine operation. Fig. 2 shows the positions of the intake and exhaust ports around the top and second rings. In order to measure the strain on the bottom sides of the rings as they slid over the ports, we used a strain gauge with a grid width of 0.84 mm and a grid length of 2.0 mm. We attached the strain gauge on the bottom sides of our rings over the center of port width. To protect the lead wires of our strain gauge from breaking, we set terminals next to the strain gauge, then connected our lead wires and heavy-duty signal wires via terminals, as shown in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001784_11539902_140-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001784_11539902_140-Figure2-1.png", + "caption": "Fig. 2. Obstacle classification and a MISO neural-fuzzy system", + "texts": [ + " 9, and left multiplying ( J\u2020 E )T \u00b7 S\u0304T , yields M\u0304 \u00b7 x\u0308E + V\u0304 \u00b7 x\u0307E + G\u0304 = \u03c4\u0304 (10) Where M\u0304 = ( J\u2020 E )T S\u0304T MS\u0304J\u2020 E , V\u0304 = ( J\u2020 E )T S\u0304T ( M \u02d9\u0304SJ\u2020 E \u2212 MS\u0304J\u2020J\u0307 + V S\u0304J\u2020 E ) , G\u0304 = ( J\u2020 E )T S\u0304T G, \u03c4\u0304 = ( J\u2020 E )T S\u0304T B ( \u03c4 + JT Fext ) ; and ( J\u2020 E )T S\u0304T C\u03bb = 0 is eliminated. Remark 1. The following properties hold for Eq. 10: 1) For any r \u2208 n+2, rT \u00b7 M\u0304 \u00b7 r \u2265 0; 2) For any r \u2208 n+2, rT \u00b7 ( \u02d9\u0304M \u2212 2V\u0304 ) \u00b7 r = 0; 3) If J is full rank, JE = ( J\u2020 E )\u22121 = [ JT \u2223 \u2223 J\u2135 ]T ; 4) If J is full rank, M\u0304, V\u0304 , G\u0304 \u2208 \u221e. Here \u221e = {x (t) \u2208 n : \u2016x\u2016\u221e < \u221e}. According to whether on the desired end-effector trajectory or not, obstacles can be divided into two kinds: the task-consistent one and the task-inconsistent one, see Fig. 2(a). The task-consistent obstacles can be avoided by on-line adjusting self-motions. However, the task-inconsistent obstacles can not be avoided without affecting end-effector executed tasks. One solution to avoid task-inconsistent obstacles is to regenerate the desired end-effector task, which belongs to the high-level decision making problem and is beyond the discussion of this paper. In this paper, only task-consistent obstacles are concerned and redundancy of the robot is supposed to be high enough to avoid obstacles just by adjusting selfmotions", + " Assume the system is far away from singularity and physical limits, then the self-motions can be used specially for obstacle avoidance. If a point on the robot gets too close to an obstacle (\u2016dij\u2016 < dc), this point can be called a critical point, and dc is called the cut-off distance. The artificial potential function for the ith critical point and the jth obstacle can be defined by \u03c6ij(q) = { 1 2 \u00b7 k\u03c6 \u00b7 ( 1 \u2016dij\u2016 \u2212 1 dc )2 , \u2016dij\u2016 < dc 0, \u2016dij\u2016 \u2265 dc (11) Where k\u03c6 > 0 is a constant coefficient, dij = xci \u2212 xoj is the nearest distance between the ith critical point and the jth obstacle, as shown in Fig. 2(a). Here, xci = [pcix pciy pciz]T and xoj = [pojx pojy pojz]T are position vectors for the ith critical point and the jth obstacle with respect to OBXBYBZB. To avoid obstacles in a realtime manner, the self-motions can be planned to optimize the following function: \u03a6 (q) = No\u2211 j=1 Nc\u2211 i=1 \u03c6ij (q) (12) Where No and Nc are the numbers of obstacles and critical points respectively. Then q\u0307sd = \u2212\u2202\u03a6 (q) \u2202q = \u2212 No\u2211 j=1 Nc\u2211 i=1 \u2202\u03c6ij (q) \u2202q (13) Where \u2202\u03c6ij(q) \u2202q can be derived from Eq. 12. \u2202\u03c6ij (q) \u2202q = \u23a7 \u23a8 \u23a9 \u2212 [ k\u03c6 \u00b7 ( 1 \u2016dij\u2016 \u2212 1 dc ) \u00b7 dT ij \u2016dij\u20163 \u00b7 ( \u2202xci \u2202qT \u2212 \u2202xoj \u2202qT )]T , \u2016dij\u2016 < dc 0, \u2016dij\u2016 \u2265 dc (14) Then, x\u2135d, x\u0307\u2135d and x\u0308\u2135d can be determined. Theorem 1. (Universal Approximation Theorem [12]) The multiple inputs single output (MISO) fuzzy logic system (FLS) with center average defuzzifier, product inference rule and singleton fuzzifier, and Gaussian membership function can uniformly approximate any nonlinear functions over a compact set U \u2208 n to any degree of accuracy. If the FLS described above is realized by a neural network (NN), a neural fuzzy system (NFS) can be obtained as shown in Fig. 2(b). Output of this NFS is given by fNFS = Nr\u2211 j=1 { wj \u00b7 Ni\u220f i=1 { exp [ \u2212 ( xi\u2212 ji \u03c3ji )2 ]}} Nr\u2211 j=1 { Ni\u220f i=1 exp [ \u2212 ( xi\u2212 ji \u03c3ji )2 ]} (15) Where xi is the ith input variable, wj denotes the jth centroids for the output fuzzy sets, i = 1, 2, \u00b7 \u00b7 \u00b7 , Ni, j = 1, 2, \u00b7 \u00b7 \u00b7 , Nr, here Ni and Nr represent the number of input variables and rules respectively. ji and \u03c3ji are the mean and standard derivation of the Gaussian membership functions accordingly. Let xd, x\u0307d and x\u0308d be desired task-space position, velocity and acceleration" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002964_wsc.2007.4419836-Figure3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002964_wsc.2007.4419836-Figure3-1.png", + "caption": "Figure 3: The DoFs of a Crane", + "texts": [ + " A loaded crane has a maximum of eight degrees of freedom (DoFs) (Reddy and Varghese 2002), and path planning for manipulators having more than four DoFs is considered to be complex (Hwang and Ahuja 1992). As mentioned by Reddy and Varghese (2002), there can be multiple solutions to configure the DoFs of the manipulator for a particular location of the endeffector (i.e., the hook); therefore, simplifying the representation and avoiding the inverse kinematics should be considered. The scope of the present work is limited to four DoFs, as shown in Figure 3. The engineering constraints of cranes are mainly from the working range and the load charts. The working range shows the minimum and maximum boom angle according to the length of the boom and the counterweight. Load charts give the lifting capacity based on the boom length, boom angel to the ground and the counterweight. Three major criteria should be taken into account: Lift path clearances, capacity during lift, and ground support during lift. Rules should be developed to represent these constraints which are stored in a database" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001550_iros.2005.1545550-Figure5-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001550_iros.2005.1545550-Figure5-1.png", + "caption": "Fig. 5. Driving mechanism of forceps;(a) Rotation. (b) Translation.", + "texts": [ + " We adopted an ultrasonic motor with rotary encoder (custom order, Fukoku, Japan) to drive the outer case of friction wheel because it has advantages in compact size, light weight, high holding torque, and suitable for hollow-shaft one. b) Rotation: For the rotation of the forceps around its shaft, we rotate both friction wheels in the same direction. In this case, rollers holding the shaft do not rotate, and do not make spiral motion. Thus, the shaft held by rollers rotates at the same speed as the FWM (Fig.5(a)). c) Translation: In the case of translation, we rotate each FWM in the opposite direction (Fig.5(b)). Each rotational motion cancels each other, and only translation remains. Thus translation along its shaft is realized. In the former prototype, because of the low machining accuracy, each center of friction wheels was not located in the same straight line, thus, FWM did not move smoothly. In this study, we modified the housing where friction wheels were mounted, and arrayed each center of friction wheels in the same line. Because the rotational speed of ultrasonic motor was not stable depending on the load, we implemented optical rotary encoder to get the rotational speed information" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003740_10402000903491283-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003740_10402000903491283-Figure2-1.png", + "caption": "Fig. 2\u2014Redesigned slide rod.", + "texts": [ + " Compared with the traditional tripod joint, its input shaft and tripod link up through three slide rods instead of three spherical or cylindrical rollers. However, the premature failures induced by the lubricating insufficiency of the main mating surfaces restrain the popularization and utilization thereof. Especially, severe failures of the mating surfaces between the sleeves and slide rods were frequently found in that they are the main elements transmitting torque. In order to improve the lubricating performance between them, therefore, the slide rod was redesigned into a novel one with several annular bumps around a column (Fig. 2) by authors. This is obviously an elastohydrodynamical lubrication (EHL) problem under reciprocating motion with only relative motion considered between the sleeves and slide rods. This problem has been studied by several researchers. Petrousevitch, et al. (15) accomplished the theoretical and experimental studies on the film thickness and shape for the point contact of a reversible ball against a plane. Hooke (16) observed the minimum film thickness under EHL line contacts during a reversal of entrainment and obtained the film thickness formulae on the basis of dimensionless groups" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001785_piee.1971.0221-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001785_piee.1971.0221-Figure1-1.png", + "caption": "Fig. 1A Conductors on level ground", + "texts": [ + "1 Preliminary considerations Carson6 gives expressions for the mutual impedance between two parallel circuits with earth return, and for the self impedance of a circuit witn earth return, in terms of the corresponding impedances with perfectly conducting earth, plus functions of two dimensionless parameters r and 6; r pertains to the attenuation of electromagnetic waves propagated through the earth, and 6 is an angle defined by the geometry of the system. Equivalent expressions are given by Pollaczek.7 Referring to Fig. 1A, which represents a cross-section through the parallel conductors G and N, assumed to be horizontal, and the surface of the ground assumed to be hori- zontal and plane, the X axis of rectangular co-ordinates is the straight line on the Earth's surface perpendicular to G and N, and the Y axis is the straight line through N perpendicular to the X axis. N is considered as fixed in position with coordinates (0, yn) metres, while G is a variable, with co-ordinates (xng, yg) metres. 1227 In general, the surface of the ground is not a smooth plane, and by convention hypothetical supports are assumed to be in the plane of the cross-section. The X axis is then taken as the straight line through the feet of the supports, and the surface of the ground is taken as the plane parallel to the conductors and containing the X axis. The quantities r, X and Y may be considered as 'electrical lengths', as distinguished from the corresponding physical lengths, Dng, xng and yg + yn. Conductors on sloping ground Where the feet of the supports are on different levels, as in Fig. 1B (exaggerated for clarity), the co-ordinates of N and G on the Y axis, namely yn, yg, are not the same as the respective heights hn, hg, and xng is not the same as the distance between the feet of the supports; however, in practice, the differences are usually trivial. If a conductor lies below the X axis, as in the case of a buried cable, the corresponding Y ordinate has a negative sign. (Note that the actual values for ym yg, hn, hg, used in calculations for aerial lines are less than the dimensions at the supports, as allowance has to be made for the sag of the conductors" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003676_978-1-4020-9402-6_15-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003676_978-1-4020-9402-6_15-Figure2-1.png", + "caption": "Fig. 2 Schematic representation of Kagome-structure, taken from [2]. The face-sheet is shown in blue, the tetrahedral core in green and the Kagome back-plane in red", + "texts": [ + " For all analyses conducted within the frame of the present work, the value of 17% was adopted for the relative density. Using L = 1mm, from Eqs. (1)\u2013(3) we obtain tT = 0.05mm, tH = 0.15mm and tK = 0.1mm, respectively. It is desired to assess the effect of notch\u2019s presence in the tensile behavior of the lattices. The Kagome structure is a very attractive structure for application as a shapemorphing material because of its unique combination of low actuation resistance and high passive strength. Hence, it can be easily actuated into relatively intricate surface shapes. As shown in the schematic of Fig. 2, it consists of a solid face-sheet with a Kagome, active back-plane and a tetrahedral core. Shape change of the facesheet is achieved by replacing various struts in the back-plane with linear actuators. During service, the structure is loaded simultaneously by passive loads and actuation loads. The length of the struts in both the back-plane and the core is 51 mm while their diameters are 3.0 and 1.4 mm, respectively. All members of the structure are assumed to be made of steel with a Young\u2019s modulus of 200 GPa, Poisson ratio of 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000560_robot.1988.12083-Figure6-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000560_robot.1988.12083-Figure6-1.png", + "caption": "Figure 6: A six-d.0.f. Cartesian type pantograph manipulator with a pair of differential mechanism.", + "texts": [ + " Since these pitching and yawing motions affect the orientation of the wrist, the inverse position analysis of the last three axis is more complicated. The inverse velocity and acceleration analyses become very complicated because the first and second derivatives of the angle a2 have very complicated forms (refer to Equation (43) in Appendix). Thus, the advantage of simple kinematics diminishes. Hirose and Umetani' introduced a set of differential mechanism which can eliminate the unwanted pitching and yawing motions of the wrist. A modified version of this set of differential mechanism is shown in Figure 6. During any motion of the first three axes, the bottom bevel gear at joint U does not rotate and, the unwanted yawing motion of the wrist is eliminated via the differential mechanism. The unwanted pitching motion is eliminated by through the pulley and chain system. Hence, the kinematics of the last three d.0.f. is fully decoupled from the first 417 # X,,X. F i g u r e 7 : A k i n e m a t i c c y l i n d r i c a l model o f a s i x - d . 0 . f . t y p e pantograph m a n i p u l a t o r a t a genera l p o s i t i o n " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001231_sice.2006.314734-Figure4-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001231_sice.2006.314734-Figure4-1.png", + "caption": "Fig. 4. (a) Complete bearing (b) defective inner raceway (c) defective ball (d) defective outer raceway.", + "texts": [ + " A laboratory set is prepared to examine theoretical results, which consists of a vibration sensor and its driver circuit, amplifying and filter circuits, a data acquisition card and a digital computer. To obtain the vibration of defective bearings an apparatus is made consisting of a shaft which two bearings are assembled on its ends and two housings to retain the bearings, accompanied by two induction machines working as motor and generator. By proper selection of pulleys and belts the shaft can rotate in different velocities. All of these parts are mounted on a frame (Fig. 3). The applied bearing can be easily disassembled. The complete bearing and its defective parts are shown in Fig. 4. Bearing has 28 balls in two rows, ball diameter is 7.9mm, inner raceway diameter is 35mm, and outer raceway diameter is 57mm. The vibration sensor is an accelerometer, with a bandwidth of more than 10 kHz, the selected data acquisition card, is a 16-bit card with a maximum sampling rate of 250ks/s for all channels. The circuits consist of a current source to drive the vibration sensor and a low-pass 6th order Butterworth filter with a cut-off frequency of 1 kHz. This circuit amplifies the signal too" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003622_s0026-0657(09)70201-2-Figure5-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003622_s0026-0657(09)70201-2-Figure5-1.png", + "caption": "Figure 5. Optimal design of a three dimensional channel system.", + "texts": [ + " Case study \u2013 PE bottles In this case study small DirectTool inserts with conformal cooling channels were built and integrated into a conventionally manufactured tool in order to extract the heat from these parts more quickly. The inserts reduced cycle times from 15 down to 9 seconds. This enables a 75 % increase in productivity for a 4 bottle blow mould without sacrificing on quality. The cycle-time and productivity of this type of tool were limited by the time it took to cool down the bottle necks as wall thickness reaches a maximum there. Figure 5 shows three examples using conformal cooling channels. Figure 5a shows a tool for blow-moulding PE bottles, while figure 5b shows a cooling pin for cooling an injection point \u2013 a classical hot spot. Conformal cooling in this case reduces cycle times by two thirds. Figure 5c depicts a core with spiral conformal cooling channels inside the dome. By conventionally milling the lower part of the tool and limiting DMLS to the part with conformal cooling, costs were reduced. A 0.3mm machining allowance was added for finishmachining of the outer surface." + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003326_1464052.1464063-Figure7-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003326_1464052.1464063-Figure7-1.png", + "caption": "Figure 7. Solenoid Driver Circuit.", + "texts": [ + " A second unit now under construction, uses a Golay code and its redundant bits are generated serially, with higher speed logic to partially compensate for the multiple logic cycles. In that unit, with modest access time, it was found more economical to supply a sepa- 15 14 13 12 11 10 9 8 7 6 5 4 3 2 1 MODULO - 2 - SUM rate, faster clock and control counter, all in strumented in higher speed logic, than to gen erate the Golay code in a parallel. Drive Solenoids The drive solenoids are operated in pairs, with their respective windings connected in parallel so that for one given drive polarity the solenoid flux polarities are opposite as shown in Figure 7. This balanced configuration achieves an approximation of a closed magnetic circuit without the need for an actual closure. Although, because of the air gaps, the mutual inductance between the two solenoids of a pair is not large, the superposition of the individual solenoid fields radically reduces the stray flux. Further, although the correlation used for ad dressing is very non-critical, the drive pattern sensitivity of individually driven solenoids would be unacceptable. Thus, to minimize the need for magnetic shielding between the drive and pick-up arrays, to minimize drive pattern sensitivity and to somewhat improve mutual coupling to the data planes, paired solenoids can be fully justified", + " To give the structure mechanical strength, the ferrites are inserted into a phenolic-paper tube and appropriately glued in, and the wind ings are laid into a shallow threaded groove cut on the outside of the tube. The entire as sembly is then appropriately varnished or epoxy coated. Solenoid Drivers The solenoid drivers are designed to supply a 16 volt, half microsecond long pulse to an inductive load of about twenty microhenries, with ample margin. They are also designed to withstand an inductive overshoot equal to the drive pulse, or about 32 volts peak. As was mentioned earlier, each solenoid pair is connected in parallel, as shown in Figure 7, so that the solenoids in each pair always have opposite polarities. A second consideration is that a \"one\" input should drive the pair one way, and a \"zero\" the other. In earlier designs, a transformer with two input drive windings of opposite winding polarity, one from each of two separate drivers, coupled to an output winding that was connected to the solenoid load. When \"one\" was asserted, one switch closed and drove the transformer, and when a zero was asserted, the other switch and wind ing drove the transformer, thereby generating opposite drives on the output winding for the two states" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001198_esda2006-95466-Figure8-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001198_esda2006-95466-Figure8-1.png", + "caption": "Figure 8. VOLUME VARIATION AT THE TOOTH TIP", + "texts": [ + " Because of the clearance between the bush and the case, the isolated tooth space is put in communication with the high or low pressure. So, for the tooth space that is connected with the low pressure chamber there is a flow that comes out of the tooth space, on the contrary, for the tooth space that is connected with the high pressure chamber there is a flow that comes into the tooth space. The volume variation dVi of the tooth space i [5], for an angle rotation d\u03b8 of the gear, is obtained by the difference of the shaded areas in figure 8 multiplied by the thickness b of the gear: 6 Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 02/07/2016 Te dVi d\u03b8 = rextb(hi+1 \u2212hi) (9) So, taking into account all the previous terms and substituting in equations (2), the continuity equation for the generic tooth space i can be obtained. Moreover, in order to calculate the pressure distribution, the model takes into account the pressure variation in the low and high pressure chamber and the trapped volume too. The equation (1) has been applied to all the control volumes, considering for the terms on the left side of equation (1) all the volumetric flows depicted in figure 4 and where the volume variation is geometrically obtained using equation (9)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000210_iros.1996.571043-Figure11-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000210_iros.1996.571043-Figure11-1.png", + "caption": "Figure 11: Mobile manipulator used for experiment", + "texts": [ + " Next, we fix the desired position of the reference point on an extended line from the origin to the contact point with a distance 1, from the contact point, where I, is the distance from the first joint to the arm tip on XOYO plane when the manipulability[l5] is maximum (see Figure 10). Then the desired trajectory of the reference point is obtained from that of the object. 7 Experiment 7.1 Experimental system In this section, experimental result of pushing with a mobile manipulator will be presented to show the effectiveness of the proposed method. The robot used for the experiment is a mobile robot with two independent driving wheels on which a horizontal paralleldrive two-degrees-of-freedom manipulator is mounted. Figure 11(a) shows an overview of the mobile manipulator, and Figure l l ( b ) shows its parameters. Rotational angles of wheels are measured by optical encoders, and rotational velocity, position, and orientation of the mobile robot are calculated from the output of the encoders. These computations are performed on a board computer with 68000 CPU, and the data is sent to the host computer with i4861i487 CPU(66MHz) through DPM(Dual Port Memory). Each wheel is driven by a motor which is controlled by the board computer according to the instruction sent from the host computer through DPM" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002188_ijtc2008-71096-Figure6-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002188_ijtc2008-71096-Figure6-1.png", + "caption": "Fig. 6. Test stand for detailed research on single bearing (Sbts). 1 \u2013 main shaft \u00f8100mm ; 2 \u2013 tested bearing bush; 3 \u2013 baring sleeve; 4- covers with sealings; 5 \u2013 static load leaver; 6 roller bearings; A \u2013 pressure sensor ; B \u2013 eddy current proximity sensors; C - torqmetter;", + "texts": [], + "surrounding_texts": [ + "Downloaded\nProceedings of the STLE/ASME International Joint Tribology Conference IJTC2008\nOctober 20-22, 2008, Miami, Florida USA\nIJTC2008-71096\nMARINE WATER LUBRICATED STERN TUBE BEARINGS \u2013 CALCULATIONS AND MESUREMENTS OF HEAVY LOADED BEARINGS\nWojciech Litwin\nFaculty of Ocean Engineering and Ship Technology Gdansk University of Technology,\nUl. Narutowicza 11/12; 80-952 Gdansk; Poland\nINTRODUCTION Water lubricated bearings are often used to support propeller shaft of modern ships. These environmentally friendly bearings are less expensive than those oil lubricated because of simplicity of the bearing unit and lower material costs [1]. However, there were many serious failures associated with this kind of bearings. In some cases a rapid wear process can result from instability of the hydrodynamic film, poor material properties, or unfavorable working conditions. Theoretically, water lubricated stern tube bearings operate in hydrodynamic regime. Usually, the unit load is low because of high bearing length to shaft diameter ratio (L/D). Sliding speed depends on the engine type, main gear ratio and size of the ship. Generally, this speed can exceed 3 m/s, which should be sufficient for hydrodynamic film to develop. The following three causes can be behind failures of these bearings: design flaws, ship specific problems such as long shaft line on flexible ship like double ended ferry, and harsh sea environment [2].\nOne of the main problems is to define maximum hydrodynamic film capacity. Because of very thin hydrodynamic films and significant bearing bush deformation the EHL model should be used for calculations. It is especially important for flexible polymer materials.\nTEST BEARING\nA polymer bearing for marine application was tested. The\nbush had five open grooves. The bearing geometry, which is typical for propeller shaft bearings, is illustrated in Figure 1.\nshaft diameter diameter clearance bearings length bearings load Bearing bush module of elasticity\n[mm] [mm] [mm] [MPa] [MPa] 1 0,16 2 0.28 3 0.41 4 99,88 0,46 300 0.52 800\nTable. 1. Tested bearing specifications\nFrom: http://proceedings.asmedigitalcollection.asme.org/pdfaccess.ashx?ur\nFig.1. Test bearing\nThe tests were conducted for conditions typical for a\nmarine propeller shaft aft bearing. Shaft rotation speed ranged from 0 to 11 rev/s. The static radial load was higher than in classic bearing, ranging from 0.16 to 0.52 MPa.\nCALCULATIONS\nAll calculations were based on hydrodynamic theory for the EHL isothermal model (Fig. 2) because of very low influence of changing temperature on water viscosity. Bearing bush deformation was calculated in the FEM commercial code.\nFig.2. Calculated temperature in water film in CFD software, shaft\ndiameter 100 mm, shaft rotation speed 660 rpm, diameter clearance 0,27mm, minimum film thickness 7 \u00b5m, bearing length 300mm\n1 Copyright \u00a9 2008 by ASME\nl=/data/conferences/ijtc2008/70338/ on 06/20/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use", + "Dow\nThe algorithm for the calculations consisted of the following two major parts: \u2022 calculation of pressure field in the water film in\nwater film, from a two-dimensional Reynolds equation,\n\u2022 calculation of bearing bush deformation using commercial FEM package. Both modules were solved simultaneously in an iterative process. The convergence criterion used was 0.1%. An example of the calculation convergence is illustrated in Figure 3. It shows that the fifth iteration produced a satisfactory solution. Relatively small variations of water viscosity with temperature justified using an isothermal elastohydrodynamic model of lubrication (Figure 2).\n2750\n2800\n2850\n2900\n2950\n3000\n3050\n3100\n1 2 3 4 5 6 7 8\niteration step\nB ea\nrin g\nhy dr\nod yn\nam ic\nlo ad\n[N ]\nFig.3. Calculated values of hydrodynamic pressure; step 1- Bering without bush deformation; steps 5\u00f78 results iteration, final result of\nbearing hydrodynamic capacity 2857N\nRESULTS OF CALCULATIONS\nA strong effect of bearing bush deformation resulting from hydrodynamic pressure was particularly evident at heavy bearing loads and thin hydrodynamic films. This is illustrated in Figures 4 and 5. As expected, the maximum deformation occurred in the bearing middle plane, while at both edges deformation was significantly smaller because of the zero pressure at the bearing side edges.\n1 9\n17 25 33 41 49 57 65 73 81 89 97 S1\n0,0\n1,0\n2,0\n3,0\n4,0\n5,0\n6,0\n7,0\nH yd\nro dy\nna m\nic P\nre ss\nur e\n[M P\na]\nBearing perimeter\nB er\nin g\nle ng\nht\nFig.4. Calculated hydrodynamic pressure in water lubricated bearing.\nBearing load 0.52MPa, shaft speed 600 rpm.\nnloaded From: http://proceedings.asmedigitalcollection.asme.org/pdfaccess.ashx?ur\n1 9\n17 25 33 41 49 57 65 73 81 89 97 S1\n0,00\n0,01\n0,02\n0,03\n0,04\n0,05\nB ea\nrin g\nbu sh\nd ef\nor m\nat io\nn [m\nm ]\nBering perimeter\nB ea\nrin g\nle ng\nht\nFig.5. Calculated bearing bush deformation, load 0.52MPa, shaft speed\n660 rpm.\nMEASUREMENTS The experimental rig has been described elsewhere [3]. Here is its schematic.\nDuring the tests, friction torque, water film pressure, and shaft center path were recorded. Coefficient of friction was calculated from friction torque (Figure 7). The test rig offered also a unique possibility of measurement of clearance circle without need for disassembly of the bearing unit. That allowed for monitoring of bearing clearance variations resulting from water soaking and wears (Figures 8 and 9).\n2 Copyright \u00a9 2008 by ASME\nl=/data/conferences/ijtc2008/70338/ on 06/20/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use", + "Do\n0\n0,01\n0,02\n0,03\n0,04\n0,05\n1 2 3 4 5 6 7 8 9 10 11\nshaft rotation speed [rev/s]\nco ef\nfic ie\nnt o\nf f ric\ntio n\n0,55 MPa 0,41 MPa 0,28 MPa\nFig.7. Measured coefficient of friction\n500\n550\n600\n650\n700\n750 500600700800900\n[\u00b5m]\n[\u00b5 m\n]\n3 [rev/s]\n7 [rev/s]\n11 [rev/s]\ncircle\nFig.8. Measured shaft center trajectories and clearance circle: effect of\nbearing bush deformation. Bearing load 0.16MPa\n500\n600\n700\n800 500600700800900\n[\u00b5m]\n[\u00b5 m\n]\n3 [rev/s]\n7 [rev/s]\n11 [rev/s]\ncircle\nFig.9. Measured shaft center path and clearance circle \u2013 effect of\nbearing bush deformation. Bearing load 0.52MPa\nwnloaded From: http://proceedings.asmedigitalcollection.asme.org/pdfaccess.ashx?ur\n0,0141 0,01503718\n0\n0,01\n0,02\n0,03\n0,04\n0,05\nC oe\nffi ci\nen t o f fri ct io n\nbearing circumference bearing lenght\nFig.10. Measured hydrodynamic pressure in bearing gap.\nCONCLUSIONS The experimental results indicate that at loads above 0.16 MPa and for rotation speed under 5 rev/s the test bearing operated in mixed lubrication conditions. Coefficient of friction is relatively low but the shaft center trajectories are outside the clearance circle. This proves the significant bearing bush deformation resulting from load and flexible material properties. The measurements of pressure in the bearing gap seem to indicate lack of fluid film between the shaft and the bush. This is a conclusion drawn form Figure10. The profile is relatively flat in the pressure zone, and is different from those typical for a hydrodynamic lubrication. It is a result of blocking pressure sensor which rotates with the shaft.\nThe calculations showed full hydrodynamic lubrication in the center of the bearing. At the edges, however, it is most likely that mixed lubrication conditions existed because the film thickness was significantly lower than the theoretically predicted minimum film thickness.\nREFERENCES [1] Litwin Wojciech Marine water lubricated stern tube\nbearings \u2013 design and operations problems STLE/ASME International Joint Tribology Conference 2007; San Diego USA\n[2] Dymarski C. Litwin W. Influence of surface roughness of\nbearing bush on properties of water lubricated main shaft bearings \u2013 experimental tests NORDTRIB 2008 TAMPERE FINLAND\n[3] Litwin W The tests stands for water lubricated marine\nmain haft bearings NORDTRIB 2006, Helsingor Denmark\n3 Copyright \u00a9 2008 by ASME\nl=/data/conferences/ijtc2008/70338/ on 06/20/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use" + ] + }, + { + "image_filename": "designv11_69_0003652_vetecf.2010.5594384-Figure5-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003652_vetecf.2010.5594384-Figure5-1.png", + "caption": "Fig. 5. Reference frames for the UAV", + "texts": [ + " The proportional-derivative (PD) gains are chosen as kp1 > 0 and kd1 > 0 for i = 1, 2. It is considered that m = 2 helicopter models are monitored by n = 2 different ground stations. The distributed filtering architecture is shown in Fig. 4. To succeed accurate localization of each UAV it is necessary to fuse the GPS measurements with the IMU measurements of the UAV or with measurements from visual sensors (visual odometry) [15]. The inertial coordinates system OXY is defined. Furthermore the coordinates system O\u2032X \u2032Y \u2032 is considered (Fig. 5). O\u2032X \u2032Y \u2032 results from OXY if it is rotated by an angle \u03b8. The coordinates of the center of symmetry of the UAV with respect to OXY are (x, y), while the coordinates of the GPS or visual sensor that is mounted on the UAV, with respect to O\u2032X \u2032Y \u2032 are x \u2032 i, y \u2032 i. The orientation of the GPS (or visual sensor) with respect to OX \u2032Y \u2032 is \u03b8 \u2032 i. Thus the coordinates of the GPS or visual sensor with respect to OXY are (xi, yi) and its orientation is \u03b8i, and are given by: xi(k) = x(k) + x \u2032 isin(\u03b8(k)) + y \u2032 icos(\u03b8(k)), yi(k) = y(k) \u2212 x \u2032 icos(\u03b8(k)) + y \u2032 isin(\u03b8(k)), \u03b8i(k) = \u03b8(k) + \u03b8i. The GPS sensor (or visual sensor i) is at position xi(k), yi(k) with respect to the inertial coordinates system OXY and its orientation is \u03b8i(k). Using the above notation, the distance of the GPS (or visual sensor i), from the plane P j is represented by P j r , P j n where (i) P j r is the normal distance of the plane from the origin O, (ii) P j n is the angle between the normal line to the plane and the x-direction: dj i (k) = P j r \u2212xi(k)cos(P j n)\u2212 yi(k)sin(P j n) (see Fig. 5). Results on the UIF and DPF performance in estimating the state vectors of multiple UAVs when observed by distributed processing units is given in Fig. 6 and Fig. 7, respectively. The advantages of using Distributed Filtering are as follows: (i) there is robust state estimation (which in the DPF case is not constrained by the assumption of Gaussian noises). The fusion between the local filters compensates for deviations in state estimates (which in the UIF case can be due to linearization errors while in the DPF case can be due to outlier particles) thus resulting in an aggregate state distribution that confines with accuracy the real state vector of each UAV" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001950_1-84628-179-2_5-Figure5.6-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001950_1-84628-179-2_5-Figure5.6-1.png", + "caption": "Fig. 5.6. Coaxial helicopter configuration.", + "texts": [ + "40) Finally, the total torque applied to the tandem rotor helicopter (expressed in the body fixed frame) is given by \u03c4 = \u03c4NT + |QN |E3 \u2212 |QT |E3 (5.41) Substituting the total force (5.35) and total torque in a Newton basic model, we have \u03be\u0307 = \u03c5 (5.42) m\u03c5\u0307 = (|TN |\u03b2N + |TT |\u03b2T )RE2 \u2212 (|TN | + |TT |)RE3 + mgEz (5.43) R\u0307 = R\u2126\u0302 (5.44) I\u2126\u0307 = \u2212\u2126 \u00d7 I\u2126 + |QN |E3 \u2212 |QT |E3 + \u03c4NT (5.45) A coaxial helicopter is again a twin main rotor helicopter configuration that uses two contrarotating rotors of equal size and loading and with concentric shafts (see Figure 5.6). Some vertical separation of the rotor disks is required to accommodate lateral flapping. Pitch and roll control is achieved by the cyclic pitch of the swash plate. The height control is achieved by the collective pitch. The yaw control mechanism is more subtle. When a rotor turns, it has to overcome air resistance, so a reactive force acts on the rotor in the direction opposite to the rotation of the rotor. As long as all rotors produce the same torque, they produce the same reactive torque. This torque is mostly a function of rotation speed and rotor blade pitch", + " It is important to note that this operation has no effect on translation in the x or y direction in a coaxial helicopter configuration. For the sake of simplicity we present here the dynamic model of a coaxial helicopter in hovering. We use the same assumptions taken for the standard helicopter, adding/changing only the following assumptions: 1C The above rotor blades are assumed to rotate in an anti-clockwise direction when viewed from above and the down rotor blades rotate in a clockwise direction, see Figure 5.6. 2C The operation of two or more rotors in close proximity will modify the flow field at each, and hence the performance of the rotor system will not be the same as for the isolated rotors. We will not consider this phenomenon to simplify the dynamic model. We have separated the aerodynamic forces into two parts. The first part is composed of the total translational forces applied to the coaxial helicopter and the second part is related to the sum of the rotational torques. More details on coaxial helicopter dynamics can be found in [78]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000512_810105-Figure8-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000512_810105-Figure8-1.png", + "caption": "Fig. 8 - Original point setting device", + "texts": [ + " The positioning accuracy of the machine is 1 \u0302 m and the angular accuracy of the rotational table is 2 sec. The enlarged view of the apparatus in operation is shown in Fig. 6, where we see the probe detecting the deviation of the tooth surface in y direction. The probe is held rigidly in the lateral direction (x direction) to prevent it from deflection when placed on the inclined tooth surfaces. It has been confirmed that the lateral force of 50 grw causes only 1 /im deflection as shown in Fig. 7. The setting device of the gear is shown in Fig. 8 and Fig. 9. It has a setting plane and a setting ball which are used to determine the original point, 0, of the measurement. The x and z coordinates of the center of the ball, Xe and Ze, and the y coordinate of the setting plane, Ye, have been measured in advance. At first the y-setting plane of the device is set parallel to the x axis. The x coordinates Xi and X2 are measured from an arbitrary point for two positions of the probe which give the same reading of the deflection of the probe in y direction" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000729_bf02644088-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000729_bf02644088-Figure1-1.png", + "caption": "Fig. 1--An antiphase boundary tube as suggested by Vidoz and", + "texts": [ + " 1 as a function of aging t ime at 650~ and 540~ The s p e c i m e n aged at 650~ for 20 min had a flow s t r e s s of 53 K $ / m m z, which c o r r e s p o n d s a p p r o x i - m a t e l y to the m a x i m u m va lue ob ta ined in the s e r i e s of the s p e c i m e n s aged at 650~ and was about 20 K g / m m 2 h igher than the va lue ob ta ined f rom the so lu - t ion t r e a t e d s p e c i m e n . The s a m e value of flow s t r e s s was a t t a ined in the s p e c i m e n aged at 540~ for 10 min, ~\" 9 0 ~ , E E \"\" 80 ~' Aged ot 540% -~ zo O 5O I I ,=4O,,, 3 d ;<-- Solution treoted u_ I0 I0 = I0 ~ 104 Aging time (ram) Fig. 1 - -F low s t r e s s as a function of the aging t ime at 650~ and 540~ 20P ~-,. m ~ 4 - .7. Fig. 2--Scanning e l ec t ron m i c r o g r a p h showing duct i le d imple s u r f a c e a p p e a r a n c e of a s p e c i m e n aged at 650~ for 20 min. METALLURGICAL TRANSACTIONS A VOLUME 7A,JUNE 1976-889" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003293_s0263574710000421-Figure3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003293_s0263574710000421-Figure3-1.png", + "caption": "Fig. 3. The 5R manipulator and some of its singular configurations.", + "texts": [ + " Since the base module is a 3SPS+3RRRS decoupled parallel manipulator, the singularity analysis is approached by analyzing separately the spherical and translational parallel manipulators. The singularity analysis of the 3SPS parallel manipulator was successfully approached in Alici and Shirinzadeh46 and, therefore, it is unnecessary to include it here. On the other hand, it is evident that the singularities of the translational parallel manipulator can be investigated by considering it as a 5R closed chain; see Fig. 3. Initially, the manipulator under study is modeled as an open serial chain in which the velocity state of body 5 with respect to body 0 is given by 0\u03c91 0$1 + 1\u03c92 1$2 + 2\u03c93 2$3 + 3\u03c94 3$4 + 4\u03c95 4$5 = 0V 5, (23) where the screws are reduced to three-dimensional vectors. As demonstrated by Rico et al.,47 in a serial manipulator the screws connecting it to the base link and the end-effector are not responsible to fall or to escape the manipulator from a singular configuration and therefore the joint rates 0\u03c91 and 4\u03c95, associated with the active joints q1 and q2, respectively, must be disregarded immediately from the analysis, even though they are the motors of the manipulator", + " On the other hand, if body 5 is joined to body 0, then the open manipulator becomes a closed chain where 0V 5 = 0. Thereafter, Eq. (23) is rewritten in a matrix\u2013vector form as Js \u23a1 \u23a3 1\u03c92 2\u03c93 3\u03c94 \u23a4 \u23a6 = \u2212 0\u03c91 0$1 \u2212 4\u03c95 4$5, (24) where Js = [1$2, 2$3, 3$4]. In order to solve (24) it is necessary that det(Js) = 0. In other words, the closed chain is at a singular configuration if the screws 1$2, 2$3, and 3$4 are linearly independent, which implies that dim(Js) < 3, which occurs mainly when the revolute joints of such screws are aligned; see Fig. 3. In what follows, it is shown how the closed chain can escape from a singularity by means of a simple case. To this end, consider the singular configuration depicted in Fig. 4. In order to escape from the singularity, the following steps are suggested: (1) Detect the singularity. Since det(Js) = 0, the closed chain is at a singular configuration. (2) Lock the revolute joint q1, e.g. q\u03071 = 0, and unlock q2 such that this revolute joint becomes a passive element. (3) Consider the third limb, containing q3, as an active leg which implies that point C \u2032 can move along a circular trajectory" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001327_1.2080521-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001327_1.2080521-Figure1-1.png", + "caption": "FIGURE 1 \u2014 Schematic drawing of bowing phenomenon in biaxial stretching. The distortion of the principal axis of refractive index in the TD is depicted.", + "texts": [ + " However, it is most difficult to fabricate due to the complicated manufacturing processes. The conventional fabrication process is biaxial stretching, in which a film is stretched in two orthogonal directions, a machine direction (MD) and a transverse direction (TD). For instance, during stretching in the TD, a tensile strain also exists in the MD. Since the width of the film is finite, the tensile strain in the MD depends on the position of the width. Therefore, the principal axis of refractive indices of biaxial films in the TD is inclined, called \u201cbowing,\u201d3 as shown in Fig. 1. Bowing occurs the most in the TD stretching process. It is reported that a deformed helix cholesteric network can create birefringence in the plane of the film and become biaxial.4 The deformed cholesteric LC nanostructure is induced by the exposure of polarized ultraviolet (UV) light with a rod-like dichroic photoinitiator. The principal axes in a plane of films can be uniformly controlled by a polarization plane. A VA-LCD compensated by using a deformed cholesteric film has been proposed5; however, the deformed cholesteric LC nanostructure is not directly investigated" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000006_jsvi.2002.5135-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000006_jsvi.2002.5135-Figure1-1.png", + "caption": "Figure 1. Schematic representation of motion of four-beam flexible system with a central rigid body: (a) initial configuration; (b) initial movement; (c) motion after first locking; (d) end configuration.", + "texts": [ + " A brief description of the flexible deployment system is given in section 2 and the dynamic equations of flexible multi-body system with internal impact are established in section 3. A \u2018\u2018longitudinal constraint\u2019\u2019 is suggested to decrease the stiffness of the equations in section 4. The simulation results of the deployment system are presented in section 5. Finally, the experimental set-up and results are presented and are compared with numerical simulation. The deployment system investigated in this paper is shown in Figure 1. It includes a central rigid body and four articulated flexible beams. The system is folded into an initial position, as shown in Figure 1(a), and is expanded by releasing the compressed torsion springs in the joints. In the deployment, the beams experience both large rigid-body motions and elastic vibrations, as shown in Figure 1(a)\u20131(d). When joints are locked at appropriate positions, the impacts in the joints will induce elastic vibrations. The purpose of this section is to establish the dynamic equations of the deployment system described in the above section. Section 3.1 deals with the finite segment modelling of the flexible beams that experience large rigid-body motions. Section 3.2 introduces the mathematical model of the locking impacts in the joints and section 3.3 establishes the dynamic equations of motion of the flexible deployment system" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003121_9780470027318.a9083-Figure6-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003121_9780470027318.a9083-Figure6-1.png", + "caption": "Figure 6 Interaction of active site of the enzyme with the electrode surface. (a) Simple contact between the active site and the electrode. (b) Contact combined with enzymatic reaction. (c) Stepwise ET between the active site and the electrode via multiple cofactors, e.g. Fe\u2013S clusters, hemes etc. Distance between the active site and the electrode surface is approximately 7\u201330 A\u030a (10\u221210 m). The ET may occur in both directions, depending on the applied potential.", + "texts": [ + " Reconstitution can be obtained either directly in solution by implementing the cofactor into the apoprotein, or first by immobilizing the cofactor unit on the electrode surface, followed by the interaction with the respective apoprotein. Reconstitution method was demonstrated for various proteins, including flavoproteins (e.g. glucose oxidase (GOX),(91) quinoproteins,(92) or hemoproteins(93)). The mechanism of flux of electrons during DET may be described in several ways. The simplest way of protein\u2013electrode interaction is through the movement of the electron from the redox center to the electrode (Figure 6(a)). Another mechanism is depicted in Figure 6(b), when the DET is coupled with an enzyme reaction. An even more complicated system is presented in Figure 6(c). Many enzymes capable of DET contain more than one redox center and thus the electron goes through a cascade of coenzymes on its way to the electrode. Several redox enzymes exhibiting DET are listed in Table 1. Some examples of individual proteins are now described in greater detail. The copper redox state in the catalytic centers of proteins \u2013 mostly enzymes \u2013 is either Cu2+ or Cu+. According to the mode of binding of the metal, there are three types(73): T1, blue copper type; T2, normal copper type; and T3, coupled binuclear copper centers" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003010_20100906-5-jp-2022.00015-Figure3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003010_20100906-5-jp-2022.00015-Figure3-1.png", + "caption": "Fig. 3. Aerosonde (image credit to NASA)", + "texts": [ + " The MPC uses the linear discrete-time model for prediction whereas the turn coordination and the dampers act on the non-linear system in continuous time. This model mismatch might potentially lead to constraint violations, in particular for steep manoeuvres. The proposed three-layer control structure is now demonstrated for the Aerosonde, a small fixed-wing umanned aircraft. The objective is to demonstrate the implementation of each layer of the proposed framework and the overall control structure. The Aersonde (Fig. 3) is a small fixed wing unmanned civil aircraft developed and operated by Aerosonde Pty Ltd and Aerosonde North America (www.aerosonde.com). Its wingspan is 2.9m and its maximum airspeed is around 32m/s. For simulation we use the Aerosim blockset for Matlab/Simulink (www.u-dynamics.com). Even though the Aerosonde has a V-tail, this blockset uses the traditional setup with elevators and rudder, accounting for it in the aerodynamic coefficients. The Aerosonde\u2019s dynamic behaviour is described by a standard non-linear six degree of freedom model (Stevens and Lewis (2003))" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001164_1-84628-559-3_20-Figure7-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001164_1-84628-559-3_20-Figure7-1.png", + "caption": "Fig. 7. Positioning stage and base.", + "texts": [ + " The friction adjusting mechanism set on the top surface of the stage is used for providing a suitable holding fricitonal force for the rotaional stage. Figure 6 shows a photograph of the modualrized springmounted PZT actuator, in which the PZT actuator has the dimension of 5 5 10 mm (Tokin). The stiffness of the spring is 0.023 N/mm. Six actuating units are symmetrically mounted to the positioning stage having a radius on the bottom side. The positioning stage made of stainless steel having a mass of 2.6 kg is then set on the base having the same radius on the contace surfaces. Figure 7 shows the positioning stage and the base. Figure 8 shows the 3-DOF positioning stage. The waveform of applied voltage was generated by LabVIEW, which is a Windows supported graphical programming language. A 16-bit DA/AD converter was used to transform the pulse driving waveform to the power amplifier and then to the PZT actuator. Due to the number limitation of measuring probes, two kinds of gap sensors were used to detect the simultaneous motion behaviors of the positioning stage along three rotation axes" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002979_978-3-540-70534-5_7-Figure11.3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002979_978-3-540-70534-5_7-Figure11.3-1.png", + "caption": "Figure 11.3: Andy Droid humanoid robot", + "texts": [ + " Three servos are used to bend the leg at the ankle, knee, and hip joints, all in the same plane. One servo is used to turn the Biped Robot Design 169 leg in the hip, in order to allow the robot to turn. Both robots have an additional dof for bending the torso sideways as a counterweight. Jack is also equipped with arms, a single dof per arm enabling it to swing its arms, either for balance or for touching any objects. A second-generation humanoid robot is Andy Droid, developed by InroSoft (see Figure 11.3). This robot differs from the first-generation design in a number of ways [Br\u00e4unl, Sutherland, Unkelbach 2002]: \u2022 Five dof per leg Allowing the robot to bend the leg and also to lean sideways. \u2022 Lightweight design Using the minimum amount of aluminum and steel to reduce weight. \u2022 Separate power supplies for controller and motors To eliminate incorrect sensor readings due to high currents and voltage fluctuations during walking. Figure 11.3, left, shows Andy without its arms and head, but with its second generation foot design. Each foot consists of three adjustable toes, each equipped with a strain gauge. With this sensor feedback, the on-board controller can directly determine the robot\u2019s pressure point over each foot\u2019s support area and therefore immediately counteract to an imbalance or adjust the walking gait parameters (Figure 11.3, right, [Zimmermann 2004]). Andy has a total of 13 dof, five per leg, one per arm, and one optional dof for the camera head. The robot\u2019s five dof per leg comprise three servos for bending the leg at the ankle, knee, and hips joints, all in the same plane (same as for Johnny). Two additional servos per leg are used to bend each leg side- Walking Robots 170 11 ways in the ankle and hip position, allowing the robot to lean sideways while keeping the torso level. There is no servo for turning a leg in the hip" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002690_robio.2007.4522277-Figure8-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002690_robio.2007.4522277-Figure8-1.png", + "caption": "Fig. 8. Manipulating a thin and light object by two robot fingers in 2D plane", + "texts": [ + "(31), it is also possible to derive the following equation: \u03b8\u0307 = 1 l1 + l2 { (x\u030701 \u2212 x\u030702) sin \u03b8 + (y\u030701 \u2212 y\u030702) cos \u03b8 \u2212 (r1 \u2212\u0394x1) q\u030711 + (r2 \u2212\u0394x2) q\u030721 } (40) Then, by virtue of the convergence that q\u0307i1 \u2192 0(i = 1, 2) as t \u2192 \u221e, we can show that \u03b8\u0307 of equation (40) converges to zero as t\u2192\u221e. Finally, we can prove that as t\u2192\u221e{ q\u0307i(t)\u2192 0, \u0394x\u0307i(t) \u2192 0, x\u0307(t)\u2192 0, \u03b8\u0307(t) \u2192 0 qi(t)\u2192 qd, \u0394xi(t) \u2192 \u0394xid, x(t) \u2192 xd, \u03b8(t)\u2192 \u03b8d (41) and, at the same time, the following equation concerning the total moment of pinched object must be satisfied as t\u2192\u221e: fi (Y1 \u2212 Y2) + (\u22121)i\u03bbi (l1 + l2) = 0, i = 1, 2 (42) However, in order to analyze stability of dynamic grasp of a thin and light rigid object as shown in the finger-object model of Fig.8, a smaller width li(i = 1, 2) of the object makes eq.(40) useless, because \u03b8\u0307 of eq.(40) may diverge to infinity. Therefore, we can not prove the convergence of \u03b8\u0307 with the aid of LaSalle\u2019s invariance theorem when li becomes far small. However, in the next section, we show that two robot fingers can pinch a thin and light rigid object stably, that will be validated by numerical simulation in this paper. From this observation, the control signals of eq.(35) must be applied even for the case of grasping a thin rigid object, but we must develop a new stability proof including exponential convergence of q\u0307i1(t), x\u0307(t), and \u03b8\u0307(t) to zero as t\u2192\u221e. The details of this stability proof will be presented in a future work. Numerical simulation is carried out by using the fingerobject model as depicted in Fig.8. The purpose of this simulation is to show how important is the visco-elastic characteristics of finger-tip material in stabilization of dynamic grasp of various kinds of objects. Finger models with soft finger tip characteristics are shown in Table V and parameters of control signals are given in Table VI. Then, it is interesting to compare two simulation cases of a soft tip and a rigid one by using the same parameters of the fingers and object except finger-tip characteristics. In this case of grasp of such a thin and light object, stiffness parameter k, damping coefficients c\u0394 and c\u03b8 are chosen reasonably small specified in Table V (compare this table with Table III)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002418_s1052618807020082-Figure3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002418_s1052618807020082-Figure3-1.png", + "caption": "Fig. 3. (a) Diagram of the action of forces in a friction pendulum bearing and (b) the hysteresis loop. Here, fd is the dynamic friction coefficient, D is the projection displacement, F = fdW + [W/R]D is the maximum horizontal force, Ki = fdW/Dy is the initial rigidity, Keff = F/D is the effective rigidity, K = W/R is the rigidity of the friction pendulum bearing, and Dy = 0.24 cm is the initial displacement at the start.", + "texts": [ + " Operation of a friction pendulum bearing: (a) pendulum motion, (b) the sliding pendulum motion of a friction pendulum bearing, (c) the operation of a bearing, (d) schematic diagram of the location of a friction pendulum bearing under a structure. 1 34 2 Fig. 2. Diagram of the arrangement of a friction pendulum bearing. (1) Plate with a concave surface; (2) plate of the case; (3) swing sliding block; (4) self-lubricating material on the surface of the swing sliding block. JOURNAL OF MACHINERY MANUFACTURE AND RELIABILITY Vol. 36 No. 2 2007 THE EFFECT OF EARTHQUAKE PARAMETERS 149 Figure 3 demonstrates the properties of the friction pendulum bearing depending on friction, construction parameters, and vertical load [9, 10]. Friction pendulum bearings can be modeled as bilinear hysteresis elements. The lateral restoring rigidity of the friction pendulum bearing is given by K = W/R, where W is the vertical load and R is the curvature radius of the concave surface. The natural period of horizontal vibrations for the support with the friction pendulum bearing is determined from the pendulum equation T = , where l is the length of the curvature radius R (the equivalent pendulum length)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001384_j.ijmachtools.2006.02.011-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001384_j.ijmachtools.2006.02.011-Figure1-1.png", + "caption": "Fig. 1. Concept of heat ablation using a hot tool.", + "texts": [ + " Finally, the practical applicability of the RHA process is then demonstrated by the results of these fabrications for a hemispherical shape and a standard test part for machining in terms of geometrical conformity, volume of remaining material, ablating time and dimensional accuracy. Heat ablation process is used with readily meltable workpiece. The principle of heat ablation is to melt and decompose any remaining material while a hot tool, heated to above the decomposition temperature, moves along a predetermined path, as illustrated in Fig. 1. The sequence of material removal mechanism is as follows. The heated tool makes the thermal field of material. According to temperature distribution by the thermal field, material phase is changed at each section. Material at the section of decomposition in the vicinity of the tool is vaporized. The ARTICLE IN PRESS H.C. Kim et al. / International Journal of Machine Tools & Manufacture 47 (2007) 124\u2013132126 molten material is mechanically removed by the movement of the tool. Consequently, the sections of melting and decomposition are removed", + " / International Journal of Machine Tools & Manufacture 47 (2007) 124\u2013132128 (3) Insulator: An insulator connected with a tool of surface temperature over 5001C directly should have good heat resistance and stability at high temperatures. A ceramic material with the required characteristics was chosen for the insulator. (4) Tool: In order to minimize the radius of the material removal zone during the process, the hot tool has tangential grooves separated into two regions. The workpiece is decomposed in the inner region and melted in the outer region, as shown in Fig. 1. In the present work, copper was used as the tool material for effective heat transference. In the RHA process, when the heated tool ablates the material, the kerfwidth, the ablated width of the workpiece, is strongly related to the heat-affected zone resulting from the hot tool. Therefore, the heat radius should be investigated to control the kerfwidth for precise cutting. During a cutting process using a heat source, such as laser, plasma, or hotwire, the main process parameters that determine the kerfwidth are generally the cutting speed and the heat input [10,12,13]", + " In the RHA process, the thermal characteristics in the Z-direction as well as in the X-direction should be investigated, because the hot tool ablates the workpiece moving on an XZ plane. Moreover, the thermal characteristics in the Z-direction are the main influence on the surface quality of a part. In order to predict the virtual tool shape formed by the heat radius, the heat radius in the Z-direction should be investigated when a toolpath is generated. Experiments were carried out with ablation at various depths and effective heat inputs. As illustrated in Fig. 1, the term a is defined as the heat radius minus the tool diameter from the kerfwidth in the X-direction. The term b is defined as the heat radius minus the tool length from depth of cut in the Z-direction. The relationship between the heat radius in the Xdirection and the effective heat input was obtained through experiments. Therefore, the relationship between the heat radius in the Z-direction and the process parameter can be obtained through comparison with the heat radius in the X-direction" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000910_detc2004-57046-Figure3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000910_detc2004-57046-Figure3-1.png", + "caption": "Figure 3. TROCAR POINT KINEMATICS", + "texts": [ + " The signals from the sensors are amplified and transmitted via CAN-bus to a PC system. Sensor readings are blurred with noise, hence we have applied digital filters to stabilize the results. Since we know the position and orientation of the instruments, we can transform occurring forces back to the coordinate system of the PHANToM devices. Therefore the user has the impression of direct haptic immersion. The basic idea of minimally invasive surgery is, that only small openings have to be made into the surface of the patient\u2019s thorax (so-called keyholes, Fig. 3). That means the translational movements of the instruments are essentially restricted by shifts and rotations about these holes. In order to provide the surgeon with a comfortable environment, it is desirable to map the move- Copyright c 2004 by ASME rms of Use: http://www.asme.org/about-asme/terms-of-use ments of the stylus at the input device directly to instrument motions. Therefore we have to consider the inverse kinematics of our system. That means we have to find a mapping of an arbitrary posture of the instrument\u2019s tip to a position of the motors that control the eight degrees of freedom" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000429_6.2003-5652-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000429_6.2003-5652-Figure2-1.png", + "caption": "Fig. 2 Engagement geometry.", + "texts": [ + "15/)(0216.01)(0404.01( )0197.1)(0096.1)(007.01( )( 22 ++++ ++\u2212 == ssss sss sG K a c P \u03b4 (7) The gains cK , tK and K in the above TFs are obtained from the steady state conditions and are used as scale factors. The step responses of the TFs of Eqs. (5)-(7) are plotted in Fig. 1. 2.2 Guidance dynamics The guidance problem is treated in this paper as planar. It is assumed that during the endgame deviations from collision triangle are small, justifying linearization. The engagement geometry is plotted in Fig. 2 where the X axis is along the initial line of sight and Y is perpendicular to it. Note that the subscript P denotes the pursuer (the interceptor) and the subscript E denotes the evader (the target). For the guidance laws development simplified closedloop flight control dynamics is assumed. Fig. 3 shows the classical guidance and control architecture where the guidance law issues a single maneuvering command. The flight control system is responsible for the distribution of the commands to the two actuators following a heuristic rule (e", + " Hence tciVu ci ,; 2 3 == \u03bbD (45) 4. PERFORMANCE ANALYSIS 4.1 Tools 4.1.1 Guidance dynamics A general block diagram representing the guidance dynamics chosen for the guidance law evaluation is given in Fig. 6. Note that unlike the simplified guidance dynamics of Figs. 3 and 4 used for the guidance law development, here high order interceptor dynamics and an estimator are used. It is assumed that the interceptor acquires the following measurement wrywM +\u2245+= /\u03bb (46) at a given rate f, where r is the range (see Fig. 2) and w is the measurement error, modeled as a white Gaussian noise: w ~ N(0, 2 \u03bb\u03c3 ) (47) For the estimator design, the well known exponentially correlated acceleration (ECA) shaping filter, associated with the Singer model [14], is used. The state vector of the filter is T],,[ Eayy D=W (48) The equations of the filter model are: \u03c9wPww a CBWAW ++= (49) where = \u2212= \u2212 = 1 0 0 ; 0 1 0 ; /100 100 010 ww a w CBA \u03c4 (50) and ),0(N~ Q\u03c9 (51) The parameter a\u03c4 is the assumed average time between the changes in the piecewise constant target acceleration levels" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000638_sensor.2003.1217016-Figure3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000638_sensor.2003.1217016-Figure3-1.png", + "caption": "Figure 3: (a) Case 1 and (b) Case 2.", + "texts": [ + " For this case, the pressure in the bulge of the capillary fluid is higher than that of the assembly fluid, which stabilizes the z force balance. The angle of tilt can either increase or decrease until equilibrium is reached. Figure 6 shows a sequence of video frames of this phenomenon. The bulge curvature changes slightly with time, and the tilting of the micropart decreases from 18\" until it reaches an equilibrium position at 16\". When the initial capillary volume is less than the critical volume and bulging of the capillary fluid does not occur in that plane as shown in Figure 3h, til! correction and squeeze-film settling proceeds quite quickly as predicted by theory. Video frames of such a case are presented in Figure TRANSDUCERS '03 The 121h International Conference On Solid State Sensors, Actuators and MicrOSyStemS, Boston. June 8-12, 2003 7. The initial angle of tilt is approximately 3\u201d-5\u201d, which is within the range of the small angle approximation taken in the models. Not much appears to happen for the first 3 seconds. Then the majority of the tilt is removed within one frame of the next second, which is in agreement with the previously presented theories. The video capture equipment used for these experiments is only capable of 30 kames per second, which corresponds to a resolution of approximately 35 ms. Therefore, 11 was not possible to capture any slow motion videos of these processes. The capillary fluid volume is minimized along the edge in which tilt correction nrrllrF Reducing the volume of capillary liquid coated onto the substrate binding site below the critical volume prevents tilted equilibrium positions from forming, as shown in Figure 3b. Tilt correction only occurs when the volume of capillary liquid initially coated onto the substrate is at or below a critical volume. For a maximum initial angle of 15\u2019 and the given parameters, the critical volume is 7.6 nL. Three distinct processes (tilt correction, force stabilization, and squeeze-film settling) have been mathematically represented in a single model to describe micropart equilibrium positions and tilt correction. The model presented describes these processes for microparts assembled by capillaq forces and its prediction is found to be consistent with experiment" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000373_1.3453903-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000373_1.3453903-Figure1-1.png", + "caption": "Fig. 1 Cross section of rotor assembly", + "texts": [ + "1vity effects on the vertical rotor arrangement. The current experimental study was undertaken to investigate the influence of radial cleara1lce, unbalance and inlet oil .prcssme on the performance of a central groove oil-film damper for both concentric and eccentric orbits. The modifications of the rig, the instrumentation, experimental procedure and some typical test results are reported in this paper, The Oil-Film Damper Rig A cross-sectional view of the squeeze film bearing arrange ment is iihown in Fig. 1. The onter race of the duplex ball bearing is shrunk ilt inside the damper journal which is held in the damper bearing by a retaining ring and has complete lateral freedom of movement. The damper journal has a nom inal outside diameter of 7.6 em (3 in.) and lL length of 2,6 cm (1 in.), A central oil-groove of about 1.25 cm width divides the journal into two equal lands. 'fhi\" rather wide groove was requiled to ,fit the off-centered oil supply holes while rett1ining a symmetric damper, The oil film is limited in length at each end by a steel pis tOil 8el11 ring" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001712_nme.1428-Figure3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001712_nme.1428-Figure3-1.png", + "caption": "Figure 3. Frictional Hertzian contact.", + "texts": [ + " The special cases discussed in Section 5 serve as an illustration of the correctness of the proposed Gauss\u2013Jacobi quadrature formulae in special cases. In this section two further numerical examples drawn from typical sliding contact problems will be considered. Firstly the sliding Hertzian contact in the presence of friction is solved. Secondly, fully sliding contact is considered between a rigid flat punch with rounded corner and an elastic substrate. Frictionless contact for this case has been investigated by Ciavarella et al. [20]. The problem of a rigid cylindrical punch sliding on an elastic half-plane is considered in Figure 3. The problem is solved by the numerical method presently proposed. Note, however, that this problem may also be solved in closed form analytically. Comparing the results obtained Copyright 2005 John Wiley & Sons, Ltd. Int. J. Numer. Meth. Engng 2005; 64:1236\u20131255 by the two different approaches provides a good method for the evaluation of the accuracy of the numerical method proposed in this paper. Since the punch is rigid we are concerned solely with the deformation of the half-plane. The normal displacement of the surface, v(x), is related to the normal and shear tractions p(x) and q(x) as follows [21] \u22121 \u2212 2 2 q(x) + 1 \u2212 \u222b a1 \u2212a1 p(t) t \u2212 x dt = v(x) x = \u2212 (x \u2212 e) R (53) where is Poisson\u2019s ratio, is the modulus of rigidity, R is the radius of the contacting cylinder and e is the eccentricity" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000039_004-Figure7-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000039_004-Figure7-1.png", + "caption": "Figure 7 Magnification of virtual image", + "texts": [ + "55 This content was downloaded on 29/09/2013 at 16:02 Please note that terms and conditions apply. 1979 Phys. Educ. 14 301 (http://iopscience.iop.org/0031-9120/14/5/004) View the table of contents for this issue, or go to the journal homepage for more Home Search Collections Journals About Contact us My IOPscience Phys Educ V01 14, 1979 Printed In Great Brllain Figure 6 Comparison of holography with optical imaging using lenses. In (a) the image is real; in (b) and (c) the image is virtual of magnifying the virtual image is to move the hologram along. the expanding laser beam (figure 7a); the laser light is diffracted at the same angle, but the distance AB increases. The dimensions of the image also depend on the form of the incident wave (figure 7b, c, d). When the incident wave is convergent, then for a certain angle the virtual image disappears (the real image, of course, still exists but becomes smaller). This is so because the expanding wave of the virtual image is transformed into a plane wave (figure 7e). The experiments described above lead to a better understanding not only of the holographic method of producing images, but also of optical imaging by other optical systems. Holography may be used as an attractive means of teaching the principles of optics, especially in the newer physics courses. Born M and Wolf E 1964 Principles of Optics (Oxford: Butters J 1971 Holography and its Technology (London) Halliday D and Resnick R 1962 Physics, Part 2 (New York: Hands R A 1972 Light for Advanced Courses (London: Kock E 1975 \u2018Sound visualisation and holography\u2019 Phys" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001118_0471758159.ch5-Figure5.37-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001118_0471758159.ch5-Figure5.37-1.png", + "caption": "FIGURE 5.37 (a) A simple way of winding a common-mode choke; (b) parasitic capacitance.", + "texts": [ + " In order to provide this impedance to common-mode currents, the wires must be wound around the core such that the fluxes due to the two common-mode currents add in the core whereas the fluxes due to the two differential-mode currents subtract in the core. Whether the wires have been wound properly can be checked with the right-hand rule, where, if one places the thumb of one\u2019s right hand in the direction of the current, the fingers will point in the resulting direction of the flux produced by that current. A foolproof way of winding two wires (or any number of wires) on a core to produce this effect is to wind the entire group around the core as illustrated in Fig. 5.37a. In either case one should ensure that the wires entering the winding and those exiting the winding are separated from each other on the core, or else the parasitic capacitance between the input and output will shunt the core and reduce its effectiveness, as illustrated in Fig. 5.37b. The effectiveness of the common-mode choke relied on the assumption that the self and mutual inductances are equal, L \u00bc M. High-permeability cores tend to concentrate the flux in the core and reduce any leakage flux. Symmetric windings also aid in producing this. Unfortunately, ferromagnetic materials suffer from saturation effects at high currents, as discussed earlier, and their permeabilities tend to deteriorate with increasing frequency more than low-permeability cores. One of the most important advantages of the common-mode choke is that fluxes due to high differential-mode currents cancel in the core and do not saturate it" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002597_acc.2008.4586601-Figure11-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002597_acc.2008.4586601-Figure11-1.png", + "caption": "Fig. 11. Tortoise", + "texts": [ + " For the simulation it is assumed that the period T = 1sec and the initial starting angle \u03b8(0) = \u03c0/17. A standard gradient descent algorithm has been implemented to find the optimal control \u03c9(t). The results are shown in Fig. 7, Fig. 8, Fig. 9 and Fig. 10. As in Fig. 10 the creature goes backwards initially when opening its flappers and moves forward when they are closed. The net effect is a forward motion due to the differential friction. Previous discussions can be extended to other robotic devices. One could offset two flappers to obtain a tortoise Fig. 11. In this case the flapper angles \u03b81 and \u03b82 have opposite phases to accommodate locomotion. In the next section we will extend the flapper to a simplified snake model. We extend the previous analysis to a simplified model of a snake, Fig. 12. We will derive the friction forces exerted by the environment on the snake. First, consider a small piece with length dr of the snake from either the upper or lower bar, which is located at r(t) with respect to a inertial reference frame with basis {ex, ey}, Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002320_12.715862-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002320_12.715862-Figure1-1.png", + "caption": "Figure 1: Morphing aircraft (Lee et al., 2005) and (Bae et al., 2005)", + "texts": [], + "surrounding_texts": [ + "The theory of operation of the sensor network and its interaction with the wing and aircraft dynamics are introduced. Comparisons between the desired and actual shapes of the morphing wings are also presented for different control law parameters. The developed theoretical and experimental techniques provide invaluable tools for the design and operation of other classes of remotely-controlled morphing aircraft. [Work is funded by NSF].\nKeywords: wireless and distributed sensor network, morphing structures, morphing wing, shape and health monitoring, wireless feedback control.\n1. INTRODUCTION\nNew designs for bio-inspired morphing wing emerge every year to harness the flexibility of different flight regimes. The concept of the ability to optimize wing shapes to correspond with an ad hoc flight is an appealing concept that has already been adopted to some extent in use today such as the AFTI/F-111 Mission Adaptive Wing (MAW) with its variable camber wing (Cesnik et al., 2004) and NASA\u2019s hyper-elliptic cambered span (HECS) wing (Davidson et al., 2003; Manzo et al., 2005). Flexibility of wings allow for adaptable maneuvering, velocity, endurance, and payload capabilities. Morphing wings have also come in many different designs such as a variable span morphing wing (Bae et al. 2005), folding wings, and sliding skin wings (Lee et al. 2005).\nSensors and Smart Structures Technologies for Civil, Mechanical, and Aerospace Systems 2007, edited by Masayoshi Tomizuka, Chung-Bang Yun, Victor Giurgiutiu, Proc. of SPIE Vol. 6529,\n65290P, (2007) \u00b7 0277-786X/07/$18 \u00b7 doi: 10.1117/12.715862\nProc. of SPIE Vol. 6529 65290P-1\nDownloaded From: http://proceedings.spiedigitallibrary.org/ on 06/26/2016 Terms of Use: http://spiedigitallibrary.org/ss/TermsOfUse.aspx", + "z\nDesigns for morphing wings also include polyhedral wing shapes which allow for variations between dihedral and anhedral wing sections. Refinement of these sections lead to near continuous morphing cambered span wing shapes (Wiggins et al., 2004). These designs focus on vertical deflection rather than horizontal orientation. Wings retain general dimensions, but introduce gradual bends to optimize aerodynamics for the desired flight.\nProperly controlling a wing deflection requires wing health monitoring. Though many means of sensing exist, most are designed for single point input. However, the sensor posed by Akl et al., (2006) suggests a distributed sensor network that allows for interpolation allowing for assurance of aerodynamically smooth and wrinkle free surfaces. With the health monitoring capabilities of the sensor, a computer feed back can designate a control input to ensure the proper configuration of the wing. Thus, deflections due to aerodynamic loads will be compensated for any flying condition.\nThe development of the distributed sensor for monitoring the LARGE deformation of morphing structures will be based on the work of Akl et al. (2006) which employed networks of distributed wire sensors to monitor SMALL amplitudes of vibration of beams and plates. For small deflections, the sensor relies in its operation on the linear theory of finite elements to extract the transverse linear and angular deflections. But for large deflections, the proposed sensor network will be based on the non-linear theory of finite elements to extract the transverse linear and angular deflections as well as the in-plane longitudinal deflections. The concept of the proposed distributed network sensor can best be understood by considering the onedimensional flexible beam-like structure shown in Figure (3).\nThe strain \u03b5(x) in a wire embedded inside the beam at a distance a, from its neutral axis, can be determined from\nthe following von Karman strain-displacement relationships:\n( )2 , , , 1 2x x xxu w aw\u03b5 = + \u2212 (1)\nFigure (3) shows also that the wire sensor is divided into N segments in order to extract N unknown degrees of freedom, i.e. nodal deflections, as will be explained latter.\nThe axial and transverse displacements, u and w, can be approximated as follows { }{ },uu N= \u2206 and { }{ }ww N= \u2206 (2)\nwhere { }uN and { }wN are the classical shape functions for axial and transverse displacements, and { }\u2206 is the nodal deflection vector defined as\n,{ } { }T xu w w\u2206 = (3)\nSubstituting equations (2) and (3) into equation (1) yields\nProc. of SPIE Vol. 6529 65290P-2\nDownloaded From: http://proceedings.spiedigitallibrary.org/ on 06/26/2016 Terms of Use: http://spiedigitallibrary.org/ss/TermsOfUse.aspx" + ] + }, + { + "image_filename": "designv11_69_0003403_fskd.2010.5569352-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003403_fskd.2010.5569352-Figure1-1.png", + "caption": "Figure 1. Simplified model of double inverted pendulum", + "texts": [ + " The composition rules are determined using composition coefficients which are derived from the LQR feedback controller parameters of double inverted pendulum. Experiment results of numeric simulation and real-time experimental testbed system show that the proposed method has better tracking performance, disturbance resisting performance, and robustness against model parameter perturbance. II. STATE VARIABLE COMPOSITION OF DOUBLE INVERTED PENDULUM MODEL Ignoring the air flow and friction, the real double inverted pendulum can be simplified to a model composed of a cart, two linked mass-well-distributed poles and a mass, as shown in Fig. 1. The nonlinear equations of motions can be derived using Euler-Lagrange\u2019s equation [1,3]. The state and output equation of the linearized model can be further derived by linearization at the equilibrium 1 2 1 2( , , , , , , ) (0,0,0,0,0,0,0)x x x\u03b8 \u03b8 \u03b8 \u03b8 = of the above nonlinear model (with detailed actual parameter values listed in appendix table 1) : 1 1 2 2 1 1 2 2 0 0 0 1 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 1 0 77.0642 -21.1927 0 0 0 5.7012 0 -38.5321 37.8186 0 0 0 -0.0728 x x u x x \u03b8 \u03b8 \u03b8 \u03b8 \u03b8 \u03b8 \u03b8 \u03b8 \u23a1 \u23a4 \u23a1 \u23a4\u23a1 \u23a4 \u23a1 \u23a4 \u23a2 \u23a5 \u23a2 \u23a5\u23a2 \u23a5 \u23a2 \u23a5 \u23a2 \u23a5 \u23a2 \u23a5\u23a2 \u23a5 \u23a2 \u23a5 \u23a2 \u23a5 \u23a2 \u23a5\u23a2 \u23a5 \u23a2 \u23a5 = +\u23a2 \u23a5 \u23a2 \u23a5\u23a2 \u23a5 \u23a2 \u23a5 \u23a2 \u23a5 \u23a2 \u23a5\u23a2 \u23a5 \u23a2 \u23a5 \u23a2 \u23a5 \u23a2 \u23a5\u23a2 \u23a5 \u23a2 \u23a5 \u23a2 \u23a5 \u23a2 \u23a5\u23a2 \u23a5 \u23a2 \u23a5 \u23a2 \u23a5 \u23a2 \u23a5\u23a2 \u23a5 \u23a2 \u23a5\u23a3 \u23a6 \u23a3 \u23a6\u23a3 \u23a6 \u23a3 \u23a6 (1) 1 2 1 2 1 2 1 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 1 0 0 0 0 x x u x \u03b8 \u03b8 \u03b8 \u03b8 \u03b8 \u03b8 \u23a1 \u23a4 \u23a2 \u23a5 \u23a2 \u23a5\u23a1 \u23a4 \u23a1 \u23a4 \u23a1 \u23a4\u23a2 \u23a5\u23a2 \u23a5 \u23a2 \u23a5 \u23a2 \u23a5= +\u23a2 \u23a5\u23a2 \u23a5 \u23a2 \u23a5 \u23a2 \u23a5\u23a2 \u23a5\u23a2 \u23a5 \u23a2 \u23a5 \u23a2 \u23a5\u23a3 \u23a6 \u23a3 \u23a6 \u23a3 \u23a6\u23a2 \u23a5 \u23a2 \u23a5 \u23a2 \u23a5\u23a3 \u23a6 (2) There are totally six output variables for the double inverted pendulum system: x , x , 1\u03b8 , 1\u03b8 , 2\u03b8 and 2\u03b8 " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000512_810105-Figure18-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000512_810105-Figure18-1.png", + "caption": "Fig. 18 - Contour maps of profile deviation", + "texts": [ + " The profiles of two set of hypoid gears, which are good and poor in respect to the axle noise respectively are compared in Fig. 17. We notice that these test results provide more valuable information than the tooth bearing test. For example, we see that with the poor set the lengthwise profiles are more inclined to each other and the crowning of the pinion is slightly larger. So it is seen that the conjugateness of the poor set is inferior to that of the good set. In addition to the lengthwise and trans verse profiles, the measuring method enable us to draw the profiles on the whole tooth surface, as shown in Fig. 18. In the case the measurement is made at the lattice points on the whole tooth surface and the contour map of the profile deviation is drawn from the data obtained. In Fig. 19 we see another example of the contour map of the relative profile deviation with the set of hypoid gears. This gives us an easy understanding of the characteristics of gearing. 810105 10 810105 CONCLUSION A new measuring method for the hypoid gear tooth profiles has been developed to obtain the lengthwise profiles, the transverse profiles and the profiles on the whole surface of a gear and a pinion" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002482_s1068798x08070058-Figure5-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002482_s1068798x08070058-Figure5-1.png", + "caption": "Fig. 5. Wear curve of channel: (1) ball profile; (2) initial channel profile; (3) final channel profile.", + "texts": [ + " To verify the proposed method, we calculate the dimensions ao and bo of the elliptical contact areas and the maximum contact pressures at the contact points A and B between the ball and the external and internal bearing rings when Po = 3660 N (\u2206co = 10 \u00b5m). For point A, by the proposed method, ao = 2.648 mm; bo = 0.312 mm; p2o = 2200 MPa; by the Hertz method, ao = 2.586 mm; bo = 0.295 mm; p2o = 2297 MPa. For 650 RUSSIAN ENGINEERING RESEARCH Vol. 28 No. 7 2008 PAVLOV point B, by the proposed method, ao = 2.45 mm; bo = 0.248 mm; p1o = 2740 MPa; by the Hertz method, ao = 2.652 mm; bo = 0.241 mm; p1o = 2692 MPa. A typical wear curve for the bearing-ring channel is shown in Fig. 5. We see that, in the races, a single central wear crater and two lateral wear craters are formed symmetrically with respect to the plane of bearing rotation. For a total radial gap in the bearing \u2206co = 10 \u00b5m and Po = 3660 N, the wear of the rings during a single contact with the central ball is \u2206h2o = 2.65 \u00d7 10\u20137 \u00b5m at point A and \u2206h1o = 2.89 \u00d7 10\u20137 \u00b5m at point B. The total wear of the internal and external bearing rings per rotation of the internal ring, with j2 = 4.1 contacts at point A of the outer ring under a load Po and with j1 = 2" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003745_15502280902939486-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003745_15502280902939486-Figure2-1.png", + "caption": "FIG. 2. Sphere profile at different stages.", + "texts": [ + " The current authors have also proposed [17] the results for the same pertaining to the materials of (300 < E/Y < 1000) group in the following way: pm Y = { 3.529 ( \u03c9 \u03c9c )2 + 198.3 ( \u03c9 \u03c9c ) + 90.68 } {( \u03c9 \u03c9c )2 + 83.26 ( \u03c9 \u03c9c ) + 173.9 } (3) A Ac = 0.93 ( \u03c9 \u03c9c )1.4112 (4) P Pc = 1.8 ( \u03c9 \u03c9c )1.214 \u2212 0.785 (5) During the second stage of unloading, the interference \u03c9 is gradually reduced. At the end, the contact load and area reduced to zero but the geometry is not fully regained because of the development of residual stresses. As a result, the unloaded sphere ended up with a deformed shape (Figure 2). However, this deformed shape is limited to a relatively narrow zone of the contact that prevails at the beginning of unloading. This deformed profile is characterized by the residual interference (\u03c9res) and residual radius (Rres). Obviously, these \u03c9res and Rres depend on maximum interference (\u03c9max) from where unloading started. As the radial displacements of the contacting points are negligible compared to the axial displacements, simulations are performed with perfect slip conditions at the contact without considering frictional forces", + " If the stress is gradually applied, as in the present case, resilience or the work done per unit volume (uy) in stressing to yield point is equal to average yield stress times the deformation (\u03b5). uy = 1 2 Y\u03b5 = 1 2 Y Y E = Y 2 2E (6) Hence low E/Y value indicates that the material is highly resilient, i.e. capable of absorbing more work or energy in the elastic range. The present results very well agree with the results and outcome predicted in [12] for materials with E/Y > 300. Hence this work aims to place emphasis only on the distinct unloading behavior of low E/Y materials. Figure 2 shows schematic representation of the geometry of the sphere in un-deformed, fully loaded and completely unloaded states. From this analysis it is found that the residual interference essentially depends on the level of intensity of stress and strain from where unloading started. This fact is shown in Figure 3, which presents the numerical results obtained from the present finite element analysis. Figure 3 relates residual interference in non-dimensional form (\u03c9res /\u03c9max) with the corresponding maximum interference (\u03c9max /\u03c9c) from where unloading started" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003786_cca.2010.5611063-Figure3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003786_cca.2010.5611063-Figure3-1.png", + "caption": "Fig. 3. The t", + "texts": [ + " SYSTEM DE in this paper al example of e system are t re related to t roximately lin d 2 (raising m as represent the same con g axis is the ng movemen e a tilt angle g environme the left part d in the right p P hardware is surely th engineering c RTW) provide diagram draw s provided by be integrat ment and e atform. The di in this paper nd testing proc llows: In S described. Th -methodology suitability will be show SCRIPTION is illustrated a Multi-Inpu he lift forces p he control vol ear manner. T axis) are con ed in Fig. 3. I trol power, a consequence. t (change of ( 3 ), represe ntal condition of the V-cycl art of the cycl and softwar e best known ommunity. Th s an automati n in Simulink the dSPACE ed in th xtending th fferent parts o in conjunction ess. ection II, th is is followed for the given of the Exact n based on in Fig. 2 and t Multi-Outpu rovided by th tage (actuating he two angle sidered as th f the two rotor change of th By a differen 1 ) can b nted in Fig. 4 s e e e e c . e e f e - t e s e s e t e , wo output angles ngle of the system the circular paths described by 1 and 2 " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001573_icmech.2006.252519-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001573_icmech.2006.252519-Figure2-1.png", + "caption": "Figure 2 Active leg of the ULB stiff Stewart platform \\ kix/", + "texts": [ + "he legs are mounted in such a way to achieve the geometry of cubic configuration. Each active leg consists I INTRODUCTION of a force sensor (B&K 8200), an amplified piezoelectric Vibration isolation is becoming more and more stringent as actuator (Cedrat Recherche APA50s) and two flexible the mechatronic systems are advancing and developing in joints as shown in Fig.2. In the ideal situation, the hexapod space and ground applications. In order to obtain high needs to be hinged using spherical joints, but to avoid the performance from the vibration isolator, the corner problem of friction and backlash, flexible tips are used frequency should be as low as possible [1, 2]. This leads the probem of frictiontanbclsh flexible tips are use researchers to reduce the frequency by the following ways: inta ofshrcljit.TeefeilXishv h following properties: zero friction, zero backlash, high axial stiffness and relatively low bending stiffness", + " Root locus of single axis piezoelectric isolator with PI feedback system can be used as an active strut for the previously mentioned Stewart platform. As an application for the From the analytical calculation, the intermediate frequency reduction, one can imagine adaptive structures displacement xa is: that can change their resonance frequency instantaneously sx + g(as + l)x to avoid being excited when the excitation frequency x sc +g(as+)d approaches a resonance. s + g(as + This schematic drawing represents the active leg of Stewart From the foregoing equations, one can calculate the platform shown in Fig.2 where the active feedback transmissibility FRF between the disturbance displacement controller is applied by acquiring the signal measured by and the payload displacement, which is equal to: the force sensor and feeding it back to the piezoelectric actuator after being filtered and compensated with PI compensator. The governing equation of motion for the Xc system in Laplace transform is: = Xd s2[(1+ ga)/CW2 ] + S[g/ C2 ] +l Ms2x = -ms2xd = k(x - x ) =F~d n MS2Xc =-MS2xd k(xd -Xa) F Where o,, is the natural frequency of the system", + " maimum reductionhas bee obaie by inrasn th 1~~~~~~~~~~~~~~gi of th prprtoa par of th cmpensator,butthi V0 EXERMETA VEIFCAO can lead to intailt if th surudn codtin chang proEportionlmluentegal ()cotrolrtreuehe ooatnhesaetmeoincese stability margi of th stiffnesissof he mstrcueas discuss d thoretially inther sysaatiem prvou redction. sms1iit rmsiuain eut 179 aeeto h dsubnesuc bd n hto h In the same context, another experiment has been done. The same control technique has been applied to the truss the force output from the collocated force sensor. The openstructure shown in Fig.12. The truss contains two active loop FRF (before stiffness reduction) shows that the two struts like the one shown in Fig.2 replacing two passive modes are located at 8.8 and 10.5 Hz. Using this control members. These two struts are used for the purpose of technique, they have been moved to 2.6 and 5 Hz, adding active damping to the system. respectively. A potential application of this is the adaptation The signals of the two force sensors, in the two active struts of structural resonances to a narrow band disturbance of of the truss, have been filtered using the (PI) compensator variable frequency. in a Digital Signal Processor (DSP) and fed back to the piezoelectric actuators" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002645_itsc.2008.4732593-Figure3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002645_itsc.2008.4732593-Figure3-1.png", + "caption": "Fig. 3. Suspension model", + "texts": [ + " In order to be able to use This work was supported and developed by the French LCPC laboratory (Laboratoire Central des Ponts et Chausse\u0301es) in collaboration with French industrial partners, Renault Trucks, Michelin and Sodit in the framework of French project VIF (Ve\u0301hicule Lourd Interactif du Future). sliding mode observer, the vehicle model is developed. Various studies have dealt with vehicle modelling ([6], [7], [8], [9]). In this section, a vehicle model with 5 degrees of freedom and composed of a car body, four suspensions and four wheels is developed. The suspension is modelled as the combination of spring and damper elements as shown in figure 3. The chassis (with the mass M ) is suspended on its axles through two suspension systems. We can show that the tire of the wheel i is modelled by only springs with coefficients ki and the suspension is modelled by both springs with 1-4244-2112-1/08/$20.00 \u00a92008 IEEE 523 coefficient Ki and damper elements Bi. The wheel masses are given by mi (i = 1, \u00b7 \u00b7 \u00b7 , 4). At the tire contact, the road profile represented by the variable ui (i = 1, \u00b7 \u00b7 \u00b7 , 4) is considered as a heavy vehicle input. At the tire, the road profile, longitudinal and lateral slope, skid resistance and radius of curvature represent road data inputs of this model" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000410_bf02326646-Figure5-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000410_bf02326646-Figure5-1.png", + "caption": "Fig. 5--Strain in transition joint on heating from 20-560~ computed by a finite-element analysis", + "texts": [ + " These strain componen ts for a t empera ture change from 20-560~ are shown in Fig. 4, where it can be seen that the interface affects the total strain for about 30 mm on either side. The interesting features are the peaks in the strain perpendicular to the interface a few millimeters each side of the interface. There is some indicat ion of a similar d is turbance in the strain parallel to the interface. These strains can be compared with those obta ined from a finite-element analysis of the plate (Fig. 5). There is some similarity between the results, but the discont inui ty at the interface is p robably an artefact produced by the finite-element model. The moir ( measurements conf i rm that there is a dis turbance near the interface, but that , as one might expect, there is no sudden discontinuity. It should, of course, be remembered that the measure- ments are of total strain: thermal , elastic, plastic and creep. The thermal componen t is by far the largest, account ing for more than 85 percent of the total " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002979_978-3-540-70534-5_7-Figure13.3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002979_978-3-540-70534-5_7-Figure13.3-1.png", + "caption": "Figure 13.3: Mako design [Gonzalez 2004]", + "texts": [], + "surrounding_texts": [ + "[Br\u00e4unl 2000], [Br\u00e4unl, Sutherland, Unkelbach 2002] Classic PID control is used to control the robot\u2019s leaning front/back and left/right, similar to the case of static balance. However, here we do not intend to make the robot stand up straight. Instead, in a teaching stage, we record the desired front and side lean of the robot\u2019s body during all phases of its gait. Later, when controlling the walking gait, we try to achieve this offset of front and side lean by using a PID controller. The following parameters can be set in a standard walking gate to achieve this leaning: \u2022 Step length \u2022 Height of leg lift \u2022 Walking speed \u2022 Amount of forward lean of torso \u2022 Maximal amount of side swing 6. Fuzzy control [Unpublished] We are working on an adaptation of the PID control, replacing the classic PID control by fuzzy logic for dynamic balance. 7. Artificial horizon [Wicke 2001] This innovative approach does not use any of the kinetics sensors of the other approaches, but a monocular grayscale camera. In the simple version, a black line on white ground (an \u201cartificial horizon\u201d) is placed in the visual field of the robot. We can then measure the robot\u2019s orientation by changes of the line\u2019s position and orientation in the image. For example, the line will move to the top if the robot is falling forward, it will be slanted at an angle if the robot is leaning left, and so on (Figure 11.8). With a more powerful controller for image processing, the same principle can be applied even without the need for an artificial horizon. As long as there is enough texture in the background, general optical flow can be used to determine the robot\u2019s movements. Figure 11.9 shows Johnny Walker during a walking cycle. Note the typical side-swing of the torso to counterbalance the leg-lifting movement. This creates a large momentum around the robot\u2019s center of mass, which can cause problems with stability due to the limited accuracy of the servos used as actuators. Dynamic Balance 179 Figure 11.10 shows a similar walking sequence with Andy Droid. Here, the robot performs a much smoother and better controlled walking gait, since the mechanical design of the hip area allows a smoother shift of weight toward the side than in Johnny\u2019s case. 11.5.2 Alternative Biped Designs All the biped robots we have discussed so far are using servos as actuators. This allows an efficient mechanical and electronic design of a robot and therefore is a frequent design approach in many research groups, as can be seen from the group photo of FIRA HuroSot World Cup Competition in 2002 [Baltes, Br\u00e4unl 2002]. With the exception of one robot, all robots were using servos (see Figure 11.11). Other biped robot designs also using the EyeCon controller are Tao Pie Pie from University of Auckland, New Zealand, and University of Manitoba, Canada, [Lam, Baltes 2002] and ZORC from Universit\u00e4t Dortmund, Germany [Ziegler et al. 2001]. Walking Robots 180 11 As has been mentioned before, servos have severe disadvantages for a number of reasons, most importantly because of their lack of external feedback. The construction of a biped robot with DC motors, encoders, and endswitches, however, is much more expensive, requires additional motor driver electronics, and is considerably more demanding in software development. So instead of redesigning a biped robot by replacing servos with DC motors and keeping the same number of degrees of freedom, we decided to go for a minimal approach. Although Andy has 10 dof in both legs, it utilizes only three independent dof: bending each leg up and down, and leaning the whole body left or right. Therefore, it should be possible to build a robot that uses only three motors and uses mechanical gears or pulleys to achieve the articulated joint motion. 181 The CAD designs following this approach and the finished robot are shown in Figure 11.12 [Jungpakdee 2002]. Each leg is driven by only one motor, while the mechanical arrangement lets the foot perform an ellipsoid curve for each motor revolution. The feet are only point contacts, so the robot has to keep moving continuously, in order to maintain dynamic balance. Only one motor is used for shifting a counterweight in the robot\u2019s torso sideways (the original drawing in Figure 11.12 specified two motors). Figure 11.13 shows the simulation of a dynamic walking sequence [Jungpakdee 2002]. Walking Robots 182 11 11.6 References BALTES. J., BR\u00c4UNL, T. HuroSot - Laws of the Game, FIRA 1st Humanoid Ro- bot Soccer Workshop (HuroSot), Daejeon Korea, Jan. 2002, pp. 43-68 (26) BOEING, A., BR\u00c4UNL, T. Evolving Splines: An alternative locomotion controller for a bipedal robot, Seventh International Conference on Control, Automation, Robotics and Vision, ICARV 2002, CD-ROM, Singapore, Dec. 2002, pp. 1-5 (5) BOEING, A., BR\u00c4UNL, T. Evolving a Controller for Bipedal Locomotion, Proceedings of the Second International Symposium on Autonomous Minirobots for Research and Edutainment, AMiRE 2003, Brisbane, Feb. 2003, pp. 43-52 (10) BR\u00c4UNL, T. Design of Low-Cost Android Robots, Proceedings of the First IEEE-RAS International Conference on Humanoid Robots, Humanoids 2000, MIT, Boston, Sept. 2000, pp. 1-6 (6) BR\u00c4UNL, T., SUTHERLAND, A., UNKELBACH, A. Dynamic Balancing of a Humanoid Robot, FIRA 1st Humanoid Robot Soccer Workshop (HuroSot), Daejeon Korea, Jan. 2002, pp. 19-23 (5) CAUX, S., MATEO, E., ZAPATA, R. Balance of biped robots: special double-inverted pendulum, IEEE International Conference on Systems, Man, and Cybernetics, 1998, pp. 3691-3696 (6) CHO, H., LEE, J.-J. (Eds.) Proceedings 2002 FIRA Robot World Congress, Seoul, Korea, May 2002 DOERSCHUK, P., NGUYEN, V., LI, A. Neural network control of a three-link leg, in Proceedings of the International Conference on Tools with Artificial Intelligence, 1995, pp. 278-281 (4) FUJIMOTO, Y., KAWAMURA, A. Simulation of an autonomous biped walking robot including environmental force interaction, IEEE Robotics and Automation Magazine, June 1998, pp. 33-42 (10) GODDARD, R., ZHENG, Y., HEMAMI, H. Control of the heel-off to toe-off motion of a dynamic biped gait, IEEE Transactions on Systems, Man, and Cybernetics, vol. 22, no. 1, 1992, pp. 92-102 (11) HARADA, H. Andy-2 Visualization Video, http://robotics.ee.uwa.edu.au /eyebot/mpg/walk-2leg/, 2006 JUNGPAKDEE, K., Design and construction of a minimal biped walking mechanism, B.E. Honours Thesis, The Univ. of Western Australia, Dept. of Mechanical Eng., supervised by T. Br\u00e4unl and K. Miller, 2002 KAJITA, S., YAMAURA, T., KOBAYASHI, A. Dynamic walking control of a biped robot along a potential energy conserving orbit, IEEE Transactions on Robotics and Automation, Aug. 1992, pp. 431-438 (8) References 183 KUN, A., MILLER III, W. Adaptive dynamic balance of a biped using neural networks, in Proceedings of the 1996 IEEE International Conference on Robotics and Automation, Apr. 1996, pp. 240-245 (6) LAM, P., BALTES, J. Development of Walking Gaits for a Small Humanoid Robot, Proceedings 2002 FIRA Robot World Congress, Seoul, Korea, May 2002, pp. 694-697 (4) MILLER III, W. Real-time neural network control of a biped walking robot, IEEE Control Systems, Feb. 1994, pp. 41-48 (8) MONTGOMERY, G. Robo Crop - Inside our AI Labs, Australian Personal Computer, Issue 274, Oct. 2001, pp. 80-92 (13) NICHOLLS, E. Bipedal Dynamic Walking in Robotics, B.E. Honours Thesis, The Univ. of Western Australia, Electrical and Computer Eng., supervised by T. Br\u00e4unl, 1998 PARK, J.H., KIM, K.D. Bipedal Robot Walking Using Gravity-Compensated Inverted Pendulum Mode and Computed Torque Control, IEEE International Conference on Robotics and Automation, 1998, pp. 3528-3533 (6) R\u00dcCKERT, U., SITTE, J., WITKOWSKI, U. (Eds.) Autonomous Minirobots for Research and Edutainment \u2013 AMiRE2001, Proceedings of the 5th International Heinz Nixdorf Symposium, HNI-Verlagsschriftenreihe, no. 97, Univ. Paderborn, Oct. 2001 SUTHERLAND, A., BR\u00c4UNL, T. Learning to Balance an Unknown System, Proceedings of the IEEE-RAS International Conference on Humanoid Robots, Humanoids 2001, Waseda University, Tokyo, Nov. 2001, pp. 385-391 (7) TAKANISHI, A., ISHIDA, M., YAMAZAKI, Y., KATO, I. The realization of dynamic walking by the biped walking robot WL-10RD, in ICAR\u201985, 1985, pp. 459-466 (8) UNKELBACH, A. Analysis of sensor data for balancing and walking of a biped robot, Project Thesis, Univ. Kaiserslautern / The Univ. of Western Australia, supervised by T. Br\u00e4unl and D. Henrich, 2002 WICKE, M. Bipedal Walking, Project Thesis, Univ. Kaiserslautern / The Univ. of Western Australia, supervised by T. Br\u00e4unl, M. Kasper, and E. von Puttkamer, 2001 ZIEGLER, J., WOLFF, K., NORDIN, P., BANZHAF, W. Constructing a Small Humanoid Walking Robot as a Platform for the Genetic Evolution of Walking, Proceedings of the 5th International Heinz Nixdorf Symposium, Autonomous Minirobots for Research and Edutainment, AMiRE 2001, HNI-Verlagsschriftenreihe, no. 97, Univ. Paderborn, Oct. 2001, pp. 51-59 (9) Walking Robots 184 11 ZIMMERMANN, J., Balancing of a Biped Robot using Force Feedback, Diploma Thesis, FH Koblenz / The Univ. of Western Australia, supervised by T. Br\u00e4unl, 2004 185185 . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . AUTONOMOUS PLANES uilding an autonomous model airplane is a considerably more difficult undertaking than the previously described autonomous driving or walking robots. Model planes or helicopters require a significantly higher level of safety, not only because the model plane with its expensive equipment might be lost, but more importantly to prevent endangering people on the ground. A number of autonomous planes or UAVs (Unmanned Aerial Vehicles) have been built in the past for surveillance tasks, for example Aerosonde [Aerosonde 2006]. These projects usually have multi-million-dollar budgets, which cannot be compared to the smaller-scale projects shown here. Two projects with similar scale and scope to the one presented here are \u201cMicroPilot\u201d [MicroPilot 2006], a commercial hobbyist system for model planes, and \u201cFireMite\u201d [Hennessey 2002], an autonomous model plane designed for competing in the International Aerial Robotics Competition [AUVS 2006]. 12.1 Application Low-budget autopilot Our goal was to modify a remote controlled model airplane for autonomous flying to a given sequence of waypoints (autopilot). \u2022 The plane takes off under remote control. \u2022 Once in the air, the plane is switched to autopilot and flies to a previ- ously recorded sequence of waypoints using GPS (global positioning system) data. \u2022 The plane is switched back to remote control and landed. So the most difficult tasks of take-off and landing are handled by a pilot using the remote control. The plane requires an embedded controller to interface to the GPS and additional sensors and to generate output driving the servos. There are basically two design options for constructing an autopilot system for such a project (see Figure 12.1): Autonomous Planes 186 12 A. The embedded controller drives the plane\u2019s servos at all times. It receives sensor input as well as input from the ground transmitter. B. A central (and remote controlled) multiplexer switches between ground transmitter control and autopilot control of the plane\u2019s servos. Design option A is the simpler and more direct solution. The controller reads data from its sensors including the GPS and the plane\u2019s receiver. Ground control can switch between autopilot and manual operation by a separate channel. The controller is at all times connected to the plane\u2019s servos and generates their PWM control signals. However, when in manual mode, the controller reads the receiver\u2019s servo output and regenerates identical signals. Design option B requires a four-way multiplexer as an additional hardware component. (Design A has a similar multiplexer implemented in software.) The multiplexer connects either the controller\u2019s four servo outputs or the receiver\u2019s four servo outputs to the plane\u2019s servos. A special receiver channel is used for toggling the multiplexer state under remote control. Although design A is the superior solution in principle, it requires that the controller operates with highest reliability. Any fault in either controller hardware or controller software, for example the \u201changing\u201d of an application pro- A GPS 5 5 4 Controller Receiver Application 187 gram, will lead to the immediate loss of all control surfaces and therefore the loss of the plane. For this reason we opted to implement design B. Although it requires a custom-built multiplexer as additional hardware, this is a rather simple electro-magnetic device that can be directly operated via remote control and is not subject to possible software faults. Figure 12.2 shows photos of the construction and during flight of our first autonomous plane. This plane had the EyeCon controller and the multiplexer unit mounted on opposite sides of the fuselage. Autonomous Planes 188 12 12.2 Control System and Sensors Black box An EyeCon system is used as on-board flight controller. Before take-off, GPS waypoints for the desired flight path are downloaded to the controller. After the landing, flight data from all on-board sensors is uploaded, similar to the operation of a \u201cblack box\u201d data recorder on a real plane. The EyeCon\u2019s timing processor outputs generate PWM signals that can directly drive servos. In this application, they are one set of inputs for the multiplexer, while the second set of inputs comes from the plane\u2019s receiver. Two serial ports are used on the EyeCon, one for upload/download of data and programs, and one for continuous GPS data input. Although the GPS is the main sensor for autonomous flight, it is insufficient because it delivers a very slow update of 0.5Hz .. 1.0Hz and it cannot determine the plane\u2019s orientation. We are therefore experimenting with a number of additional sensors (see Chapter 3 for details of these sensors): \u2022 Digital compass Although the GPS gives directional data, its update rates are insufficient when flying in a curve around a waypoint. \u2022 Piezo gyroscope and inclinometer Gyroscopes give the rate of change, while inclinometers return the absolute orientation. The combined use of both sensor types helps reduce the problems with each individual sensor. \u2022 Altimeter and air-speed sensor Both altimeter and air-speed sensor have been built by using air pressure sensors. These sensors need to be calibrated for different heights and temperatures. The construction of an air-speed sensor requires the combination of two sensors measuring the air pressure at the wing tip with a so-called \u201cPitot tube\u201d, and comparing the result with a third air pressure sensor inside the fuselage, which can double as a height sensor. Figure 12.3 shows the \u201cEyeBox\u201d, which contains most equipment required for autonomous flight, EyeCon controller, multiplexer unit, and rechargeable battery, but none of the sensors. The box itself is an important component, since it is made out of lead-alloy and is required to shield the plane\u2019s receiver from any radiation from either the controller or the multiplexer. Since the standard radio control carrier frequency of 35MHz is in the same range as the EyeCon\u2019s operating speed, shielding is essential. Another consequence of the decision for design B is that the plane has to remain within remote control range. If the plane was to leave this range, unpredictable switching between the multiplexer inputs would occur, switching control of the plane back and forth between the correct autopilot settings and noise signals. A similar problem would exist for design A as well; however, the controller could use plausibility checks to distinguish noise from proper remote control signals. By effectively determining transmitter strength, the controller could fly the plane even outside the remote transmitter\u2019s range. Flight Program Autonomous Planes 190 12 Program 12.1 shows sample NMEA output. After the start-up sequence, we get regular code strings for position and time, but we only decode the lines starting with $GPGGA. In the beginning, the GPS has not yet logged on to a sufficient number of satellites, so it still reports the geographical position as 0 N and 0 E. The quality indicator in the sixth position (following \u201cE\u201d) is 0, so the coordinates are invalid. In the second part of Program 12.1, the $GPRMC string has quality indicator 1 and the proper coordinates of Western Australia. In our current flight system we are using approach A, to be more flexible in flight path generation. For the first implementation, we are only switching the rudder between autopilot and remote control, not all of the plane\u2019s surfaces. Motor and elevator stay on remote control for safety reasons, while the ailerons are automatically operated by a gyroscope to eliminate any roll. Turns under autopilot therefore have to be completely flown using the rudder, which requires a much larger radius than turns using ailerons and elevator. The remaining control surfaces of the plane can be added step by step to the autopilot system. The flight controller has to perform a number of tasks, which need to be accessible through its user interface: Program 12.1: NMEA sample output $TOW: 0 $WK: 1151 $POS: 6378137 0 0 $CLK: 96000 $CHNL:12 $Baud rate: 4800 System clock: 12.277MHz $HW Type: S2AR $GPGGA,235948.000,0000.0000,N,00000.0000,E,0,00,50.0,0.0,M,,,,0000*3A $GPGSA,A,1,,,,,,,,,,,,,50.0,50.0,50.0*05 $GPRMC,235948.000,V,0000.0000,N,00000.0000,E,,,260102,,*12 $GPGGA,235948.999,0000.0000,N,00000.0000,E,0,00,50.0,0.0,M,,,,0000*33 $GPGSA,A,1,,,,,,,,,,,,,50.0,50.0,50.0*05 $GPRMC,235948.999,V,0000.0000,N,00000.0000,E,,,260102,,*1B $GPGGA,235949.999,0000.0000,N,00000.0000,E,0,00,50.0,0.0,M,,,,0000*32 $GPGSA,A,1,,,,,,,,,,,,,50.0,50.0,50.0*05 ... $GPRMC,071540.282,A,3152.6047,S,11554.2536,E,0.49,201.69,171202,,*11 $GPGGA,071541.282,3152.6044,S,11554.2536,E,1,04,5.5,3.7,M,,,,0000*19 $GPGSA,A,2,20,01,25,13,,,,,,,,,6.0,5.5,2.5*34 $GPRMC,071541.282,A,3152.6044,S,11554.2536,E,0.53,196.76,171202,,*1B $GPGGA,071542.282,3152.6046,S,11554.2535,E,1,04,5.5,3.2,M,,,,0000*1E $GPGSA,A,2,20,01,25,13,,,,,,,,,6.0,5.5,2.5*34 $GPRMC,071542.282,A,3152.6046,S,11554.2535,E,0.37,197.32,171202,,*1A $GPGGA,071543.282,3152.6050,S,11554.2534,E,1,04,5.5,3.3,M,,,,0000*18 $GPGSA,A,2,20,01,25,13,,,,,,,,,6.0,5.5,2.5*34 $GPGSV,3,1,10,01,67,190,42,20,62,128,42,13,45,270,41,04,38,228,*7B $GPGSV,3,2,10,11,38,008,,29,34,135,,27,18,339,,25,13,138,37*7F $GPGSV,3,3,10,22,10,095,,07,07,254,*76 Flight Program 191 Pre-flight \u2022 Initialize and test all sensors, calibrate sensors. \u2022 Initialize and test all servos, enable setting of zero positions of servos, enable setting of maximum angles of servos. \u2022 Perform waypoint download \u2013 (only for technique A). In-flight (continuous loop) \u2022 Generate desired heading \u2013 (only for technique A). \u2022 Set plane servos according to desired heading. \u2022 Record flight data from sensors. Post-flight \u2022 Perform flight data upload. These tasks and settings can be activated by navigating through several flight system menus, as shown in Figure 12.4 [Hines 2001]. They can be displayed and operated through button presses either directly on the EyeCon or Autonomous Planes 192 12 remotely via a serial link cable on a PDA (Personal Digital Assistant, for example Compaq IPAQ). The link between the EyeCon and the PDA has been developed to be able to remote-control (via cable) the flight controller pre-flight and post-flight, especially to download waypoints before take-off and upload flight data after landing. All pre-start diagnostics, for example correct operation of all sensors or the satellite log-on of the GPS, are transmitted from the EyeCon to the handheld PDA screen. After completion of the flight, all sensor data from the flight together with time stamps are uploaded from the EyeCon to the PDA and can be graphically displayed. Figure 12.5 [Purdie 2002] shows an example of an uploaded flight path ([x, y] coordinates from the GPS sensor); however, all other sensor data is being logged as well for post-flight analysis. A desirable extension to this setup is the inclusion of wireless data transmission from the plane to the ground (see also Chapter 7). This would allow us to receive instantaneous data from the plane\u2019s sensors and the controller\u2019s status as opposed to doing a post-flight analysis. However, because of interference problems with the other autopilot components, wireless data transmission has been left until a later stage of the project. 12.4 References AEROSONDE, Global Robotic Observation System Aerosonde, http://www. aerosonde.com, 2006 AUVS, International Aerial Robotics Competition, Association for Unmanned Vehicle Systems, http://avdil.gtri.gatech.edu/AUVS/IARC LaunchPoint.html, 2006 HENNESSEY, G. The FireMite Project, http://www.craighennessey.com/ firemite/, May 2002 References 193 HINES, N. Autonomous Plane Project 2001 \u2013 Compass, GPS & Logging Subsystems, B.E. Honours Thesis, The University of Western Australia, Electrical and Computer Eng., supervised by T. Br\u00e4unl and C. Croft, 2001 MAGELLAN, GPS 315/320 User Manual, Magellan Satellite Access Technology, San Dimas CA, 1999 MICROPILOT, MicroPilot UAV Autopilots, http://www.micropilot.com, 2006 PURDIE, J. Autonomous Flight Control for Radio Controlled Aircraft, B.E. Honours Thesis, The University of Western Australia, Electrical and Computer Eng., supervised by T. Br\u00e4unl and C. Croft, 2002 ROJONE, MicroGenius 3 User Manual, Rojone Pty. Ltd. CD-ROM, Sydney Australia, 2002 195195 13AUTONOMOUS VESSELS . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . AND UNDERWATER VEHICLES he design of an autonomous vessel or underwater vehicle requires one additional skill compared to the robot designs discussed previously: watertightness. This is a challenge especially for autonomous under- water vehicles (AUVs), as they have to cope with increasing water pressure when diving and they require watertight connections to actuators and sensors outside the AUV\u2019s hull. In this chapter, we will concentrate on AUVs, since autonomous vessels or boats can be seen as AUVs without the diving functionality. The area of AUVs looks very promising to advance commercially, given the current boom of the resource industry combined with the immense cost of either manned or remotely operated underwater missions. 13.1 Application Unlike many other areas of mobile robots, AUVs have an immediate application area conducting various sub-sea surveillance and manipulation tasks for the resource industry. In the following, we want to concentrate on intelligent control and not on general engineering tasks such as constructing AUVs that can go to great depths, as there are industrial ROV (remotely operated vehicle) solutions available that have solved these problems. While most autonomous mobile robot applications can also use wireless communication to a host station, this is a lot harder for an AUV. Once submerged, none of the standard communication methods work; Bluetooth or WLAN only operate up to a water depth of about 50cm. The only wireless communication method available is sonar with a very low data rate, but unfortunately these systems have been designed for the open ocean and can usually not cope with signal reflections as they occur when using them in a pool. So unless some wire-bound communication method is used, AUV applications have to be truly autonomous. T Autonomous Vessels and Underwater Vehicles 196 13 AUV Competition The Association for Unmanned Vehicles International (AUVSI) organizes annual competitions for autonomous aerial vehicles and for autonomous underwater vehicles [AUVSI 2006]. Unfortunately, the tasks are very demanding, so it is difficult for new research groups to enter. Therefore, we decided to develop a set of simplified tasks, which could be used for a regional or entrylevel AUV competition (Figure 13.1). We further developed the AUV simulation system SubSim (see Section 15.6), which allows to design AUVs and implement control programs for the individual tasks without having to build a physical AUV. This simulation system could serve as the platform for a simulation track of an AUV competition. The four suggested tasks to be completed in an olympic size swimming pool are: 1. Wall Following The AUV is placed close to a corner of the pool and has to follow the pool wall without touching it. The AUV should perform one lap around the pool, return to the starting position, then stop. 2. Pipeline Following A plastic pipe is placed along the bottom of the pool, starting on one side of the pool and terminating on the opposite side. The pipe is made out of straight pieces and 90 degree angles. The AUV is placed over the start of the pipe on one side of the pool and has to follow the pipe on the ground until the opposite wall has been reached. Dynamic Model 197 3. Target Finding The AUV has to locate a target plate with a distinctive texture that is placed at a random position within a 3m diameter from the center of the pool. 4. Object Mapping A number of simple objects (balls or boxes of distinctive color) are placed at the bottom of the pool, distributed over the whole pool area. The AUV has to survey the whole pool area, e.g. by diving along a sweeping pattern, and record all objects found at the bottom of the pool. Finally, the AUV has to return to its start corner and upload the coordinates of all objects found. 13.2 Dynamic Model The dynamic model of an AUV describes the AUV\u2019s motions as a result of its shape, mass distribution, forces/torques exerted by the AUV\u2019s motors, and external forces/torques (e.g. ocean currents). Since we are operating at relatively low speeds, we can disregarding the Coriolis force and present a simplified dynamic model [Gonzalez 2004]: with: M mass and inertia matrix v linear and angular velocity vector D hydrodynamic damping matrix G gravitational and buoyancy vector force and torque vector (AUV motors and eternal forces/torques) D can be further simplified as a diagonal matrix with zero entries for y (AUV can only move forward/backward along x, and dive/surface along z, but not move sideways), and zero entries for rotations about x and y (AUV can actively rotate only about z, while its self-righting movement, see Section 13.3, greatly eliminates rotations about x and y). G is non-zero only in its z component, which is the sum of the AUV\u2019s gravity and buoyancy vectors. is the product of the force vector combining all of an AUV\u2019s motors, with a pose matrix that defines each motor\u2019s position and orientation based on the AUV\u2019s local coordinate system. 13.3 AUV Design Mako The Mako (Figure 13.2) was designed from scratch as a dual PVC hull containing all electronics and batteries, linked by an aluminum frame and propelled by 4 trolling motors, 2 of which are for active diving. The advantages of this design over competing proposals are [Br\u00e4unl et al. 2004], [Gonzalez 2004]: M v\u00b7 D v( ) v G\u2022\u22c5\u2022\u22c5 \u03c4\u2022+ + = \u03c4 \u03c4 Autonomous Vessels and Underwater Vehicles 198 13 \u2022 Ease in machining and construction due to its simple structure \u2022 Relative ease in ensuring watertight integrity because of the lack of rotating mechanical devices such as bow planes and rudders \u2022 Substantial internal space owing to the existence of two hulls \u2022 High modularity due to the relative ease with which components can be attached to the skeletal frame \u2022 Cost-effectiveness because of the availability and use of common materials and components \u2022 Relative ease in software control implementation when compared to using a ballast tank and single thruster system \u2022 Ease in submerging with two vertical thrusters \u2022 Static stability due to the separation of the centers of mass and buoy- ancy, and dynamic stability due to the alignment of thrusters Simplicity and modularity were key goals in both the mechanical and electrical system designs. With the vehicle not intended for use below 5m depth, pressure did not pose a major problem. The Mako AUV measures 1.34 m long, 64.5 cm wide and 46 cm tall. The vehicle comprises two watertight PVC hulls mounted to a supporting aluminum skeletal frame. Two thrusters are mounted on the port and starboard sides of the vehicle for longitudinal movement, while two others are mounted vertically on the bow and stern for depth control. The Mako\u2019s vertical thruster diving system is not power conservative, however, when a comparison is made with ballast systems that involve complex mechanical devices, the advantages such as precision and simplicity that comes with using these two thrusters far outweighs those of a ballast system. AUV Design Mako 199 Propulsion is provided by four modified 12V, 7A trolling motors that allow horizontal and vertical movement of the vehicle. These motors were chosen for their small size and the fact that they are intended for underwater use; a feature that minimized construction complexity substantially and provided watertight integrity. Autonomous Vessels and Underwater Vehicles 200 13 The starboard and port motors provide both forward and reverse movement while the stern and bow motors provide depth control in both downward and upward directions. Roll is passively controlled by the vehicle\u2019s innate righting moment (Figure 13.5). The top hull contains mostly air besides light electronics equipment, the bottom hull contains heavy batteries. Therefore mainly a buoyancy force pulls the top cylinder up and gravity pulls the bottom cylinder down. If for whatever reason, the AUV rolls as in Figure 13.5, right, these two forces ensure that the AUV will right itself. Overall, this provides the vehicle with 4DOF that can be actively controlled. These 4DOF provide an ample range of motion suited to accomplishing a wide range of tasks. Controllers The control system of the Mako is separated into two controllers; an EyeBot microcontroller and a mini-PC. The EyeBot\u2019s purpose is controlling the AUV\u2019s movement through its four thrusters and its sensors. It can run a completely autonomous mission without the secondary controller. The mini PC is a Cyrix 233MHz processor, 32Mb of RAM and a 5GB hard drive, running Linux. Its sole function is to provide processing power for the computationally intensive vision system. Motor controllers designed and built specifically for the thrusters provide both speed and direction control. Each motor controller interfaces with the EyeBot controller via two servo ports. Due to the high current used by the thrusters, each motor controller produces a large amount of heat. To keep the temperature inside the hull from rising too high and damaging electronic components, a heat sink attached to the motor controller circuit on the outer hull was devised. Hence, the water continuously cools the heat sink and allows the temperature inside the hull to remain at an acceptable level. Sensors The sonar/navigation system utilizes an array of Navman Depth2100 echo sounders, operating at 200 kHz. One of these sensors is facing down and thereby providing an effective depth sensor (assuming the pool depth is known), while the other three sensors are used as distance sensors pointing forward, left, and right. An auxiliary control board, based on a PIC controller, has been designed to multiplex the four sonars and connect to the EyeBot [Alfirevich 2005]. AUV Design USAL 201 A low-cost Vector 2X digital magnetic compass module provides for yaw or heading control. A simple dual water detector circuit connected to analogue-to-digital converter (ADC) channels on the EyeBot controller is used to detect a possible hull breach. Two probes run along the bottom of each hull, which allows for the location (upper or lower hull) of the leak to be known. The EyeBot periodically monitors whether or not the hull integrity of the vehicle has been compromised, and if so immediately surfaces the vehicle. Another ADC input of the EyeBot is used for a power monitor that will ensure that the system voltage remains at an acceptable level. Figure 13.6 shows the Mako in operation. 13.4 AUV Design USAL The USAL AUV uses a commercial ROV as a basis, which was heavily modified and extended (Figure 13.7). All original electronics were taken out and replaced by an EyeBot controller (Figure 13.8). The hull was split and extended by a trolling motor for active diving, which allows the AUV to hover, while the original ROV had to use active rudder control during a forward motion for diving, [Gerl 2006], [Drtil 2006]. Figure 13.8 shows USAL\u2019s complete electronics subsystem. For simplicity and cost reasons, we decided to trial infrared PSD sensors (see Section 3.6) for the USAL instead of the echo sounders used on the Mako. Since the front part of the hull was made out of clear perspex, we were able to place the PSD sensors inside the AUV hull, so we did not have to worry about waterproofing sensors and cabling. Figure 13.9 shows the results of measurements conducted in [Drtil 2006], using this sensor setup in air (through the hull), and in different grades of water quality. Assuming good water quality, as can be expected in a swimming pool, the sensor setup returns reliable results up to a distance of about 1.1 m, which is sufficient for using it as a collision avoidance sensor, but too short for using it as a navigation aid in a large pool. Autonomous Vessels and Underwater Vehicles 202 13 AUV Design USAL 203 The USAL system overview is shown in Figure 13.10. Numerous sensors are connected to the EyeBot on-board controller. These include a digital camera, four analog PSD infrared distance sensors, a digital compass, a three-axes solid state accelerometer and a depth pressure sensor. The Bluetooth wireless communication system can only be used when the AUV has surfaced or is diving close to the surface. The energy control subsystem contains voltage regulators and level converters, additional voltage and leakage sensors, as well as motor drivers for the stern main driving motor, the rudder servo, the diving trolling motor, and the bow thruster pump. Figure 13.11 shows the arrangement of the three thrusters and the stern rudder, together with a typical turning movement of the USAL. Autonomous Vessels and Underwater Vehicles 204 13 13.5 References ALFIREVICH, E. Depth and Position Sensing for an Autonomous Underwater Vehicle, B.E. Honours Thesis, The University of Western Australia, Electrical and Computer Eng., supervised by T. Br\u00e4unl, 2005 AUVSI, AUVSI and ONR's 9th International Autonomous Underwater Vehicle Competition, Association for Unmanned Vehicle Systems International, http://www.auvsi.org/competitions/water.cfm, 2006 BR\u00c4UNL, T., BOEING, A., GONZALES, L., KOESTLER, A., NGUYEN, M., PETITT, J. The Autonomous Underwater Vehicle Initiative \u2013 Project Mako, 2004 IEEE Conference on Robotics, Automation, and Mechatronics (IEEE-RAM), Dec. 2004, Singapore, pp. 446-451 (6) DRTIL, M. Electronics and Sensor Design of an Autonomous Underwater Vehicle, Diploma Thesis, The University of Western Australia and FH Koblenz, Electrical and Computer Eng., supervised by T. Br\u00e4unl, 2006 GERL, B. Development of an Autonomous Underwater Vehicle in an Interdisciplinary Context, Diploma Thesis, The University of Western Australia and Technical Univ. M\u00fcnchen, Electrical and Computer Eng., supervised by T. Br\u00e4unl, 2006 GONZALEZ, L. Design, Modelling and Control of an Autonomous Underwater Vehicle, B.E. Honours Thesis, The University of Western Australia, Electrical and Computer Eng., supervised by T. Br\u00e4unl, 2004 205205 ROBOT . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . MANIPULATORS he main focus on the robotics side of this book is on autonomous mobile robots. However, we also want to give a brief introduction to the area of stationary manipulators, as they still form the vast majority of all commercial robot systems. Traditional applications of robot manipulators are spot welding and spray painting (Figure 14.1, Figure 14.2), e.g., in the automotive industry, as well as packaging and filling tasks, e.g. in the chemical and pharmaceutical industry. Robot manipulators can work in hazardous environments (e.g., nuclear power plants), and can conduct a variety of tasks from simple repetitive movements to complex sensor-based assemblies. Robot Manipulators 206 14 14.1 Homogeneous Coordinates Any design work involving robot manipulators is based on kinematics. Homogeneous coordinates are a necessary prerequisite for this." + ] + }, + { + "image_filename": "designv11_69_0000694_b:tels.0000029042.75697.f0-Figure11-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000694_b:tels.0000029042.75697.f0-Figure11-1.png", + "caption": "Figure 11. Scanning mirror rotates around zsr axis. The look-down angle \u03b8sr t , of the radar beam is fixed. Analysis of the FFT power spectrum reveals the target range.", + "texts": [ + " MMW radar front end and horn-lens antenna. The transmitter and receiver in the MMW radar front end are configured to operate at 77 GHz with a transmit power of 10 mW. Equation (23) is an empirical formula that relates 3 dB beamwidth to antenna diameter D for a typical antenna [Brooker et al., 4]. From this equation we can conclude that for a given frequency (77 GHz, \u03bb \u2248 4 mm), the only parameter we can modify to confine the electromagnetic radiation beamwidth is the antenna aperture. As pictured in figure 11, beamwidth defines the angular resolution. \u03b83 dB = 70\u03bb D deg. (23) The physical size of the UAV is the major constraint in the determination of the antenna diameter. A 150 mm horn-lens antenna is used on the Brumby Mark.III UAV, producing a 1.8\u25e6 3 dB beamwidth. Analogue to digital converter. The MMW radar front end is connected to a high speed 12 bit Analogue to Digital Converter (ADC) unit. Digitised data is stored on a 4096 word buffer and then transferred to DSP board memory through the high speed SHARC link ports", + " The gimbals have two degrees of freedom; one on the roll and the other on the pitch axis. The gimbals are controlled in real-time to counter the UAV\u2019s motions on these axes. The shaped scanning mirror above the gimbal reflects the millimetre waves radiated from the horn-lens antenna, as well as shaping the beam pattern to produce a wider elevation beamwidth. Received echo signals follow the same path from target to a receiver module via the scanning mirror and horn-lens antenna while the scanning mirror rotates at a frequency of 2.5 Hz. A simplified scan pattern is shown in figure 11. Servo controller and brushless servo motor drivers. The gimbal axis and scanning mirror are driven by separate brushless mini servo motors. Each motor is controlled by a 20 MHz PIC microcontroller based controller and driver board. These boards are daisy chained on a RS485 serial bus with commands sent by the 266 MHz Pentium II control computer as shown in figure 9. Control computer. The DSP unit passes the result of the FFT analysis (i.e. measured target ranges), to the control computer. The Control computer combines the range data with the gimbal and scanning mirror positions acquired from the servo controllers to form RBE data packets to be later converted into information form for DDF" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000168_fie.2002.1158229-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000168_fie.2002.1158229-Figure2-1.png", + "caption": "FIGURE 2 MULTI TOOL COMPONENTS AND FEATURES A S CAPTURED FROM PnolE", + "texts": [ + " The pliers have serrated jaws and a wire cutter, and there are #2 Philips and flat blade screwdrivers. They are made of tool steel that will be heat treated as part of the manufacturing process. Session 1: Safety and Material Preparation This session starts in the design suite, outlining the day's activities. The team is given the instruction packet and drawing package, then is asked to review this information. There is a discussion about solid modeling and drawings as a form of technical communication. An image of the ProE multi-tool assembly is shown in Figure 2, and partslfeatures are labeled for later reference in this paper. Expectations for drawings relating to their capstone project are explained by using the multi-tool drawings as an example. The first item to be covered in h e shop is safety. Maintaining an accident free shop is a top priority. After going over general shop safety rules the team is led on a tour of the shop. They see all the major equipment in the shop and with quick lessons on safety for each machine, and general instructions on their use", + " I This session will see the design suite again and discuss the questions from the previous week that haven't been answered yet can be fielded then, Machinery used in this session is: Ovens for heat treatment end ofmachining on the pieces, and shown in Figure 4 and is a set of replacement jaws for the preparation for heat treatment. The team and mentor meet in for the day. Manual mill #2 to create finish screwdriver tips Drill press to create holes for screwdriver pivots Manual mill #3 for making rounds Manual mill #2 to counter bore for rivet heads Students begin by finishing the screwdriver heads. They begin by cutting the flat blade screwdriver tip (Figure 2, part D) on the manual mill using tooling plate (fixture H) in Figure 5 to hold the stock. The angled cuts on the Philips head screwdriver are done on. the same fixture (fixture I). When both screwdriver tips are finished the students can measurements of tip thickness. FIGURE 4 REPLACEMENTViSEJhWSUSEV TO PREPhREPIllLlPSSCREWVRlV~RAND check their fit heads and JAW SERRATlONS 0-7803-7444-4/02/$17.00 0 2002 IEEE November 6-9,2002,Boston, MA 32\"' ASEEllEEE Frontiers in Education Conference F 4 D 9 The rest of the machining processes deal with cosmetics and preparation for riveting. Holes for the pivot on each screwdriver and mating center section (Figure 2, feature G) are drilled on a manual drill press with a fixture seen in Figure 6. The fixture has dowels to locate the ends of the parts and a guide hole in top locates the drill bit in the appropriate location for holes in the center sections (parts A and B) and screwdrivers (parts C and D). Manual mill #3 uses the tooling plate shown in Figure 7 to hold all four pieces together in the closed position. A 118\" radius cutter is used to round over the back edges (feature F) of the multitool. At this point in the session, manual mill #2 is available and the students work on spot facing and countersinking the holes to receive the rivets" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003996_cca.2009.5281178-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003996_cca.2009.5281178-Figure1-1.png", + "caption": "Fig. 1 Mechanical transmission system", + "texts": [ + " INTRODUCTION T is known that mechanical systems often contain some non-smooth nonlinearities such as dead-zone, friction, backlash, and backlash-like hysteresis. Among those non-smooth nonlinearities, backlash-like hysteresis is even more complex than the others due to its asymmetric feature with multi-valued mapping. Usually, in mechanical transmission system, the phenomenon of backlash comes from the space between the gear teeth. If the gears wear, it may cause larger but asymmetrical space when the gears move forward and backward. In this case, the asymmetric phenomenon of the backlash is called as backlash-like hysteresis. Fig. 1 shows a mechanical transmission system servomotor, gearbox, screw and work platform. In this system, the servomotor is used to drive a gearbox connected with a mechanical work platform through a screw. In this system, u is the servomotor rotational angle, x is the angle of the Ruili Dong is with College of Mechanical and Electronic Engineering, Shanghai Normal University, Shanghai 201814, China, e-mail: npu_dongruili@ hotmail.com). Yonghong Tan is with College of Mechanical and Electronic Engineering, Shanghai Normal University, Shanghai 201814, China", + " In section II, the Hammerstein model with backlash-like hysteresis is used to describe the nonlinear and dynamic characteristic of the mechanical transmission system. Then, in section III, the recursive algorithm based on the gradient technique to identify the parameters of the Hammerstein model with backlash-like hysteresis is developed. The bundle method is used to search for the optimizing direction. The simulation results are finally illustrated in section IV. The mechanical system demonstrated in Fig.1 can be described by a Hammerstein system with backlash shown in Fig.2. In fact, the motor rotational angle can be considered as the measurable input of the system, i.e. u(k). The measurable output, i.e. y(k) represents the displacement of the work platform. In this system, x(k), the rotational angle of the gearbox is the internal variable, which is supposed to be un-measurable. It is assumed that the smooth linear dynamic subsystem, i.e. L(\u00b7) is stable and the time-delay q in L(\u00b7) is known. It is also supposed that coefficient b0 in L(\u00b7) is equal to unity for unique representation", + " If a valid \u03b7(k) is found, and namely there occurs a significant decrease of the objective function f(\u00b7), then we can proceed in direction ( )kd from \u02c6( 1)k \u2212\u03b8 to \u02c6( )k\u03b8 (serious step). Otherwise, ( )kd is not a valid direction, and namely the objective function f(\u00b7) is increased. Then, the estimated parameters will not be updated. Hence, the algorithm approximates the effectively searching direction at non-smooth points based on the bundle method, which is what the smooth optimization methods can not be realized. In this section, the proposed control method is used for a simulated mechanical transmission system shown in Fig.1. In this system, the servomotor is used to drive a gearbox connected with a mechanical work platform through a screw. Hence, this mechanical system can be described by the above-stated model. Suppose the parameters of the backlash-like hysteresis in the system are respectively m1=1, m2=1.5, c1=0.5 and c2=1. The linear part of the work platform can be described by: 2 1.20 12.03( ) 2.01 17.22 sG s s s + = + + . (11) In this simulation, the signal to excite the system is a random sequence with variance \u03c32=12" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001712_nme.1428-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001712_nme.1428-Figure2-1.png", + "caption": "Figure 2. The triangular traction distribution method.", + "texts": [ + " Evaluate the derivative with respect to the tangential co-ordinate of the relative surface displacement of substrate surface due to a standard triangular distribution of normal and tangential tractions. 2. Assume a derivative with respect to the tangential co-ordinate of the vertical relative displacement (i.e. the gap function) of the indenter and substrate surface due to the combination of normal and tangential tractions. 3. Express this assumed gap function as a combination of standard triangular elements with appropriate peak values (Figure 2). 4. Use an iterative procedure involving the adjustment of the gap function values and the contact width until the required applied normal load and moment equilibrium is achieved. The triangular element approach to discretization of the unknown continuous traction distribution is in fact equivalent to the trapezoidal rule for subinterval in numerical integration methods. In comparison with the Gaussian quadrature approach this method has somewhat lower accuracy for the same degree of discretization" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000881_robot.2004.1308085-Figure4-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000881_robot.2004.1308085-Figure4-1.png", + "caption": "Fig. 4. The Pivoting Dynamics model simplifies the RRRobot-ona-plane model (see Figure 3) into two pans: (a) RRRobot pivoted at its geomemc center on a spherical joint and (b) a sphere on a plane.", + "texts": [ + " To analyze just the interaction between leg motion and body attitude, we simplify the RRRobot-on-a-plane model by pivoting the robot on a spherical joint and ignoring the effect of translation on body attitude dynamics. Once we compute the body attitude motion for a certain leg trajectory, we use the contact kinematics equations to approximately predict RRRobot translation in the plane. Thus, this model, called the Pivoting Dynamics model, approximately reduces the RRRobot system into two parts (see Fig. 4): 1) The dynamics of RRRobot rotating about a spherical joint, 2) The contact kinematics of a sphere on the plane. The configuration of the Pivoting Dynamics model qp consists of the sphere's orientation R(6'1,6'~, 6'3) with respect to a spatial frame and the configuration of its legs (41 .d~) . Thus, qp = (R(O1.82,@3),41.dz)~ E SO(3) x R2. model take the form The equations of motion for the Pivoting Dynamics A f ( q p ) i i p + C(qp, Yp)Qp + G(qp) = 7. (3) where A f ( q p ) E is the positive-definite nondiagonal variable mass matrix, C(qp, &) E R5 is the vector of Coriolis and centrifugal terms, G(qp) E R5 is the vector of gravitational terms, and T = (O,O, O , T ~ , ~ 2 ) ~ is the generalized force" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002597_acc.2008.4586601-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002597_acc.2008.4586601-Figure2-1.png", + "caption": "Fig. 2. The Flapper", + "texts": [ + " In the differential friction model, it is assumed that the friction coefficients satisfy the following ordering: \u00b5BW >> \u00b5t > \u00b5FW > 0, where \u00b5BW is much larger than the other friction coefficients. This implies that the backwards friction FBW is much larger when the scales slide backwards, which agrees with the geometry of the scales of the body. As a shorthand notation, we introduce a function \u00b5A(v), which describes the axial friction coefficient with the positive direction shown in Fig. 1. Thus, \u00b5a(v) = { \u00b5FW if v > 0; \u00b5BW if v < 0. (1) In this section we consider our first simple toy creature. Consider the flapper system in Fig. 2. Two (inflexible) rods are hinged at O with the scales orientations as shown. We assume that the instantaneous velocity of the flapper aOb is directed towards the left. The half-opening angle is \u03b8. Let \u03b8\u0307 = \u03c9 be the angular velocity of rod Oa. At a point P, which is a distance s away from the hinge O, the resulting linear velocity is \u03c9s. The combined velocity component in the axial and transversal direction of the section of the rod Oa at P is va = v cos \u03b8 and vt = \u03c9s + v sin \u03b8. This results in an axial and transversal friction force at this point Fa(s) = \u2212\u00b5a(v) v cos \u03b8 and Ft(s) = \u2212\u00b5t (v sin \u03b8+\u03c9s)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002459_s10778-009-0111-0-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002459_s10778-009-0111-0-Figure1-1.png", + "caption": "Fig. 1", + "texts": [ + " Despite the fact that the dimensions of vehicles, the minimum turning radii, and the geometry of obstacles are different, we can develop a general strategy of motion that would combine nearly minimum number of maneuvers. The present paper addresses [4] the motion of a rectangular (in plan) robot along an L-shaped holding alley with a \u2013 /2-turn. For other configurations of obstacles, including corners other than /2, similar formulas can be used after some correction. The robot is considered to have three wheels, which make it statically determinate. The first (I) and second (II) legs of the alley are of different widths a1 and a2 (Fig. 1). The results outlined below can be extended to holding alleys of other configurations. For different possible maneuvers starting from some position, trigonometric formulas are derived to describe the motion of the vehicle to different obstacles and change of its heading. This allows us to choose a specific maneuver or a group of maneuvers starting from a specific position. We will consider the following maneuvers: translations and turns around pointsO1 andO2 located in plan to the left and to the right of the vehicle\u2019s longitudinal line of symmetry [8]. All these maneuvers can be executed in the forward and backward directions, the radius of curvature of the trajectory being generally constant. We will use an example to compare the algorithms proposed here and in [4]. 1. Geometric Characteristics of Limiting Trajectories. An idealized vehicle in the form of a straight-line segment A1A4 with all wheels steerable can move on a horizontal plane from one leg of the L-shaped alley to the other (Fig. 1) if the vehicle length l and the dimensions a1 and a2 satisfy the inequality 1 0 , 0 1 0 2 3 2 0 2 3 3 2 ( ) , / , / / a a , (1.1) where a a l i i i, / , ,0 1 2 . The right-hand side of (1.1) corresponds to such relative parameters a i ,0 and 0 at which the robot A1A4 in a limiting position touches the alley walls by three points simultaneously: A1, A4, and C (which is in contact with the corner point C of the alley). All the three perpendiculars to walls 1 and 2 at the points A1 and A4 and to the robot body at the point C intersect at one International Applied Mechanics, Vol", + " Whether the robot is capable of traversing the L-shaped alley can be judged from inequality (1.1), which ensures that the robot contacts the alley walls at no more than two points to provide freedom and static determinacy. Given the leg lengths a1 and a2, the angle between the horizontal axis and the robot of minimum length 0 l that does not allow it to traverse the alley can be found as follows: arctan a a 2 0 1 0 3 , , . (1.2) For a real vehicle rectangular in plan, the limiting trajectory of the right (Fig. 1) side (A2A3) cannot pass closer to the axes of xOy than the envelope of the corresponding family of straight-lines segments A2A3 when the robot touches walls 1 and 2 by the points A1 and A4 simultaneously. This envelope is described by the following parametric equation in the coordinate system xOy (Fig. 1): x b d b0 0 2 3 2 0 01 ( ) / , y b d b0 0 3 0 0 2 1 2 1 ( ) / , (1.3) where x x l 0 and y y l 0 . The vehicle width d d l 0 . The parameter b0 , which is negative (from \u20131 to 0), is equal to the ratio of the ordinate of the point A4 to the length l of the vehicle A1A2A3A4 that touches the points A4 and A1 of walls 1 and 2 (Fig. 1). Equations (1.3) are obtained from the simultaneous solution of the equation of the family of straight-line segments A2A3 with varying parameter b y A 4 (i.e., b y l A0 4 / ) and partial derivative with respect to b, i.e., F x y b( , , ) 0, F x y b b ( , , ) 0. (1.4) The angular coefficient of the envelope (1.3) at a current point is defined by dy dx b b 0 0 0 0 2 1 . (1.5) The envelope of left side A1A4 can be obtained similarly to (1.4), but for the family of straight-line segments A1A4. After the parameter b is eliminated, we get the envelope equation y x0 0 2 3 3 2 1 ( ) / / , (1", + " For a rectangular (in plan) vehicle to traverse the holding alley, the corner point C should be beyond the region bounded by walls 1 and 2 and envelope 4 (Fig. 2), according to Eqs. (1.3). It should be noted that since curve 4 is ambiguous, the envelope should be used if d x l , l y d. The sequential maneuvers executed by the rectangular vehicle are turns around the center O1 or O2, either clockwise or counterclockwise, and translations in either forward or backward direction [4, 8]. The turning radii R1 and R2 in Fig. 1 correspond to the segment BO1 or BO2, where the point B is at the intersection of the vehicle\u2019s longitudinal axis of symmetry and the axle of the nonsteerable wheels. A mobile robot with all four wheels steerable and turnable by / 2 is theoretically capable of following a trajectory arbitrarily close to (1.3), with continuous change in the angle . The instantaneous centers of velocities of this robot are arranged on a circle of radius l centered at the point O (the origin of the coordinate system)", + " With all other conditions being equal, decisive factors are the arrangement of steerable and nonsteerable wheels (whether they are front or rear wheels) and the position of the vehicle in the alley (within which leg of the alley, the wider or narrower one, the vehicle starts moving). It should be noted that turns with minimum radius (R i = Rmin) do not generally ensure the minimum number of maneuvers to traverse the alley. Thus, to choose a program trajectory, it is necessary to establish a sequence of maneuvers, whether they are completely or partially executed before an obstacle, and values of turning angles or translations. The vehicle is considered as a body (Fig. 1) on a plane with nonholonomic constraints [2, 3, 7] at contact points between the wheels and the surface. The system has one degree of freedom and its position is determined at any instant by three generalized coordinates, which are the Cartesian coordinates x B and y B of the point B and the angle between the vehicle\u2019s longitudinal axis of symmetry and the Ox-axis. The coordinates of the other necessary points (A A A A E1 2 3 4, , , , ) are functions of these generalized coordinates. The body and wheels of the vehicle and the walls of the alley are considered rigid; the frictional constraints between the wheels and the surface are considered as bilateral", + " For example, for a turn around O j , we have x x R B B j j( ) ( ) ( ) ( ) ( )( ) (sin sin ) 1 1 11 , y y R B B j j( ) ( ) ( ) ( ) ( )( ) (cos cos ) 1 1 11 , (2.3) where R j ( ) 1 are the radii of turning around O j1 1( ) and O j2 2( ) in the ( 1)th maneuver. Values of R j ( ) 1 can be either given or determined from the strategy of motion in the presence of geometrical obstacles. We introduce the following notation for the maneuvers of turning around the centers O1 and O2 and trigonometric formulas for the maximum angles ( )( ) 1 : 1,1 \u2014 clockwise turn around O1 until the point A4 touches wall 1 (Fig. 1), ( , )( ) ( ) ( ) . 1 1 1 1 1 12 0 5 arctan arccos l l R d x R R d l l B ( ) ( ) ( ) ( ) sin ( . ) ( ) 1 1 1 1 2 1 2 0 5 , (2.4) 1,2 \u2014 counter-clockwise turn around O1 until the point A1 touches wall 2, ( , )( ) ( ) ( ) . 1 2 1 1 1 12 0 5 arctan arcsin l R d y B ( ) ( ) ( ) ( ) cos ( . ) R R d l 1 1 1 1 2 1 2 0 5 , (2.5) 1,3 \u2014 counter-clockwise turn around O1 until the point A2 touches wall 2, ( , )( ) ( ) ( ) . 1 3 1 1 1 12 0 5 arctan arcsin l R d y B R R d l 1 1 1 1 2 1 2 0 5 ( ) ( ) ( ) cos ( ", + "18) for complete or partial motion in both directions in the presence of s constraints imposed by the alley. It is very difficult to use the minimum number of maneuvers when the trajectory is not smooth at connection points; therefore, it is expedient to use a little greater number of maneuvers than the minimum possible, which would considerably simplify the preparation and calculation of various combinations of maneuvers. In synthesizing a trajectory based on (2.4)\u2013(2.18) for an L-shaped alley (Fig. 1), it should be borne in mind that the most difficult case (requiring a great number of maneuvers) is the passage of a vehicle with steerable front wheels from the wider leg to the narrower leg of the alley (a a1 2 ). It is well to start forming a maneuver sequence from the consideration of the robot position in this alley (Fig. 3). The robot touches wall 2 by the point A1, and the point E (which is the intersection of the side A A2 3 and the axes of the nonsteerable rear wheels) coincides with the corner point C", + " Much less maneuvers are needed if m defined by (3.2) is complex, i.e., if the points C and E coincide, then the robot cannot touch wall 2 by the point A1 for any value of . 4. Numerical Example. We used the approach outlined above and MATLAB software to simulate maneuvers of a robot in an L-shaped alley with the following geometrical parameters [4]: a1 1.6, a2 1.3, l 1.5, l1 1.2, d 1, Rmin 1, x C 1.6, y C \u20131.3. Since conditions (3.3) fail, some initial maneuvers are executed to move the robot to the upper left corner (Fig. 1). Figure 4a shows a maneuver sequence executed by a robot with steerable front wheels to traverse a holding alley. The total number of maneuvers is 11, which is nine maneuvers less than in [4]. The sequence includes the following maneuvers: 1,5; 3,4; 2,1; 1,1; 2,2; 3,3 (partial maneuver); 1,2 (partial maneuver); 2,2; 3,3 (partial maneuver); 1,2 (partial maneuver); 2,1 (partial maneuver); turning radii are equal to 2,7 and 3,4 in the 10th and 11th maneuvers, respectively, and to 1 in the other maneuvers" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000270_iemdc.1999.769039-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000270_iemdc.1999.769039-Figure1-1.png", + "caption": "Fig. 1. Rotor cross section.", + "texts": [], + "surrounding_texts": [ + "I. INTRODUCTION\nPrediction of motor parameters when the motor is still in the stage of construction is one of the most important tasks of every motor designer. Generally, motor parameters are not constant but vary depending on load conditions and the level of saturation of flux paths in the motor [I].\nMagnetic conditions at different load conditions can be easily simulated by the 2D Finite Element Method.\nConsequently, the authors proposed an efficient procedure for calculation of the parameters of the two-axis model of a synchronous motor with permanent magnets by postprocessing the static magnetic field calculation results by 2D Finite Element Method (FEM).\n0-7803-5293-9/99 $10.00 0 1999 IEEE 98\n11. METHOD OF ANALYSIS The magnetic conditions in the motor were computed by\n(1) where v denotes the reluctivity, A is the magnetic vector potential, J, is the current density, and M is the magnetization.\nIn the case of sinusoidal balanced supply, the current of the first phase is given by\n2D FEM using the basic equation\nrot( v rot(A)) = J,, + rot Y M.\ni, = J?:-I-cos(ot+n. (2)\nwhere I is R M S value of stator current, p initial phase angle and o electrical angular velocity. In the remaining two phases the phase shift of currents is f120\" in the relation to the first phase.\nRotor starting position wt=O is chosen in such a way that the magnet axis of phase a is in alignment with the rotor direct axis. Then the initial phase angle p of the stator current is also the electrical angle between the stator MMF and rotor direct axis. For steady state conditions at constant load, constant voltage and frequency, electrical angle p is constant. Different operating modes can be simulated just by shifting the initial angle p. For each point of load the discrete time forms of phase voltages were calculated from the average values of the vector magnetic potential in stator slots for different discrete time moments by moving the rotor body and simultaneously changing the stator excitation over a half of electrical cycle. The end winding contribution was included in the modelling with constant value of end winding inductance Le. The instantaneous value of the phase voltage in the winding of phase a is given by\nwhere R is phase resistance,y/, is the instantaneous flux linkage of the a phase winding per pole, p is the number of pole pairs and c is the number of parallel circuit of the phase winding. The RMS values of phase voltages were calculated from their calculated time forms.\nThe load angle 6and the phase angle p were calculated from the calculated time form of the phase voltage and known time form of supply current. In the calculation of the phase winding flux linkage per pole the skewing of the stator slots is taken into account by calculation of average value of flux linkage of three slices of the stator package along z axis.", + "From the phasor diagram the components of the induced voltage in quadrature and direct axis Ek and Ea are defined by equations (4) and (5):\nEiq = Ei .COSSi= Eo +I\u2019cosp.xd\nEid = Ei -s insi= I - s inp -X, .\n(4)\n( 5 ) where Xd and Xq are direct and quadrature synchronous reactance, EO induced EMF due to the magnets, 6i internal load angle between the phasor of induced voltage Ei and quadrature axis and p torque angle between the phasor of stator current and direct axis.\nThe components of induced voltage Eh and Eid are determined as\n(6)\n(7)\nThe parameter X, can be obtained from (5), but & cannot be determined from (4) without the assumption of constant value of Eo. The uniqueness of the separate determination of EO and Xd depends on the superposition which cannot be applied under saturated conditions. Because of this condition the variation of the parameter Eo must be taken into account in determination of Xd.\nFor this reason an additional expression to expression (4) is needed. On the supposition of a small change of the stator current the magnetic conditions in the motor are the same and the parameters Xd and Eo remain constant.\nFrom measured or calculated data of load conditions the approximation cumes Lj(p), I v ) and E @ ) with use of orthogonal polinoms approximation were calculated.\nFor p\u2019=p +Ap ( A p =O. 1 degree) curves S;\u2019(p \u2019), I\u2019(p \u2019) and Ei\u2019(p\u2019) were made and considered in (8)\nEin = V .cos8 - I. R .cos(g, - 8)\nE,d = V - s ins+ I . R .sin((p - s).\nE; \u2018 C O S 8 := Eo + I\u2019 * C O S P \u2019 . xd. (8)\nParameter Xd and parameter Eo are then calculated from (4) and (8).\n014- 0\u20193 -. 0 12 -. -0 I 1 -. &IO-.\n3007 5006 %oos 3004 -om\ng: 1:\n002 001 000 i\n111. UNBALANCED STANDSTILL WORK TEST\nThe exact value of end winding inductance can be determined by combining unbalanced standstill work test and its simulation by FEM. The experimental arrangement of unbalanced standstill work test is presented in Fig. 3 [2].\n- n\n+tolal w e mjwtmx+ f r ~ m t leakage .Q- e d m l e a l c d g e - L e ( H ) s ~ t r l l ~ r a k test by XJ -cl+ Cnd e d m\n-. - - -. -. -. -. - -\nI LR - V b +\nFig. 3. W Unbalanced standstill work test arrangement. Fig. 4. Variation of end winding inductance L (H)." + ] + }, + { + "image_filename": "designv11_69_0001218_detc2006-99153-Figure5-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001218_detc2006-99153-Figure5-1.png", + "caption": "Figure 5. Three dimensional Reuleaux\u2019s method.", + "texts": [ + "ur Considering a pair of skew lines in space, the Plu\u0308cker coordinates of the common perpendicular between these two lines are derived in the Appedix. Let l\u0302a = (la, l0a) and l\u0302b = (lb, l0b) be a pair of lines given in their Plu\u0308cker coordinates that belong to the rigid body before a displacement. The homologous lines l\u0302\u2032a = (l\u2032a, l \u20320 a ) and l\u0302\u2032b = (l\u2032b, l \u20320 b ) correspond to the same lines after the rigid body displacement (see fig.4). A line can be defined by two points or by intersection of two planes. Figure 5 shows the three dimensional generalization of Reuleaux\u2019s method. For each pair of the homologous lines, we find the common perpendicular to the pair. We use equation (14) derived in the Appendix to find the Plu\u0308cker coordinates of the common perpendicular between the two lines. The common perpendicular between l\u0302a and l\u0302\u2032a is s\u0302a and between l\u0302b and l\u0302\u2032b is s\u0302b and they are written as (see fig. 5): s\u0302a = (sa,s0 a) = l\u0302a\u00d7 l\u0302\u2032a s\u0302b = (sb,s0 b) = l\u0302b\u00d7 l\u0302\u2032b Then we find the midline of each pair, which is the line at half distance between the two homologous lines and has half the angle between them. Midlines can be written as m\u0302a = (ma,m0 a) = 1 2 (l\u0302a + l\u0302\u2032b) m\u0302b = (mb,m0 b) = 1 2 (l\u0302b + l\u0302\u2032b) Next, we find the lines perpendicular to both the midlines m\u0302a and m\u0302b and the common perpendiculars s\u0302a and s\u0302b (see fig. Copyright c\u00a9 2006 by ASME l=/data/conferences/idetc/cie2006/71356/ on 02/27/2017 Terms of Use: http://www" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000279_iecon.2001.975548-Figure5-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000279_iecon.2001.975548-Figure5-1.png", + "caption": "Figure 5: %lation between F,\", fi and velocity v", + "texts": [ + " B.C : ~ ( o ) =s eo , 3e(q =s ef P = /; COG command in sagittal plane is obtained as shown in follow equations(Eq.12, 13). = \"lCog sin \"e(t) (12) (13) zcmd cog cmd - init zcog - zc*g %og = p x 2 0 g + 3z:0,g (14) With these trajectories, maximum acceleration of COG increase step by step till robot detects slipping. In stage 2, proper trajectory planning not to slip on slippery floor is considered. Relation between maximum force in each step and velocity of support foot is expressed as shown in Fig.5. F, and Fg are the force which is added to floor by robot in x-axis and z-axis respectively. F,\" is the largest F, in n-th step, and Fl is friction force between foot and floor. v is the velocity of support leg foot in world coordinate when the robot is slipping. Here, the biped robot is supposed to slip in (n+l)th step as shown in Fig.5. It is clear that if the biped robot is not slipping F,\" is less than F,,,, in each step. It shows the possibility for biped robot to walk with maximum acceleration of n-th step. To realize \"walk not to slip\" and \"walk with the fastest walking speed available\" at the same time, it is advisable to walk with conditions of n-th step, constant stride S,, and constant walking frequency f. Stage 2 is for planning trajectory to walk as fast as possible, but not to slip. If nothing happens, it can be said that this walking trajectory decided in stage 2 is stable" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000251_iis.1997.645430-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000251_iis.1997.645430-Figure1-1.png", + "caption": "Figure 1: Example Constrained Manipulator", + "texts": [ + " Also, constraint forces are not represented as Lagrange multipliers, but as the generalized forces associated with the chosen task coordinates. An example constrained robot is depicted in Fig- 'The noncontact case may be regarded as a special case of compliant contact for which the contact stiffness is zero. 0-8186-8218-3/97 $10.00 0 1997 IEEE ure 1. An end effector mounted on the distal link of the manipulator is in contact with a constraint. The end effector of the planar manipulator shown in Figure 1 is simply a point in two dimensional space, and is constrained to a line. More generally, we model the end effector as a rigid body in three dimensional space and the constraint as a mechanism that allows certain end effector positions and orientations but not others. The position and orientation of the end effector is parameterized by a vector of generalized coordinates p E R\", where n is the number of degrees of freedom (DOF) of the end effector when unconstrained. For example, n = 6 for a three dimensional rigid body, n = 3 for a planar rigid body, and n = 2 for a single point in a plane (as in Figure 1). In the case of a three dimensional rigid body end effector, p may be chosen as p = ( E ) , where pt is the three Cartesian coordinates of a point on the end effector, and p, is the roll, pitch, and yaw angles of the end effector. The constraint is modeled as a set of smooth algebraic equations in p . That is, we assume there is a smooth (i.e. C - ) function 9, : R\" + R\" such that a P ( p ) = 0 and that rank -(p) a@, = m aP throughout the region of interest. Note that the constraint shown in Figure 1 is actually of the form a P ( p ) 5 0, since it permits 'loss of contact'. However, we will use the 'bilateral' form ( l ) , since we wish to compare our stability results to those reported in [5] for bilateral elastic constraints. As well as moving on the constraint, the end effector can apply a generalized force Fp E Rn to the constraint, where the subscript p signifies that Fp is the generalized force corresponding to the generalized coordinates p . In the general case where n = 6 and p is chosen as three Cartesian position coordinates and three orientation angles, Fp consists of three force components and three torque components. We assume that the constraint is workless, i.e., Fp does no work as the end effector moves on the constraint. It is convenient to define a new set of generalized coordinates x E R\" that are naturally related to the constraint. In particular, x is chosen such that there is a smooth diffeomorphism x = X,(p) in the region of interest and 22 = 0 o:n the constraint, where 22 is defined by the partition Such coordinates are called task coordinates for the constraint, and are illustrated in Figure 1. Let Fxl E R\"-\" and Fx2 E R\" be the generalized forces of constraint corresponding to x1 and 22 respectively, so that F, = ( 2; ). It is shown in [12] that the workless constraint assumption implies that Fzl = 0, which agrees wi.th Figure 1 for our two degree of freedom example. When expressed in task coordinates, the constraint model is thus 22 = 0 (3) Fzl = 0. (4) Note that since x2 and ,F,1 are fixed, they cannot be controlled. The goal of hybrid position/force control is thus to control the independent variables x1 and Fx2 Task coordinates can be found for any holonomic constraint. A general method for choosing task coordinates is given in [ll]. However, for a given constraint, the task coordin.ates are not unique. For example, the method of [Ll] applied to the constraint of Figure 1 gives not Cartesian task coordinates, but curvilinear task coordinattes that depend on the robot kinematics. Also, for certain constraints, Cartesian task coordinates cannot be used [12, 131. Let q E Rn be the vector of manipulator joint coordinates, which we assume to be related to p by a smooth diffeomorphism. Then q and a: are diffeomorphic as well. To be precise, let 0, and 0, be open subsets of R\", and let X be a smooth diffeormorphism from 0, onto 0,. We assume that q and z are related by 2 = X ( q ) , vq E 0,", + " Therefore, by Lyapunov's linearization theorem ([15], page 179), the nonlinear system (33), (36) is exponentially stable. Exponential stability implies that limt+m z l ( t ) = ad and l im+m 5 l ( t ) = 0. Since g j in (36) is a C\" function of z1 and XI, andl g j ( O , O , z d ) = 0, we have l im+w g j (zl(i!), ;l(t), z d ) = 0. Therefore (36) gives limt+m F,z(t) = Fd. U Remark 1 The system is exponentially stable for all z d E Ozl, i.e. at all1 positions on the constraint for which the diffeomorphism z = X ( q ) is defined. For example, in Figure 1, the system is exponentially stable at all reachable points on the constraint, less the two singular points where 8X-'(z)/c?z (i.e. the Jacobian inverse) is undefined. Remark 2 Since Theorem 1 holds when k j = 0, we observe that even when force feedback is not used, J-'-control is exponentially stable and k - + w FzZ(t) = Fd. The term 'kinematic st,ability', as used in [4] and [5], refers to the stability of hybrid position/force controls in the absence of force feedback. Thus, it follows specifically from Remark 2 that J-l-control is always kinematically stable when applied to rigid constraints" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001231_sice.2006.314734-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001231_sice.2006.314734-Figure1-1.png", + "caption": "Fig. 1. Bearing structure and fundamental frequencies.", + "texts": [ + " BEARING FAULT FREQUENCIES Due to single point defects, certain characteristic fault frequencies appear in machine vibration. These frequencies are predictable and depend on which surface of the bearing contains the fault. There are five basic frequencies that are used to describe the dynamic of bearing elements: Shaft rotation frequency (FS ), cage frequency (FC ), ball pass inner raceway frequency ( FBPI ), ball pass outer raceway frequency ( FBPO ), and ball rotational frequency ( FB) as shown in Fig. 1. These frequencies are given by: FC = Fs (1 - DbCosOlDc)/2 (1) FBPI = NBFs (1+ DbCosO DC)/2 (2) FBO =NBFs(1 DbCosOlDc)/2 (3) FB= (DCFs/Db)(1-D 2CoS2ID7)/2 (4) Where, Db is the ball diameter, D is the bearing cage diameter measured from one ball center to the opposite ball center, and 0 is the contact angle of the bearing. These frequencies have small amplitudes in healthy bearings. When a raceway fault occurs, the frequency associated with the defective part of bearing ( FBPO or FBPJ ) and its harmonics are excited" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000677_ps.2780060511-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000677_ps.2780060511-Figure1-1.png", + "caption": "Figure 1. Response of chromogenic strips to various concentrations of phosphine (strips exposed for 30 min). - x -, Cresol red, strip (a); ---.--, dimethyl yellow, strip (b); -0-, cresol red-kdimethyl yellow, strip (C).", + "texts": [ + " A few experiments were also carried out to test the response of the strips in the presence of the stoichiometric volume of ammonia (1 : 3 v/v) that is released from the Phostoxin@ (aluminium phosphide) tablets along with PH3, and also a slightly higher concentration of ammonia (i.e. 1 : 6 ratio). There were four replicates in addition to the control for each treatment and each experiment was repeated once. As the concentration of phosphine increased from 0.05 to 0.9 part/106 the intensity (optical density) of colour also increased (Figure 1). In all the concentrations of phosphine tested the intensity of colour developed by the strips impregnated with a mixture of mercuric chloride, cresol red and dimethyl yellow was far more than the strips impregnated with either mercuric chloride plus cresol red or mercuric chloride Indicator for phosphine 513 plus dimethyl yellow. At the lowest dosage of phosphine tested (i.e. 0.05 part/l06) and at concentrations up to 0.3 part/l06 strip (c) was three times as sensitive as the other strips. The development of colour was also much quicker in the case of strip (c) than with the strips containing single indicators (Figure 2)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000976_cp:20040396-Figure5-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000976_cp:20040396-Figure5-1.png", + "caption": "Fig. 5 Typical flux distributions related to 15-slot, IO-pole motor.", + "texts": [], + "surrounding_texts": [ + "The superposition method is now extended to sum thr contributions of each stator slot to the resultant cogging torque. The two IO-pole motors.whose parameters are given in Table I , and which are assumed to have a pole-arc to polepitch ratio of 1 .O, are considered The predicted results for the IO-pole, 15-slot motor are shown in Figs. 5 and 6. The cogging torque waveform which results with a 3-slot stator is synthesized from that which results with a single slot stator by phase shifting its cogging torque waveform by 0, I20 and 240 degrees mechanical and summing, as shown in Fig. 6a. 111 a similar manner, the cogging torque waveform which results with the 15-slot stator is obtained, Fig. 6b. It will be noted that when there is only one slot on the stator of a 10-pole motor, the least common multiple N, is IO, while N, is 30 for both a 3-slot stator and a 15-slot stator. Hence, the periodicity of the cogging torque waveform for the I-slot stator is 36 mechanical degrees, while that for the 3-slot and 15-slot stators is 12 mechanical degrees, as confirmed by the finite element predictions shown in Fig. 6. It can also be seen that excellent agreement is achieved between the synthesized and the resultant finite element calculated cogging torque waveforms. The same synthesis procedure has been applied to the 10- pole, 12-slot motor, as shown in Figs. 7 and 8, in which the cogging torque which results with a .2-slot stator is also shown. The least common multiple N, is now I O for both the I-slot and 2-slot stators, while for the 12-slot stator N , is 60. Therefore, the periodicity of the cogging torque waveforms, which result with the I-slot and 2-slot stators is the same, viz. 36 mechanical degrees, while that for the 12-slot stator is only 6 mechanical degrees. Again, excellent agreement is achieved between the analytically synthesized and the resultant finite element calculated cogging torque waveforms. 0 10 20 30 4a 1\" fQ A o ~ u l s r Polifion (harat) (a) I-magnet and 2-magnet rotors I 0 6 l i 18 21 10 36. *\"gular Position (Degree, (a) I-slot and 3-slot stators 832 Since the width of the stator slot openings is usually much smaller than the slot-pitch, the interaction between two adjacent slots is usually negligible. Hence, in general, the resultant cogging torque waveform can be synthesized from the cogging torque waveform which results with a single slot to a high accuracy. 4 Comparison of Predicted and Measured Cogging Torque Waveforms In order to validate the accuracy of the foregoing calculations, cogging torque waveforms were measured on a 4-pole, 6-slot motor, Fig.9. Fig.10 compares the predicted and measured cogging torque waveforms when the motor is equipped with each of the rotors shown in Figs.9b-d for which the magnet pole-arc to pole-pitch ratio is 0.7, as well as with rotors having a pole-arc to pole-pitch ratio of 1.0, the corresponding flux distributions being shown in Figs. 1 and 3. It should be noted that the predicted results are deduced from finite element analysis, rather than being synthesized. As can be seen, excellent agreement is achieved between predictions and measurements for both pole-arc to pole-pitch ratios, which confirms the discussion presented in section 2. 20 *\u201clYI., P..ili\u201d. , \u201d C s r r , (a) Pole-arc to pole-pitch ratio=0.7 *\u201cI\u201c,.. P O m m l\u201d. (3.135) By defining an electrostatic energy density, we=(l/2)D\u00c8 = (y2)e\\\u00c9\\2=(l/2)\\D\\2/e, (3.136) We=j wedv. (3.137) where EXAMPLE 3.7 Static fields in a coaxial cable: A coaxial cable consists of a long, cylindrical conductor of radius ra, an air space between ra and r,\u201e a dielectric insulator between rh and rc, an air space between r, and rd, and a grounded conducting sheath between rd and re, as shown in Figure 3.33. As a function of the radius, r, plot the magnitude of the following: a. electric field intensity, E b. electric flux density, D \u0441 polarization field, P d. electric potential, V e. Compute the capacitance per unit length, C,, of a coaxial cable. 96 Chapter 3 Static Electric Fields SOLUTION a. The electric field intensity inside a conductor is zero so |\u00a3| = 0 for 0 0, I\u0304 \u2208 R3\u00d73, is the net inertia matrix of the entire craft. \u0307 \u2261 d dt ( ) . \u03c4 \u2208 R3 is the aerodynamic torque. Ir \u2208 R1 is the rotational inertia of a fan\u2019s rotor about its own axis, which is parallel to \u21c0 kA. \u03c9i \u2208 R1, (i = 1, 2, 3, 4) are the fan angular velocities, relative to frame A, with the positive convention shown in Figure 1. i , (i = 1, 2, 3, 4) are the four control inputs to the system (1)\u2013(7), and they denote motor torques. Remark 2.1: The translation equation of motion (1) does not contain any aero-dynamic drag forces, as would exist on the fuselage of the vehicle. If these forces are sufficiently large, the ability to switch from the Phase I controller herein to the Phase II controller may require additional actuation inputs. In particular, two inputs that can exert forces on the vehicle perpendicular to the thrust vector TRe3 would be appropriate" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003309_icems.2009.5382765-Figure3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003309_icems.2009.5382765-Figure3-1.png", + "caption": "Fig. 3: Supplied voltage deviation \u03b4V c and current response \u0394ic in", + "texts": [ + " The voltage deviations \u03b4vc q , \u03b4vc d from the steady-state voltage are constant in time and are given by \u03b4V c q and \u03b4V c d respectively. It is shown in [13] that the resulting current variations \u0394icq, \u0394icd after a time \u03c4 << min(\u03c4q, \u03c4d), with \u03c4q, \u03c4d the synchronous time constants of qand d-axis, can be approximated by \u0394icq = \u03c4 Lq \u03b4V c q and \u0394icd = \u03c4 Ld \u03b4V c d , (1) independent of the stator resistor Rs as well as rotor speed \u03a9. To estimate the rotor position \u03b8r, an auxiliary angle is used: the angle \u03b3 of the voltage vector \u03b4V c(\u03b4V c q , \u03b4V c d ) with respect to the q-axis, Fig. 3. As the direction of the voltage test vector is known in the stationary \u03b1\u03b2-reference frame and given by \u03b3\u2032, an estimation of the rotor angle \u03b8r is obtained as \u03b8\u0303r = \u03b3\u2032 \u2212 \u03b3\u0303 where x\u0303 denotes the estimation of x. This means that the dependence of the current response on \u03b3 instead of \u03b8r should be discussed from which an estimation of \u03b8r follows. The angle \u03b3 can be introduced as follows. By using the amplitude of \u03b4V c: \u03b4V c = \u221a (\u03b4V c q )2 + (\u03b4V c d )2, the current the qd-reference frame fixed to the rotor" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001192_s00419-006-0003-2-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001192_s00419-006-0003-2-Figure1-1.png", + "caption": "Fig. 1 The analytical model of stationary pin and rotating ring system", + "texts": [ + " The effect of thin film on the heat distribution, with taking the convective cooling from side faces under consideration was studied in paper [10]. The heat distribution, in both cases, between the pin and the disc was obtained by matching the average contact temperatures. In this paper the model proposed in papers [9,10] is generalized on the case of the rotating ring with convective cooling from external and internal surfaces. The studied system consists of the rotating steel ring sliding with regard to the end of stationary pin (Fig. 1). The ring turns fast in relation to the stationary pin with constant speed \u03c9. So, we have a state where the speed of the moving heat source is very much greater than the ratio of the thermal diffusivity and the radius of the ring [8]. As a result of friction the heat generation over the contact region in the form of a heat flows directed into the bodies, takes place. It is necessary to define temperatures on the sliding surface of a ring and a pin as well as the heat distribution ratio between them" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003275_detc2009-86970-Figure11-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003275_detc2009-86970-Figure11-1.png", + "caption": "Figure 11. Three-teeth double flank meshing model.", + "texts": [ + " 0)( )()( =\u2212 TR f MR f xtipx , (45) 0)( )()( =\u2212 TR f MR f ytipy , (46) 0)( )()( =\u2212 TL f ML f xtipx , (47) 0)( )()( =\u2212 TL f ML f ytipy . (48) 3.3 Multi-teeth meshing model and contact pair Based on the meshing model mentioned above, computer simulation programs have been developed to simulate the meshing of gear pairs in double flank gear rolling testing. Since the contact ratios of spur gear pairs are between 1 and 3 in general, a three-teeth meshing model comprising the (i)th, the (i+1)th and the (i-1)th teeth of the meshing gear pair is considered as shown in Fig.11. Notably, the cooperation of any two flanks picked from the left side and the right side tooth category may fulfil the double flank contact condition. Nine possible cases of tooth flanks collocation are listed in Table 1. As mentioned above, tooth contact types including surface-to-surface contact and tip-to-surface interference may occur in any contact position. Therefore, total 36 possible candidates are considered in the computer simulation program. During the execution of the computer simulation program, the maximum value of C\u0394 among the 36 possible candidates are recorded" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002964_wsc.2007.4419836-Figure5-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002964_wsc.2007.4419836-Figure5-1.png", + "caption": "Figure 5: Two Cranes Lifting a Panel (Zaki and Mailhot 2003)", + "texts": [ + " Through negotiation, an effective plan can be generated based on possible combinations of movements of cranes from one step to another. 4 CASE STUDY The re-decking project of Jacques Cartier Bridge in Montreal is used to demonstrate the proposed collaborative agent-based system. The deck of this bridge was replaced during 2001-2002. The existing deck was removed by saw-cutting the deck into sections. Each section was removed by two telescopic cranes and a new panel was installed using the same cranes. Figure 5 shows two telescopic cranes positioned on both sides of the section to be replaced. The case study focuses on the activities of removing existing deck sections and installing new panels on the main span of the bridge. Due to the low clearance of the truss structure on the main span of the bridge, the cranes had to work with a near-flat boom ( = 15\u00b0), which results in decreasing its lifting capacity by more than 50%. The contractor built a physical model of a part of the bridge and different types of cranes to check the feasibility and to plan the movement of each crane" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002461_bf02916353-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002461_bf02916353-Figure2-1.png", + "caption": "Fig. 2. Schematic diagram for modeling.", + "texts": [ + " In this study, a robust controller ( H controller) has been designed and applied for the horizontal control of a hydraulic-operated gimbal system that has disturbances and uncertainty of the system. The experimental system consists of the bottom (body 1) and the upper plates(body 2). The bottom plate corresponds to the deck of the vessel, and the upper plate is controlled to be maintained in horizontal state. These two plates are connected with a hydraulic cylinder making relative motion possible. Figure 1 shows the schematic diagram of the gimbal system. It is designed to control the angular move-ment of the body 2 against the motion of body 1, which corresponds the vessel. Figure 2 displays a sim-plified diagram for mathematical modeling of Fig. 1. Fig. 1. Schematic diagram of the gimbal system. axy is the fixed coordinate. ' ' 1 1 1a x y is the moving coordinate of body 1 and ' ' 1 2 2a x y is the moving coordinate of body 2. The centroids of each object 1O are 1O and respectively. With generalized coordinate for the rotating degree of body 1 and for the relative rotating degree of body 2 against body 1, the displacement and acceleration of body 2 are as in Eqs. (1) and (2)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001989_978-1-84800-239-5_34-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001989_978-1-84800-239-5_34-Figure2-1.png", + "caption": "Figure 2. a.Rack cutters definition; b. Longitudinal crowning of pinion surface", + "texts": [ + " ns is the unit normal to the shaper tooth surface. The profile of the pinion or the shaper in traditional face gear drives is standard involute, and it is generated by rack cutter with straight profile. Profile crowning to pinion or shaper can provide a predesigned parabolic function of transmission errors. Such a function is able to absorb almost linear discontinuous functions of transmission errors (caused by mis-alignment) that are the source of vibration and noise. For this reason, we use rack-cutters with parabolic profile (see Figure 2-a). ui(i=s,1) is the coordinate parameter along profile of the rack-cutter for pinion and shaper respectively. ai is the parabolic coefficient; u0 is the parameter of parabola apex. Longitudinal crowning is required for localization of bearing contact. In this paper, a new simple type of surface modification is proposed. The longitudinal crowning is illustrated by Figure 2-b. Keep the shape of profiles in sections normal to axis z1 invariable, rotate the profile toward tooth surface by a small angle . 2 2( ) /pl mida z z b (2) Here, apl is the parabolic coefficient for longitudinal crowning; z is the axis coordinate of any point on the pinion surface; zmid is the axis coordinate of midpoint on face width. b is the tooth width of pinion. This crowning can be performed on CNC machine tools. The tooth surfaces are in point contact but the contact is spread on an elliptical area due to elastic deformation of contact surfaces" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003564_1.4000520-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003564_1.4000520-Figure2-1.png", + "caption": "Fig. 2 TSSM with the assembled and disconnected third leg, respectively, and parallel rotary axes a1 ,a2", + "texts": [ + "org/ on 03/23/2 the implicit representation of yields a polynomial of degree 16 in the unknown h. As the circle c3 and have 2 12=24 intersection points, the 24\u221216=8 common points of c3 and must be located on the plane at infinity . Therefore, the two cyclic points of c3 which are the intersection points of c3 and located on the imaginary spherical circle must be fourfold points of the spin surface . 3 Consequences for Parallel Manipulators A parallel manipulator of TSSM type consists of a platform , which is connected via three spherical-prismatical-rotational legs with the base 0 see Fig. 2 , where the axes ai for i=1,2 ,3 of the rotational joints are coplanar. If we skip the assumption of coplanarity we get a more generalized class of parallel manipulators, which we will call GTSSM for the rest of this article. Theorem 1 has the following consequences for this class of parallel manipulators. Theorem 2. GTSSM manipulators with two parallel rotary axes cannot have more than 16 assembly modes except the degenerated cases with infinitely many solutions. Proof. For solving the direct kinematics of a given manipulator geometry of the platform and base as well as the length of the three legs are known we disconnect the third leg from the platform and consider the resulting RSSR mechanism with parallel rotary axes" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001164_1-84628-559-3_20-Figure3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001164_1-84628-559-3_20-Figure3-1.png", + "caption": "Fig. 3. Configuration of the 3-DOF semi-spehreical positioning stage", + "texts": [ + " 2 (d), the returning hammer may cause further movement of the stage with a distance X3. Though only one pulse waveform is applied to the PZT actuator, the cyclic process of (3-5) is iterating itself. Finally, the sliding stage will reach to a final position after the hammer damps to standstill. In this state, the PZT actuator is ready for the next actuation from step 2. Due to the special design of the modularized actuating unit, it is much easy to configure a positioning stage having multiDOF. Figure 3 shows an application example of semishperical 3-DOF positionig stage having the positioning ability along x-, y-, and z-axis. Since one actuating unit is capable of acutaing the stage in only one direction of motion, totally, six actuating units marked from (a) to (f) are required to performe the 3-DOF of positioning. For examples, the actuating units set on the top side of xy plane and marked with (a) and (b) are used for actuating the stage to rotate with respect to z-axis; simlarily, the actuating units set on the xz plane and marked with (f) and (e) are used for actuating the stage to rotate with respect to y-axis as shown in Fig. 4 which is the sectional view A-A in Fig. 3. Figure 5 shows the perspecitve schematic drawing of the semi-shperical 3-DOF positionig stage. The friction adjusting mechanism set on the top surface of the stage is used for providing a suitable holding fricitonal force for the rotaional stage. Figure 6 shows a photograph of the modualrized springmounted PZT actuator, in which the PZT actuator has the dimension of 5 5 10 mm (Tokin). The stiffness of the spring is 0.023 N/mm. Six actuating units are symmetrically mounted to the positioning stage having a radius on the bottom side" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000674_05698190490513983-Figure3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000674_05698190490513983-Figure3-1.png", + "caption": "Fig. 3\u2014Crack initiation and propagation on the surface in circular contact; arrow shows origin of cracking (Muro et al. (11)).", + "texts": [ + "5 GPa, but the direction changes variably as the loading point moves; then its contribution would be small. Therefore, it is considered that the crack propagates parallel to the surface mainly by residual tensile strain E1zxr . In the rolling contact test of the ring specimen (Muro, et al. (11)) with 50 mm diameter and with certain curvature in the axial direction, when the contact surface is circular (that is, a/b = 1), the crack initiates not at the center of the contact track but at 2a/3 away from the center of the contact track. Figure 3 shows this situation: A crack initiated 2a/3 away from the center of the contact track and then propagated at about 45\u25e6 to the rolling direction. The crack propagating toward the center of the contact track turned its direction to the rolling direction after reaching the center. The crack also propagated parallel to the surface and resulted in flaking. In the following, the crack initiation at 2a/3 away from the center of the contact track is discussed from the viewpoint of tensile strain. Figure 4 shows the result of X-ray measurement on the contact track after the test with Pmax = 5", + " The maximum elastic tensile strain exists at y = 0.7a(= 2a/3) and x = \u00b11.2b. Then we can understand that the crack initiates at the position y = 2a/3. Namely, in the case of circular contact, the crack initiation site is not at the center of the contact track but at y = 2a/3. On the other hand, this elastic tensile strain inclines by +36\u25e6 or \u221236\u25e6 to the x axis at y = 0.7a and x = \u00b11.2b. Namely, the crack can propagate at a right angle to this strain direction in both directions, \u00b154\u25e6 to the x axis. In Fig. 3, the crack propagates nearly with this angle but in one direction. The crack propagation direction is considered to be influenced by other factors, for example, the tangential force. It has been confirmed that the crack propagation directions were opposite each other between the upper and lower specimens when there was a small slip (Muro et al. (11)). TABLE 6\u2014RESIDUAL AND ELASTIC PRINCIPAL STRAINS E1xyr , E1xy AND THEIR DIRECTIONS ON xy PLANE AT z = 0.05b FOR CIRCULAR CONTACT. E1 AND E2 MEAN MAXIMUM AND MINIMUM PRINCIPAL STRAINS, RESPECTIVELY (CALCULATING CONDITION: Pmax = 5" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002439_008-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002439_008-Figure1-1.png", + "caption": "Figure 1. The Kundt tube apparatus.", + "texts": [ + " There is a need today to educate people about the basic operating principles of radio technology. A good place to begin is in the realm of resonance phenomena\u2014a simple plastic pipe in a container of water, along with a tuning fork, can show 1/4 wavelength acoustic resonance. Producing standing waves on a piece of string can lead to a discussion of transverse and longitudinal waves and of how, with a stringed musical instrument, both are involved. To actually see the direct effects of sound waves in air, a Kundt tube experiment is very useful\u2014in the apparatus shown in figure 1, a cloth powdered with violin resin excites one half of a thin brass tube which produces longitudinal vibrations along the tube. The fundamental tone of the tube is matched to the tone of a small brass disc1\u2014the disc is in resonance with the rod and the two together make a loud high-pitch tone around 4 kHz. The disc and the rod are then moved in and out of a glass tube (closed at one end) until another level of resonance occurs\u2014the airspace resonates to produce a standing wave and the strong air vibrations are enough to move a small amount of cinnamon powder to form little groups of platelets or striations (figure 2)2" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002320_12.715862-Figure7-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002320_12.715862-Figure7-1.png", + "caption": "Figure 7: E-Flite\u2122 Ascent", + "texts": [ + " This data is converted into the Universal Asynchronous Receiver and Transmitter standard and finally into a pulse-train which the controller can understand. The signal for the resulting configuration is sent by a 2.4GHz signal, converted into RS232 standard for the computer\u2019s serial port to accept. The vehicle platform includes the plant, which is the plane with its wings, the distributed strain gauge sensor and the modified servo driver. The vehicle plant for the control is an E-Flite\u2122 Ascent electric park glider (Figure 7). The fuselage is composed of light fiberglass on which the controller for the morphing wings will be mounted. Commercially manufactured flight controls include a rudder, elevator, and 20 amp propeller motor with adjustable throttle. The wings themselves are framed by a lightweight wood and shrink wrapped in a light plastic skin. The standard bihedral wings shown below that were originally packaged with the plane were modified to a morphable trihedral configuration based on the elliptical cambered wing designs of Wiggins et al" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002034_2007-01-2860-Figure14-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002034_2007-01-2860-Figure14-1.png", + "caption": "Figure 14 \u2013 Z-axis suspension twist, 100x zoom", + "texts": [ + " b) Symmetrical vertical movement - toe The main cause of toe in symmetrical movements is the relation between the moments of inertia Iz and Ix relative to the profile\u2019s centroid. The smaller Iz is in relation to Ix, the greater the crosspiece movement will be in relation to Z-axis, generating toe angles. The relation between angular deformations is given in function of the relation between Ix and Iz, i.e., if Iz is much inferior to Ix, the greater the crosspiece will tend to move in relation to Zaxis, generating toe angles as in Figure 14. Profiles with low Iz inertia tend to have greater toe angles. c) Asymmetrical vertical movement - camber In the profile twist movement, some variables affect suspension camber behavior: 1 \u2013 C-profile border movement: The suspension asymmetrical movement provokes the main crosspiece to twist. This creates an important 6 phenomenon in U and C-profile borders. Profile twist causes opposite Y- movements in the extremity of the profile, as in Figure 15, which is essential to generate camber and toe angles" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001576_icpst.2004.1460239-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001576_icpst.2004.1460239-Figure1-1.png", + "caption": "Fig. 1. Cross sectional view of n three-phase salient pole synchronous generator.", + "texts": [], + "surrounding_texts": [ + "2004 Intematronal Conference on Power System Technology - POWERCON 2004 Slngapore, 21-24 November 2004\nA Study on Magnetic Saturation Effects in a Synchronous Generator during Unbalanced Faults\nK. W. Louie, Member, IEEE\nAbsww-This paper reviews a phase-domain synchronous generator model and investigates the effects o f the non-uniform air gap saturation on the performance of a three-phase salient pole synchronous generator during a single-phase-to-ground shart circuit. Accurate representation of magnetic saturation effects in synchronous machines is required when studying their behavior closely. Modelling a synchronous machine directly in the physical phase-domain instead of the dqo-coordinates permits an easy and accurate representation of magnetic saturation in the machine. The reviewed model has been verified to be accurate and effective in representing the behavior of synchronous generators. The test results have showed the strong impact of the non-uniform air gap on the magnetic saturation in a three-phase salient pole synchronous generator,\nIndex Terms-Phase-domain synchronous machine modelling, non-uniform air gcip saturation, unbalanced faults.\nI. INTRODUCTION ULL scale power system simulators have become F increasingly important in electric power industry. A simulator is to be continuously running and should be able to perform in \"live-mode\" a full range of power system studies, fkom optimum power flow to tramient and from dynamic stability to EIXP-type fast transient analysis. These lypes of applications have particularly strong demands on a synchronous generator model. Since the simulator is running continuously, mor accumulation inherent in the prediction schemes that are applied in conventional models is not acceptable. Also since the model should be able to simulate a wide range of system phenomena, approximations valid for some types of studies (e.g. balanced three-phase transient stabifity) may not be valid for other studies (e.g. negative sequence current effects during a line-to-ground fault).\nThe reviewed synchronous generator model is based on direct phase-coordinates, as opposed to the transformed dqoplane [l]. There is less advantage in working in the dqoplane when a full range of transient phenomena (e.g. negative sequence currents) is to be considered. On the other hand, a direct solution in the phase-plane can greatly facilitate the representation of physical machine characteristics such as saliency and saturation. From a numerical solution point of view, phase-domain modelling permits the direct interface of the generator and network representations and avoids a number of numerical predictions in the electncal variables that are needed in dqo representations. Traditionally, the modelling of saturation effects in synchronous machines has\nbeen done with the dqo analysis which uses the superposition principle to solve for the d-axis path (along the poles) and the q-axis path (in the quadrature with poles) independently [2], [3], [4], [5 ] . However, superposition is not applicable in modelling saturation effects because of the non-linearity of the saturation phenomena. According to the direction of the resultant of magnetomotive force in the air gap at each time instant of the solution, the direct phase-coordinates solution can easily and accurately represent saturation by using the correct magnetizing characteristics. In a salient pole synchronous machine, the air gap is non-uniform and the reluctance of the main magnetic flux path is a function of time. Thus the magnetic fluxes along the d-axis and the qaxis are changing and the degree of saturation along these two axes is also varying. Accordingly, the saturation along both axes should be included in the machine model. However, saturation along the q-axis in a salient pole machine is usuaiIy negIected in most of the machine models.\nThrs paper focuses an the aspect of accurate d-axis and qaxis saturation representation and its effects on unbalanced fault simulation.\nII. SYNCHRONOUS GENERATOR MODEL IN THE PHASE-DOMAIN\nA three-phase synchronous generator essentially consists of a field stsucture and a set of armature coils [l], [2]. The fieid winding on the rotor carries direct current to produce a rotational magnetic field in the air gap between the stator and the rotor. In addition to the field winding, there are a Ddamper winding and a Q-damper winding on the rotor. The three-phase windings of the armature are distributed in space 120\" apart around the periphery of the inner face o f the stator, as shown in Fig. I. These six coiIs are magneticaIly coupled. In steady state operations, voltages displaced by 120\" in phase are produced in the armature windings, regardless of the speed of the rotational field. The voitagecurrent-flux relationship in this coupled six-coil system can be expressed as [I], [2], [3], [6]\nwhere [A( t ) ] = [L( t ) l [ i ( t ) ] ; [v(t)] is the vector of the voltages across the coils; [i(t)] is the vector of the currents through the coils; [h(t)] is the vector of the flux linkages linking the coils; [RI is the diagonal matrix of the coil resistances; [L(t)] is the", + "matrix of the inductances which are functions of the rotor\nThe mechanical part of the generator consists mainly of the masses associated with the rotor, the turbine (or turbines), and the exciter machine. The equation of acceleration of the mechanical part is given by [ 11, [2], [3]\nposition e .\nwhere subscript m stands for the mechanical side of the generator; [JJ is a diagonal matrix of the moments of inertia of the masses on the shaft; [0J is the vector of the position angles of the masses; [Dm] is the matrix of the damping coefficients of the fluid around the masses; [L] is the matrix of the stiffness coefficients of the amortissew springs between the masses; [Toet(t)] i s the net torque on the shaft.\nIII. INCORPORATION OF A SATURATION MODEL\nEquation (3) gives the magnitude and the phase angIe of the resultant magnetomotive force in the machine. Once the position of the rotor represented by @(t) and the position of the total mmf represented by p(t) are known, the angle a(t) defining the direction of the main magnetic path with respect to the pole axis is calculated by\naO> = - P ( 0 (4)\nThe reluctance of the main magnetic flux path depends on the lengths of the air gap, the rotor core path, and the stator core path as shown in Fig. 3. The reluctance of the main magdetic flux path for a given angle of a can be expressed as", + "Where R,+du) is the reluctance of the rotor core path; &14(a) is the reluctance of the air gap path; bj+(a) is the reluctance of the stator core path. In Fig. 3, let us define a, as And in the ranges of -n++a, 5 o( 5 -a, and a, I a 2 n-a, we get\n(13) F , (e,a,o\nR B R A - -\nsda>\nA,(B,a, i ) = a, =\u201d(?) (6)\nIn the ranges of -a, 5 a I a,and IT-a, Ia 5 x+a, &,(a) Equations (12) and (13) are referred to as magnetizing curve generating functions. Fig. 4 shows some magnetizing curves of different magnetic flux paths generated by these functions using the d-axis (a = 0) and q-axis (a = d 2 ) data from [SI. With At being the time increment, the total normalized flux linkage at time t is given by can be expressed as 1 fL&) = R A ~ R , - \u201ca0 ) (7) where *an *Fn n,(o = ~ ~ ( ~ - A ~ ) + - [ ~ , ( ~ ) - ~ , ( ~ - ~ ) ] (14) R ~ i = +{ R,,I (0) + R,,I (;) + [ Rr,l (:) - R, (O)]} ;\nRE = [ R,I (;I - - -Rrei(0) ][\u201cs%)]\nWhere WAFn is the slope at F&) on the magnetizing clwe of the main magnetic flux path.\nIf there were no saturation, the unsaturated normalized flux linkage would become\n1 + sin(a ) 1 - \u201d:a J RA = RA,[ -21 ;\n(15) 1 - S W Y , ) A,-u = LF, (0\nSimilarly, in file ranges of -x+a, I a I -ao and a, g U 5 where L 1s the unsaturated inductance. C o m b d g (l), (14), and (1 5) gives x-a, &,(a) can be found by\nA Knowing the reluctance of the main magnetic flux path, the total magnetic flux on a winding of N turns is given by a, = ;1. [L(t)X[i(t)]- [i(t - A t ) ] } ;\nAF.L\nF(B,a,i) A(B,a , i) (9) [L(t)] is the matrix of the coil inductances. The self and\nmutual inductances across the stator and rotor windings in the matrix b(t)] are functions of the rotor position 0.\n-- #(@,a, i) = Rrd (a) N\nThe total flux linkage and the total magnetomotive force can be normalized with respect to the number of tums of phase-a winding and that of the field winding as\nIV. CASESTUDY A test case of a single-phase-to-ground fault in phase-a at the terminal of a three-phase salient pole sy&hronous generator has been considered. An analytrcal solution of the short circuit can be performed using the technique of symmetrical components. During the fault, the terminal voltages of the healthy phases and the current of the faulted phase in terms of phasor quantities are given by (10) w, a, i) It,(O,a,i) = - FlOY a, i) F, (e, a, i ) = N , (1 1) Nf\nwhere subscript 11 stands for \u201cnormalized\u2019\u2019. From (7), (8), (9), n+a, we have (lo), and (11) in the ranges of -a, I aI a, and %-ao I a 5 v, =vu[(c12-u)z,+(~~-1)(x,-2.u,)] Z (17)\nv, = c [ a - a 2 ) z l Z +(o-1)(x,2his)] (18) (12) F (@,a,i)\nRB R A - -\nW a o )\ni.,(@,a,i) =\n(19) 3 v o I =-\na Z" + ] + }, + { + "image_filename": "designv11_69_0001731_iemdc.2005.195837-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001731_iemdc.2005.195837-Figure2-1.png", + "caption": "Fig. 2. Vector diagrams for asynchronous and synchronous motors", + "texts": [ + " Outputs of this part of the algorithm are current references iGs\u03b1, iGs\u03b2, in stator frame \u03b1, \u03b2. FOC is oriented on direction of definite flux vector \u03c8c. Element FM forms feedback signal \u03c8c \u2013 module of this flux vector and direction vector dc = (cos \u03b3c, sin \u03b3c) for this flux in stator frame. Controlled flux is defined usually through stator flux and current vectors: \u03c8c = \u03c8s \u2013 Le is. (1) Equivalent inductance Le is accepted usually for asynchronous drive as Le \u2248 L\u03c3 = Ls\u03c3 + Lr\u03c3, and they consider controlled flux as rotor flux. Vector diagram is represented on Fig. 2. Controlled flux is defined by magnetizing component of stator current is1. Correspondingly flux regulator FR forms reference for magnetizing component of stator current iGs1. 0-7803-8987-5/05/$20.00 \u00a92005 IEEE. 957 Torque component of current reference iGs2 is formed from torque reference. Thus current reference is formed in axes 1, 2 oriented on flux vector. In synchronous motor with electromagnet excitation torque component of stator current influence strongly on flux, and autonomous control of flux in flux frame is impossible one. Tasks of this work are as follows: investigation of synchronous drive specifics with FOC; finding and tests of simple solution for improvement of FOC for synchronous drive. We are urged to consider module of main flux (air gap flux) vector as control variable of the flux control loop for synchronous drive. Vector diagrams for steady-state regime of synchronous motors are shown on Fig. 2. Designations are as follows: d, q \u2013 rotor axes; is - stator currents vector; \u03c8s - stator flux linkages vector; \u03c8\u03b4 - vector of main magnet flux (air gap flux); \u03c1 - direction of vector \u03c8\u03b4; \u03c4 - orthogonal leading direction to vector \u03c8\u03b4; \u03d1 - angle between vector \u03c8\u03b4 and axis d; is\u03c1, is\u03c4, - magnetizing and torque components of stator current, if \u2013 excitation current, Ls\u03c3 - leakage inductance of stator, Rcd, Rcq \u2013resistances of damping loops in d, q axes. Correspondingly in relation (1) \u03c8c = \u03c8\u03b4 and Le = Ls\u03c3 " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003435_12.841847-Figure4-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003435_12.841847-Figure4-1.png", + "caption": "Figure 4. Schematic view of the fiber sensor (left) and photograph of the refixed fiber (right). The two facets of the fiber are glued into a hull with a circular opening.", + "texts": [ + " Photonics GmbH, Berlin, Germany) were machined to insert cavities with a path length of less than 100 \u00b5m. The refractive indices of the core and cladding were nCore = 2.16 and nClad = 2.14, respectively. In our experiments we used mechanically machined fibers that we cut off in half as a first step. The gap was created by mechanically sawing \u2248 30 \u00b5m deep into one of the two front facetts using a 300 \u00b5m wide blade. Afterwards, a hull with an inner diameter of 500 \u00b5m was manufactured to adjust and fix the two facetts back together again by glueing the hull onto the ends (see figure 4). For in-vitro experiments, this fiber sensor was then positioned in a temperature stabilized flow-through chamber with a volume of 2.6 mL. The chamber was combined with a programmable valve control and a pump to automatically run measurements with multiple glucose, reference and flush solutions. The flow rate for the glucose solutions was set to 10 mL/min. Mid-infrared light which originated from a quantum cascade laser was sent through the fiber and the transmitted light was detected by a pyroelectric detector using a lock-in technique" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002172_1.3616681-Figure6-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002172_1.3616681-Figure6-1.png", + "caption": "Fig. 6 (Cont.)", + "texts": [ + " The steady-state response of the system was first calculated assuming the gear to be the generator and Q = 37 for each individual inertia in the system except the gear, which was assumed to have Q = On this basis the system Q, was found to be 370, so obviously the only significant source of damping must be the gear. When the same value of Q is assigned to all individual inertias in the system, Q, is equal to the individual values of Q, and since system elements other than the gear are not expected to have a Q as low as 37, the maximum possible Q of the gear is 37. The calculation was repeated with Q = 34 for the gear and <*> for all other inertias, giving a Qs of 37. Therefore it is determined that the gear damping corresponds to a Q of 34 to 37 in this instance. The table in Fig. 6 gives the system constants and results of the calculations for the second critical frequency for each individual Q = 100 except Q = 35 for the gear, based on an input J U L Y 1 9 6 7 / 15 Downloaded From: http://gasturbinespower.asmedigitalcollection.asme.org/ on 01/29/2016 Terms of Use: http://www.asme.org/about-asme/terms-of-use Syste.- Consian-.s 1 0 ' \u00b0 k a Compressor Stage IS037 579 1 2 3 8 . 2 \u2022i \u2022\u2022 15901 33k llll)-6 \u2022i lll)21 2tO nH.6 .1 ii 10839 227 997-31 \" \" 19013 1)00 1 0 6 6 . 3 \u2022i \u2022\u2022 19293 1)07 1 0 6 6 . 3 19293 107 1 0 6 6 . 3 \" \" 22758 576 2 3 . 7 0 7 Coupling 507 1, 1 6 . 9 3 8 1/2 Gear tllio 2i)9 Turbine Siege 1097lf 230 990.47 \u2022\u2022 .i 10802 2 2 7 297-51 \" 11 12187 2 5 S 1)62.11 13301) 260 3 8 6 . 2 3 \" \" 15262 328 1)07,- 21 17293 362 10.279 0 . 6 Coupling 275 6.827 1/2 Gear IflltO 2'49 Fig. 6 generated torque of unity. The gear deflection is calculated to be 0.00014 deg per in-lb of generated torque. Since the measured gear deflection was 0.0244 deg single amplitude, the peak alternating torque generated by the gear is 0.0244/0.00014 = 175 in-lb. T h e vector diagram in Fig. 7 shows some phase relationships near the second critical. Calculations were also made to determine whether the compressor could have been the source of torque generation, assuming each individual Q = 37, showing that this centrifugal compressor would have to generate an alternating torque greater than its rated full load torque in order to explain the measured gear deflection" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000976_cp:20040396-Figure7-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000976_cp:20040396-Figure7-1.png", + "caption": "Fig. 7 Typical flux distributions related to 12-slot, IO-pole motor.", + "texts": [], + "surrounding_texts": [ + "The superposition method is now extended to sum thr contributions of each stator slot to the resultant cogging torque. The two IO-pole motors.whose parameters are given in Table I , and which are assumed to have a pole-arc to polepitch ratio of 1 .O, are considered The predicted results for the IO-pole, 15-slot motor are shown in Figs. 5 and 6. The cogging torque waveform which results with a 3-slot stator is synthesized from that which results with a single slot stator by phase shifting its cogging torque waveform by 0, I20 and 240 degrees mechanical and summing, as shown in Fig. 6a. 111 a similar manner, the cogging torque waveform which results with the 15-slot stator is obtained, Fig. 6b. It will be noted that when there is only one slot on the stator of a 10-pole motor, the least common multiple N, is IO, while N, is 30 for both a 3-slot stator and a 15-slot stator. Hence, the periodicity of the cogging torque waveform for the I-slot stator is 36 mechanical degrees, while that for the 3-slot and 15-slot stators is 12 mechanical degrees, as confirmed by the finite element predictions shown in Fig. 6. It can also be seen that excellent agreement is achieved between the synthesized and the resultant finite element calculated cogging torque waveforms. The same synthesis procedure has been applied to the 10- pole, 12-slot motor, as shown in Figs. 7 and 8, in which the cogging torque which results with a .2-slot stator is also shown. The least common multiple N, is now I O for both the I-slot and 2-slot stators, while for the 12-slot stator N , is 60. Therefore, the periodicity of the cogging torque waveforms, which result with the I-slot and 2-slot stators is the same, viz. 36 mechanical degrees, while that for the 12-slot stator is only 6 mechanical degrees. Again, excellent agreement is achieved between the analytically synthesized and the resultant finite element calculated cogging torque waveforms. 0 10 20 30 4a 1\" fQ A o ~ u l s r Polifion (harat) (a) I-magnet and 2-magnet rotors I 0 6 l i 18 21 10 36. *\"gular Position (Degree, (a) I-slot and 3-slot stators 832 Since the width of the stator slot openings is usually much smaller than the slot-pitch, the interaction between two adjacent slots is usually negligible. Hence, in general, the resultant cogging torque waveform can be synthesized from the cogging torque waveform which results with a single slot to a high accuracy. 4 Comparison of Predicted and Measured Cogging Torque Waveforms In order to validate the accuracy of the foregoing calculations, cogging torque waveforms were measured on a 4-pole, 6-slot motor, Fig.9. Fig.10 compares the predicted and measured cogging torque waveforms when the motor is equipped with each of the rotors shown in Figs.9b-d for which the magnet pole-arc to pole-pitch ratio is 0.7, as well as with rotors having a pole-arc to pole-pitch ratio of 1.0, the corresponding flux distributions being shown in Figs. 1 and 3. It should be noted that the predicted results are deduced from finite element analysis, rather than being synthesized. As can be seen, excellent agreement is achieved between predictions and measurements for both pole-arc to pole-pitch ratios, which confirms the discussion presented in section 2. 20 *\u201clYI., P..ili\u201d. , \u201d C s r r , (a) Pole-arc to pole-pitch ratio=0.7 *\u201cI\u201c,.. P O m m l\u201d 0 (28) ( 0 kp\u201d which is reached for k;k; i j > - and kz > 0 (29) k: + kz r From a physical point of view, the first inequality in (29) can be interpreted as a condition on the serial combination of the two tangential stiffnesses k i and ki" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003224_icsmc.2010.5642211-Figure18-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003224_icsmc.2010.5642211-Figure18-1.png", + "caption": "Fig. 18. Experimental electric wheelchair", + "texts": [], + "surrounding_texts": [ + "In this paper, the obstacle avoidance assist system with MVFH using reinforcement learning has been proposed, and the effectiveness has been evaluated by computer simulations. As a result, it has been shown that the proposed scheme is able to adjust to different user conditions. Moreover, the proposed scheme was applied to an experimental electric wheelchair, and the effectiveness of the proposed technique was verified in a real operating environment. The experiment verified that an electric wheelchair took safe actions in various environments." + ] + }, + { + "image_filename": "designv11_69_0003962_oceans.2009.5422068-Figure4-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003962_oceans.2009.5422068-Figure4-1.png", + "caption": "Fig. 4. Motion ranges of the ORION 7P [12]", + "texts": [ + " The workspace can be saved after the creation and the data used for the online adaptation. The workspace itself is discretised. Since the real dimension of the respective SD \u2212 WS was prior unknown, we used the parameters shown in table I. As the resolution for the cartesian and angular discretisation we used values which are below the precision achievable by the ORION 7P master-controller. The calculations have been done offline on an INTEL XEON ?? dual processor system using all 16 virtual cores. Each calculation lasted approximately 3.5h. In Fig. 4 the angular range of the ORION 7P is depicted. These graphics show the \u201dreal\u201d workingspace of the arm, without respect to dexterity. In Fig 5 two analysis of the SDWS. Depicted are SDWSFront and SDWSFrontDown. The graphs show the ratings in form of a color gradient. Looking at these graphs it is quite obvious why a automatic dexterity control system is needed. The show how much the dexterity of the system degrades inside its working space. For an operator this proves to be a serious limitation. In this section we will introduce an automatic system which allows an operator to get information about the current dexterity and to move the ROV accordingly, using the results of the previous analysis" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001292_0020-7403(71)90069-5-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001292_0020-7403(71)90069-5-Figure1-1.png", + "caption": "FIG. 1. G e o m e t r y of c l amped shal low conical shell.", + "texts": [ + " Stephens and Fulton la studied the axisymmetrical buckling behavior of clamped spherical caps under a step load of infinite duration which is distributed over a sub-region of the cap surrounding the apex. The axisymmetrical dynamic buckling behavior for clamped shallow conical shells under step loads of infinite duration with various axisymmetrical spatial distributions can be studied effectively using the numerical scheme devised by Huang. TM S T A T I C B U C K L I N G The b e h a v i o r of a t h i n c l amped shal low conical cap {Fig. 1) u n d e r g o i n g m o d e r a t e l y large deflect ions due to a n a x i s y m m e t r i c a l load can be descr ibed b y Margue r re ' s equa t i ons 14 in n o n d i m e n s i o n a l fo rm V4 w A .,, /1 , 1 ..\\ ,, /1 , 1 . \\ ,, 1 , l --.x: \u00f7tJ \u00f7ix -'(;/-J) w i t h b o u n d a r y cond i t ions a t t h e c lamped edge (x =/~) W ---- W' -~ O, z ]: P''\" f - ~ f ' - ~ f - - O, / ,, v v . . \\ ' 1 , 1 . . ,, 1 . . \" -::'-J) -x: -\u00a2\u00f7': + '\" (::) -- o. (3, 4) (5) (6) I n these express ions a, H a n d h are, respect ive ly , t he base p l ane radius , apex rise a n d shel l th ickness ", + " T h e load ing func t i on F 1 a p p e a r i n g in e q u a t i o n (1) is g iven b y t he re la t ions P ~(x--xo) 3(1 -~e ) P a 2 r 1 = w i t h p -- a n d P = 21rroP~ (8a) x ~ E H h a The n o n d i m e n s i o n a l q u a n t i t i e s a p p e a r i n g in equa t ions (1)-(6) are rela~ed to t he correspond ing phys ica l quml t i t i e s t h r o u g h the re la t ions x = - r , A = 213( i -v~) ] ,' w = W, f = (7) a - ~ 4 E H ~ h\" for the ring load and :F 1 = 4 p [ h ( x - x l ) - h ( x - x 2 ) ] with p - 3 2 E H a h q (Sb) for the band type load. In equations (8a) and (8b) xo, x~ and x~ are the nondimensional forms of r o, r, and r 2, respectively, which are described on Fig. 1. P~ is the intensity of the ring load and q is the density of the distributed load. Finally, the Dirae delta and the Heaviside flmetions appearing in equations (Sa) and (8b) are defined by the relations ~ ( x - ~ \u00b0 ) = t 0, ~' \u00a2 x0, and h(x-xi) = tO' x 0 and \u03c6 < 0, if the particle moves in an uphill and downhill direction, respectively. We find the equations of motion for the speed v and the y-component of the velocity: m dv dt = (\u2212S \u2212 F sin \u03c6), (1) m dvy dt = (\u2212F \u2212 S sin \u03c6). (2) For zero friction, S = 0, we have sin \u03c6 dvy/dt = dv/dt . Since sin \u03c6 = vy/v, it follows vy dvy/dt = v dv/dt , so that dvx/dt = 0, or x = vx,0t . This case corresponds to simple projectile motion of the particle, and we recover the familiar result that the motion along the x- and y-directions are independent of each other" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003761_15421400903112077-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003761_15421400903112077-Figure1-1.png", + "caption": "FIGURE 1 Bent core molecule.", + "texts": [ + " The liquid crystalline material is in a smectic C phase. As such there is an energetic cost to smectic layer bending. If the term in the energy corresponding to this is comparable to the Frank\u2013Oseen energy we prove that the circular configurations are stable. If however the smectic layer energy is the dominant elastic contribution we prove that circular equilibria become unstable relative to deformations with small surface undulations. Bent core liquid crystal molecules are banana-shaped as in Figure 1 and are described by two orthonormal vectors n and p. We view the molecule as two dimensional lying in the molecular plane spanned by n and p. Here n is parallel to the chord connecting the two ends of the molecule and p points toward the kink in the molecule from the midpoint of the chord. Because of its shape the molecule carries a microscopic polarization parallel to p. We do not distinguish between the microscope fields n and p and the corresponding macroscopic ones representing their local averages" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003409_j.cclet.2010.01.025-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003409_j.cclet.2010.01.025-Figure1-1.png", + "caption": "Fig. 1. Scheme of FI-CL system. (a) Sample or blank (pure water) solution; (b) CuSO4; (c) luminol; (d) H2O2; P1 and P2, peristaltic pumps; S, switching valve; Y1 and Y2, confluence points; F, flow cell; W, waste water; PMT, photomultiplier tube; PC, personal computer; and NHV, negative high voltage.", + "texts": [ + " All the chemicals were of analytical-reagent grade. The pure water (18.2 MV cm) was processed with an Ultrapure Water System (Kangning Water Treatment Solution Provider, China). * Corresponding author. 1001-8417/$ \u2013 see front matter # 2010 Yan Ming Liu. Published by Elsevier B.V. on behalf of Chinese Chemical Society. All rights reserved. doi:10.1016/j.cclet.2010.01.025 The FI-CL cell was constructed in combination with a model IFFM-E FI-CL system (Xi\u2019an Remax Analytical Instrument Co. Ltd., China) for this work as shown in Fig. 1. UV\u2013vis spectra were taken by the Uvmini-1240 UV\u2013vis spectrophotometer (Shimadzu, Japan). Our experimental results showed that 20 kinds of amino acids were found to inhibit the CL signal of luminol\u2013 H2O2\u2013CuSO4 system. Na2CO3\u2013NaHCO3 was chosen as optimal alkaline medium in luminol solution. Other conditions such as luminol, H2O2, CuSO4, Na2CO3\u2013NaHCO3 concentrations, pH of luminol solution and the flow rates were optimized as follows: 9.0 10 5 mol/L luminol prepared in 10 mmol/L Na2CO3\u2013NaHCO3(pH 9" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002122_07ias.2007.342-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002122_07ias.2007.342-Figure2-1.png", + "caption": "Fig. 2. Space vector diagram of vector-controlled motor-drive system for a case of inter-turn short circuit in one of the stator winding fault (\u03c9s is the synchronous speed).", + "texts": [ + " Meanwhile, the backward component of the stator flux space-vector, b\u03c8 , should be adjusted to cancel or mitigate the ac torque ripples produced by the second and the third term in (5) by adjusting the backward component of the applied stator voltage spacevector, bv where s f bv v v= + . At this stage, one should notice that the compensation action of the drive depends on the tuning status of its controller. The loci of the voltage space-vector, sv , stator current spacevector, si , and the stator flux space-vector, s\u03c8 under faulty conditions are depicted in Fig. 2. III. THE CONCEPT OF THE FLUX PENDULOUS OSCILLATION The concept of the magnetic field pendulous oscillation (MFPO) has been previously presented in several publications [5], [6]. This MFPO phenomenon can best be defined as the fluctuation of speed of rotation of the resultant magnetic field with respect to a synchronously rotating frame of reference, due to asymmetry either in the stator and/or the rotor circuits, which also produces torque ripples that might degrade system performance. These oscillations in the resultant rotational magnetic field are measured with respect to the synchronously rotating voltage space-vector as a reference" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001576_icpst.2004.1460239-Figure3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001576_icpst.2004.1460239-Figure3-1.png", + "caption": "Fig. 3. Main magnetic flux path of a three-phase salient pole s~chrouous genentor.", + "texts": [ + " INCORPORATION OF A SATURATION MODEL Equation (3) gives the magnitude and the phase angIe of the resultant magnetomotive force in the machine. Once the position of the rotor represented by @(t) and the position of the total mmf represented by p(t) are known, the angle a(t) defining the direction of the main magnetic path with respect to the pole axis is calculated by aO> = - P ( 0 (4) The reluctance of the main magnetic flux path depends on the lengths of the air gap, the rotor core path, and the stator core path as shown in Fig. 3. The reluctance of the main magdetic flux path for a given angle of a can be expressed as Where R,+du) is the reluctance of the rotor core path; &14(a) is the reluctance of the air gap path; bj+(a) is the reluctance of the stator core path. In Fig. 3, let us define a, as And in the ranges of -n++a, 5 o( 5 -a, and a, I a 2 n-a, we get (13) F , (e,a,o R B R A - - sda> A,(B,a, i ) = a, =\u201d(?) (6) In the ranges of -a, 5 a I a,and IT-a, Ia 5 x+a, &,(a) Equations (12) and (13) are referred to as magnetizing curve generating functions. Fig. 4 shows some magnetizing curves of different magnetic flux paths generated by these functions using the d-axis (a = 0) and q-axis (a = d 2 ) data from [SI. With At being the time increment, the total normalized flux linkage at time t is given by can be expressed as 1 fL&) = R A ~ R , - \u201ca0 ) (7) where *an *Fn n,(o = ~ ~ ( ~ - A ~ ) + - [ ~ , ( ~ ) - ~ , ( ~ - ~ ) ] (14) R ~ i = +{ R,,I (0) + R,,I (;) + [ Rr,l (:) - R, (O)]} ; RE = [ R,I (;I - - -Rrei(0) ][\u201cs%)] Where WAFn is the slope at F&) on the magnetizing clwe of the main magnetic flux path" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003926_s00707-009-0150-y-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003926_s00707-009-0150-y-Figure2-1.png", + "caption": "Fig. 2 Momentum sphere illustrating the heteroclinic orbits and the hyperbolic saddle points, curves are orbits of constant energy", + "texts": [ + " (27)\u2013(33) is found by setting \u03b5 = 0. Doing so eliminates the dependence of Eqs. (27)\u2013(29) on Eqs. (30)\u2013(33) and allows us to solve Eqs. (27)\u2013(29) independently of Eqs. (30)\u2013(33). The unperturbed system of equations corresponding to Eqs. (27)\u2013(29) are given by h\u0303\u2032 1 = 1 \u2212 r1 r1 h\u03032h\u03033, (34) h\u0303\u2032 2 = r1 \u2212 r2 r2r1 h\u03031h\u03033, (35) h\u0303\u2032 3 = r2 \u2212 1 r2 h\u03031h\u03032, (36) and are identical to Euler\u2019s rotational equations of motion for a torque-free rigid body. The phase space for the unperturbed system is a sphere shown in Fig. 2, where the non-dimensional body-fixed angular momentum components h\u03031, h\u03032 and h\u03033 are the phase variables. The phase space has six equilibrium points at {(\u00b11, 0, 0), (0, \u00b11, 0), (0, 0, \u00b11)} where the equilibrium points (\u00b11, 0, 0) and (0, 0, \u00b11) are neutrally stable centers corresponding to minor and major axis spin, respectively, and the equilibrium points at (0, \u00b11, 0) are unstable hyperbolic saddle points corresponding to intermediate axis spin. Note the presence of heteroclinic orbits jointing the pair of saddle points", + " (30)\u2013(33) so that we may substitute for \u03b1\u2032, \u03c3\u0303i into the perturbation terms of Eqs. (27)\u2013(29). The solution along the heteroclinic orbits can be found in terms of hyperbolic trigonometric functions (see [1]). These solutions are given as follows: h\u03031 = s1 X1 sech(d\u03c4), (40) h\u03032 = s2 tanh(d\u03c4), (41) h\u03033 = s3 X3 sech(d\u03c4), (42) where X1 \u221a r2(1 \u2212 r1)/(r2 \u2212 r1), X3 \u221a r1(r2 \u2212 1)/(r2 \u2212 r1), d \u221a (r1 \u2212 1)(1 \u2212 r2)/(r1r2), and s1, s2, s3 are each \u00b11 and are chosen such that the product s1s2s3 = 1 (these permutations give all four of the heteroclinic orbits shown in Fig. 2). We can solve Eqs. (31)\u2013(33), with \u03b5 = 0, for \u03c3\u0303i by substituting in the unperturbed solutions for h\u0303\u2032 i : \u03c3\u03031 = \u2212 J\u0303 h\u0303\u2032 1 r2\u00b5\u0303 , (43) \u03c3\u03032 = \u2212 J\u0303 h\u0303\u2032 2 \u00b5\u0303 , (44) \u03c3\u03033 = \u2212 J\u0303 h\u0303\u2032 3 r1\u00b5\u0303 . (45) We can solve Eq. (30), with \u03b5 = 0, for \u03b1 by substituting in the unperturbed solutions for h\u03032 and h\u03033. The approximate solution to this equation is derived by Gray et al. [1] to be \u03b1 = A sin( \u03c4 + ), where the square of the amplitude A is given by A2 = [ \u03b10 \u2212 \u03c0C 2d2 sec \u03c0 2d ]2 + { \u03b1\u2032 0 + C d + C 2d2 [ \u03c0 tanh \u03c0 2d + i ( d + i 4d ) \u2212 i ( d \u2212 i 4d )]}2 , (46) the tangent of the phase angle is given by tan = [ \u03b10 \u2212 \u03c0C 2d2 sec h \u03c0 2d ]/ { \u03b1\u2032 0 + C d + C 2d2 [ \u03c0 tanh \u03c0 2d + i ( d + i 4d ) \u2212 i ( d \u2212 i 4d )]} , (47) where \u03b10 is the initial angle of twist of the torsional appendage, \u03b1\u2032 0 is the initial twist rate, is the frequency of the appendage oscillation defined by \u221a K\u0303/(\u03bbG1r3) and \u03bb = md2/(ml2)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002354_tac.1966.1098269-Figure4-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002354_tac.1966.1098269-Figure4-1.png", + "caption": "Fig. 4. Comparison of original nonlinear function f(u) with new nonlinear function F ( i ) .", + "texts": [ + " 3 becomes unstable at points where - < - df -1 do B with the result that F ( Z ) may be discontinuous and multiple-valued. X proof of the theorem is given in the Appendix. Thus, if the condition associated a i th property 3 ) of the theorem is satisfied, the new nonlinear function F(6) will possess all the essential mathematical properties required for application of the results given in [ 3 ] and [4]. In particular, if B 2 0 . the function F ( 5 ) will always lie in the sector of the a, F(a)-plane that is bounded by the lines2 F(2) = 0 and F(i) = This latter relationship is illustrated in Fig. 4. This procedure also makes it possible, subject to the usual restrictions on A, to transform the state variable equations for the extended direct control form to the so-called Lur\u2019e First Canonic Form [ 2 ] . The recent results of Popov [SI, [ 6 ] have provided new techniques for studying the absolute stability of nonlinear control systems that can be represented by the feedback circuit shown in Fig. 1. In particular, Popov has given sufficient conditions for absolute stability which apply to a variety of special cases in which wz IWIl and 11 > I Wzl, (4) 4 and n will be of the same sign" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002614_etfa.2008.4638372-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002614_etfa.2008.4638372-Figure2-1.png", + "caption": "Figure 2. Launching of Objects", + "texts": [ + " For the calculation of a trajectory at first the flight time from the launching position to the apex must be determined: (7) Afterwards the trajectory can be calculated as (8) (9) This model has the advantage, that the launching parameters \u03b10 and v0 can be calculated for a predefined position of an apex A (xA, yA) very easily with the following equations: (10) (11) (12) (13) (14) Two different objects, a tennis ball and a pipe, were used to compare model 1 with real trajectories. Therefore these objects were thrown with a prototype of a throwing device as it is shown in Figure 2. By stretching the spring of the throwing device with different distances, different speeds v0 can be set. In the experiments the actual values of v0 were measured with two light barriers and the actual values of \u03b10 were measured with a high speed camera with 1000 fps (see Figure 2). A second high-speed camera was used to capture the trajectories of the thrown objects in a range 1,7 m \u2022 x\u2022 3 m (Figure 3 and 4). The positions )sin2( sin 0 2 0 22 0 \u03b1 \u03b1 \u22c5\u22c5+\u22c5 \u22c5 = o A vkg v y 0tan \u03b1 AE A yx x \u22c5 = xyxxyx xxxy xy AE E AE EA \u22c5\u22c5\u22c5\u2212+\u22c5 \u2212\u22c5\u22c5 = ) tan 2( tan )( )( 00 \u03b1\u03b1 )1arccos(1 2 Aykg A e kg t \u22c5\u22c5\u2212\u22c5 \u22c5 = x y yx Ty gkx T x v v vvv kgt k v e tk v T ,0 ,0 0 2 ,0 2 ,00 ,0 ,0 arctan )()( )tan(1 )1(1 = += \u22c5\u22c5\u22c5= \u2212\u22c5 \u22c5 = \u22c5\u22c5 \u03b1 )cos1ln(1)( 00 tkgv kg tx \u22c5\u22c5\u22c5\u22c5+\u22c5 \u22c5 = \u03b1 kg kv t A \u22c5 \u22c5\u22c5 = )sinarctan( 00 \u03b1 \u23aa \u23aa \u23aa \u23aa \u23ad \u23aa\u23aa \u23aa \u23aa \u23ac \u23ab \u23aa \u23aa \u23aa \u23aa \u23a9 \u23aa\u23aa \u23aa \u23aa \u23a8 \u23a7 >\u22c5\u22c5\u22c5 +\u22c5\u2212\u22c5\u2212\u22c5\u2212\u22c5 \u22c5 \u2264\u22c5\u22c5\u2212 \u22c5\u2212\u22c5\u22c5 \u22c5 = \u22c5\u2212\u22c5\u22c5\u2212 AA kttg A AA A ttforktg ekttg kg ttforktg kttg kg ty A ))))cos( 1( 2 1ln()((1 ))cos(ln ))(cos((ln1 )( )(2 of the objects R\u2019i(x\u2019R,i, y\u2019R,i) at first were determined as they could be seen on the images in the grid" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003130_978-90-481-9884-9_25-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003130_978-90-481-9884-9_25-Figure2-1.png", + "caption": "Fig. 2 Position of the centre of mass of the cylinder i", + "texts": [ + " (21) Vector P FP(kin)|B represents the force vector acting on the centre of mass of the moving platform, and P MP(kin)|B represents the moment vector acting on the moving platform, expressed in the base frame, {B}. From (19) it can be concluded that two matrices playing the roles of the inertia matrix and the Coriolis and centripetal terms matrix are: IP |B .T, (22) d dt ( IP |B .T ) . (23) If the centre of mass of each cylinder is located at a constant distance, bC , from the cylinder to base platform connecting point, Bi (Fig. 2), then its position relative to frame {B} is: BpCi |B = bC.l\u0302i + bi , (24) where l\u0302i = li \u2016li\u2016 = li li , (25) li =B xP(pos)|B +P pi|B \u2212 bi . (26) The linear velocity of the cylinder centre of mass, B p\u0307Ci |B , relative to {B} and expressed in the same frame, may be computed as: B p\u0307Ci |B =B \u03c9li |B \u00d7 bC.l\u0302i , (27) where B\u03c9li |B represents the leg angular velocity, which can be found from: B\u03c9li |B \u00d7 li =B vP |B +B \u03c9P |B \u00d7P pi|B . (28) As the leg (both the cylinder and piston) cannot rotate along its own axis, the angular velocity along l\u0302i is always zero, and vectors li and B\u03c9li |B are always perpendicular" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001981_s0091-679x(08)60713-x-Figure8-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001981_s0091-679x(08)60713-x-Figure8-1.png", + "caption": "FIG. 8. Diagram of a multiparameter reaction chamber which permits PO,, pH, and photometric changes to be studied simultaneously in a stirred medium. The addition of an ion-specific electrode (e.g., Ca2+) to the filling port enables four types of measurements to be made.", + "texts": [ + " 7, was to mount a small diameter combined pH electrode directly in the oxygen electrode holder. Although some care is required in the choice of pH electrode, several designs are available that have a more or less linear configuration and a long enough stem (such as the Thomas 4858) so that they can fit into a hole machined in the oxygen electrode carrier. Brierley (1969) used a reaction vessel which fits into an Eppendorf photometer that can monitor pH, oxygen uptake, and swelling of mitochondria (Fig. 8) . The addition of an ion-specific electrode to the chamber makes it possible to measure simultaneous changes in four different characteristics of the system. Although this chamber was designed for mitochondria1 work, it can be adapted to the measurement of either plant or animal cells, and other cell fractions. An interesting development is the construction of an oxygen electrode reaction vessel which can be used on a microscope stage. Rikmenspoel et al. (1969) designed a chamber in which simultaneous measurements of pH, oxygen uptake, sperm flagella motility, and fructolysis could be 6" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000897_cira.2003.1222216-Figure3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000897_cira.2003.1222216-Figure3-1.png", + "caption": "Figure 3: Kinematic structure and frame position of the 5 DOF robot.", + "texts": [ + " The order chosen in the presentation of the methods follows the effective temporal sequence of the development of the procedures. 3.1 Parametrical calibration The parametrical calibration is based on the develop ment of a kinematics model of the robot, to relate the joint rotations with the pose of the end-effector taking into account the robot inaccuracy. The nominal direct kinematic model can be represented as: S = F (Q, L,) ( 1 ) where S is the gripper pose, Q = [q l , q2 ,...I are joint coordinates and L, is the nominal set of ths robot parameters. The kinematic structure of the robot is depicted in figure 3, where the positioning of the frames (following D&H rules), the joint angles (9;) and the geometric p\u201d rameters are highlighted. For the considered measuring robot the goal is to relate the forced pose Sd to the actual joint coordinates Q. measured by the encoders: s d = F(Qa,Ln +AL.) (2) where AL describes the robot inaccuracy. Shortly, with reference to figure 4, the value of A S = Sd - F(Q., Ln) is used to identify the value AL for which eq.2 is satisfied. The parametrical model implemented (better described in (141) was developed using a modified D&H parametrization [7]; 25 structural parameters were considered, including joint angles offset, link length errors, axis misalignment and robot base offset" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003226_s1068366610010010-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003226_s1068366610010010-Figure2-1.png", + "caption": "Fig. 2. Pressure distribution in cross section y = 0 at Sr = 7.6 \u00d7 10\u20137; S = 4.6 \u00d7 10\u20135 and different values of parameter F: 1\u2014F = 4.38 \u00d7 10\u20136; 2\u20148.75 \u00d7 10\u20136; 3\u20141.72 \u00d7 10\u20135; 4\u20142.61 \u00d7 10\u20135; 5\u20143.48 \u00d7 10\u20135; 6\u2014F = 4.36 \u00d7 10\u20135.", + "texts": [ + " The depen dence on the values F at constant values Sr, S reflects h i 1 2 + j, 3 pi 1+ j, pi j,\u2013 \u0394xi + h i 1 2 \u2013 j, 3 pi j, pi 1\u2013 j,\u2013 \u0394xj \u2013 \u2013 \u0394xi + \u0394xi \u2013+ \u0394yj + \u0394yj \u2013+ + \u00d7 h i j 1 2 +, 3 pi j, 1+ pi j,\u2013 \u0394yj + h i j 1 2 \u2013, 3 pi j, pi j 1\u2013,\u2013 \u0394yj \u2013 \u2013\u239d \u23a0 \u239b \u239e \u2013 6 S \u03c12 h i 1 2 + j, h i j 1 2 \u2013, \u2013\u239d \u23a0 \u239b \u239e 0, 1 i Nj, 1 j M,< << <= h i 1 2 + j, 1 2 = h i 1 2 \u2013 j, 1 2 = h i j 1 2 +, 1 2 = h i 1 2 \u2013 j, 1 2 = \u0394xi + \u0394xi \u2013 \u0394yj + \u0394yj \u2013 hi j, \u0394\u2013 1 \u03c1 xi 2 yj 2+( ) 1 \u03c1 F pi j, a 2 pk 1\u2013 j, pk j,+( ) k i 1+= N \u2211+ \u23a9 \u23a8 \u23a7 + += \u00d7 Sr S xi xk 1\u2013\u2013( )\u2013\u239d \u23a0 \u239b \u239eexp Sr S xi xk\u2013( )\u2013\u239d \u23a0 \u239b \u239eexp\u2013 \u23ad \u23ac \u23ab . pNj 1\u2013 j, pNj j, , j 2 \u2026 M 1.\u2013, ,= = \u0394, the dependence on the viscoelastic layer compliance, while the dependence on the value S at the constant values Sr, F reflects the dependence on the sliding velocity. Let us consider the effect of each of the above parameters individually. Figure 2 shows the function of pressure distribution in the cross section y = 0 at the values Sr = 7.6 \u00d7 10\u20137; S = 4.6 \u00d7 10\u20135 and different values of the parameter F. It follows from the above data that the high pressure region expands when F grows. Meanwhile the maxi mum pressure drops significantly. The high pressure region grows predominantly towards the inlet bound ary of the lubricating film. Meanwhile, the middle point of the high pressure region shifts leftwards, while the maximum pressure point shifts leftwards much more than the middle point" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002765_09544100jaero169-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002765_09544100jaero169-Figure1-1.png", + "caption": "Fig. 1 The electrohydraulic actuator plant", + "texts": [ + " In section 3, the stability of the reduced order MCS control in the case of electrohydraulic actuator plant is proven by Popov\u2019s hyperstability theory. In section 4, the proportional-plus derivative feedback (P + DFB) control and the MCS control are synthesized, and the results of the comparative implementation tests are detailed in section 5. Finally, the main conclusions to the work are listed in section 6. The plant consists of a Moog 760 torque motor/ flapper operated four-way double-acting servovalve, a hydraulic pump, two accumulators, a single rod actuator, and a hydraulic arm as shown in Fig. 1. The actuator cylinder has a diameter of 32 mm, and Proc. IMechE Vol. 221 Part G: J. Aerospace Engineering JAERO169 \u00a9 IMechE 2007 at University of Birmingham on June 3, 2015pig.sagepub.comDownloaded from it has a stroke of approximately 100 mm, and the end of the arm is connected to an inertial load representing an aircraft control surface. The hydraulic pump can supply a maximum pressure of 17.2 MPa. During the tests the maximum pressure supplied by the hydraulic pump is limited to 11.0 MPa. The constant pressure hydraulic power supply is an integral part of the actuator system" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003021_jae-2010-1280-Figure4-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003021_jae-2010-1280-Figure4-1.png", + "caption": "Fig. 4. Coiled tube model.", + "texts": [ + " In order to avoid initial penetration of the contact segments, certain initial gap should be assigned between the tube skins or the adjacent layers of folded tube. Sketch of folded location is shown in Fig. 3b, and T1 is the gap distance, L is the distance of adjacent folded layers, \u03b1 is the folding angle. The skin length errors between inner skin and outer skin is \u2206L = 2 ( L\u2032 \u2212 L1 ) = 2T1/cos \u03b1 (8) Introduce an artificial wrinkles in folded locations can modify the geometric modelling error. Figure 3c and 3d show the detail of artificial wrinkles in the folded location. The coiled tube in Fig. 4 is rolled in spiral with a rolling diameter of 0.1 m in the form of an Archimedean linear spiral shape governed by the relationship, R = R0 + a\u03b8, in which, a = (T1 + T2)/2\u03c0, and T1 is the spacing of tube skins, T2 is the spacing of adjacent layers of folded tube), R0 is the initial radius, and \u03b8 is the sweep angle. The coiled model is built with five connected control volumes links to each other through dummy partitioned membranes. The base cap is the same as the Z-folded tube model. For a long coiled tube, installing Velcro strips is a good choice to improve the deployment stability" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001982_978-3-540-36841-0_154-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001982_978-3-540-36841-0_154-Figure1-1.png", + "caption": "Fig. 1. The x-ray tube generates x rays that are restricted by the aperture in the collimator. The Al filter removes low-energy x rays that would not penetrate the body. Scattered secondary radiation is trapped by the grid, whereas primary radiation strikes the screen phosphor. The resulting light exposes the film.", + "texts": [ + "FMBE Proceedings Vol. 14/2 Keywords\u2014 medical devices, computed tomography, artificial kidney, cardiac pacemaker, ventilator I. THE 10 MOST IMPORTANT BIOMEDICAL ENGINEERING DEVICES (1) The X-ray machine images internal organs and thus discovers internal abnormalities and tumors in time to remove them. A simple x-ray system consists of a high voltage generator, an x-ray tube, a collimator, the object or patient, an intensifying screen, and the film (see Fig. 1). A simple x-ray generator has a line circuit breaker, a variable autotransformer, an exposure timer and contactor, a step-up transformer and rectifier, and a filament control for the tube. Medical exposures are of the order of 80 kVp (peak kilovolts), 300 mA, 0.1 s. Power levels range up to more than 100 kW [1]. (2) Computed tomography generates slice images of internal organs with improved contrast and spatial resolution. A way to minimize obstruction of one structure by another is to expose radiographs from several directions" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003808_jrc2009-63011-Figure7-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003808_jrc2009-63011-Figure7-1.png", + "caption": "Figure 7. Service-induced residual hoop stress for braking duration of 40 min (36\u201d freight car wheel - rim thickness = 1.5 in).", + "texts": [ + " This compressive yielding occurs when the material has reduced yield strength due to the increased temperature. The temperature developed on the tread surface during the on-tread braking for 20 min is not high enough to reduce the yield strength of the material at the tread surface; therefore, the material has not yielded in compression nor developed any plastic deformation, thereby developing no additional residual stresses in the wheel rim. Copyright \u00a9 2009 by ASME ms of Use: http://www.asme.org/about-asme/terms-of-use Down Figure 7 shows the residual stress distribution in a 36\u201d freight car wheel with rim thickness 38.1 mm (1.5 in), developed during both the manufacturing process and thermal brake loading for 40 min. The results show that the residual hoop stress on the taping line is -100 MPa (-14.5 ksi). Figure 8 shows the residual stress distribution in a 36\u201d freight car wheel with rim thickness of 38.1 mm (1.5 in), developed during both the manufacturing process and thermal brake loading for 60 min. The results show that the thermal brake loading for 60 min reverses the residual hoop stresses on the taping line from compression to tension in the wheels of considered rim thicknesses" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000889_wcica.2004.1343779-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000889_wcica.2004.1343779-Figure1-1.png", + "caption": "Fig. 1 Basic Angle Relationships", + "texts": [ + " For flight control, one interesting approach towards nonlinear flight control proposed by Taeyoung Lee [2] uses a backstepping and neural networks controller. Another one proposed by Steinberg [ 5 ] uses adaptive backstepping approach on a complex flight dynamic model. The main contribution of this paper is to apply backstepping and variable control technique to terrain following controller design. 11. ADFTIVE ANGLE METHOD Adaptive angle methods have relatively broad applications in TF systems [7]. The basic angle relationships and terrain obstacle to aircraft are shown in Fig 1. If it is required that aircraft fly over obstacles at predesignated altitude hp, , the flight path inclination angle control command is where M is the current position of aircraft, H is flight altitude, D is the range from M to the highest terrain point and h, is the highest terrain point altitude. From (l), it is not yet possible to make adequate use of the maneuver capabilities. Adaptive angle method control command production on the foundation of basic angle command relationship (1) -adequately considers aircraft maneuver capabilities. Introducing a non-negative function-suppression function Fs-as well as a constant value gain-angle gain K , , we obtain the flight path inclination angle control command BFL = KFL (arcfg((h, + h,, - H ) / D ) - Fs) (2) According to the idea of adaptive angle method and Fig.1, we obtain (3) as follows: (3) where 9 , R stand for pitch angle and slant distance respectively. In (2), the effect of suppression function Fs is nothing else than reducing the control commands produced with regard to terrain far from the aircraft. It is a non-negative function. Due to control commands, it is necessary to guarantee that aircraft, using maneuver capabilities, can fly over obstacles. On basis of research experience and simulations, suppression functions Fs and slant distance can be approximately described as a three-section linear function relationship (see [7])" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002979_978-3-540-70534-5_7-Figure21.4-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002979_978-3-540-70534-5_7-Figure21.4-1.png", + "caption": "Figure 21.4: Neural network for driving a mobile robot", + "texts": [ + " The standard method is supervised learning, for example through error backpropagation (see Section 21.3). The same task is repeatedly run by the NN and the outcome judged by a supervisor. Errors made by the network are backpropagated from the output layer via the hidden layer to the input layer, amending the weights of each connection. Feed-Forward Networks 335 Evolutionary algorithms provide another method for determining the weights of a neural network. For example, a genetic algorithm (see Chapter 22) can be used to evolve an optimal set of neuron weights. Figure 21.4 shows the experimental setup for an NN that should drive a mobile robot collision-free through a maze (for example left-wall following) with constant speed. Since we are using three sensor inputs and two motor outputs and we chose six hidden neurons, our network has 3 + 6 + 2 neurons in total. The input layer receives the sensor data from the infrared PSD distance sensors and the output layer produces driving commands for the left and right motors of a robot with differential drive steering. Let us calculate the output of an NN for a simpler case with 2 + 4 + 1 neurons", + " For most large tasks, Neural Networks 344 21 the ideal mapping from input to action is not clearly specified nor readily apparent. Such tasks require a control program that must be carefully designed and tested in the robot\u2019s operational environment. The creation of these control programs is an ongoing concern in robotics as the range of viable application domains expands, increasing the complexity of tasks expected of autonomous robots. A number of questions need to be answered before the feed-forward ANN in Figure 21.4 can be implemented. Among them are: How can the success of the network be measured? The robot should perform a collision-free left-wall following. How can the training be performed? What is the desired motor output for each situation? The motor function that drives the robot close to the wall on the left-hand side and avoids collisions. Neural networks have been successfully used to mediate directly between sensors and actuators to perform certain tasks. Past research has focused on using neural net controllers to learn individual behaviors" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001218_detc2006-99153-Figure6-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001218_detc2006-99153-Figure6-1.png", + "caption": "Figure 6. Dual angle representation.", + "texts": [ + " The angle of rotation of the rigid body around the pole is the angle between the line connecting any point of the body at the first position to the pole and the line that connects the corresponding point after rotation to the pole (see fig 1). Following the same construction for the three dimensional generalization of the Reuleaux\u2019s method, we can easily see that the angle of rotation of the rigid body around the screw axis will be the angle between the common perpendicular of any line of the rigid body before displacement and the screw and the common perpendicular of the corresponding line after displacement and the screw axis (see fig. 6). 4 aded From: http://proceedings.asmedigitalcollection.asme.org/pdfaccess.ashx?url Let g\u0302 be the common perpendicular between any line of the body at the first position and the screw axis and g\u0302\u2032 be the common perpendicular between the corresponding line after the helical motion and the screw axis. Both lines will intersect the screw axis at a perpendicular angle. The angle of rotation of the rigid body around the screw axis will be the angle between g\u0302 and g\u0302\u2032, and the distance d that the body translated along the screw will be the distance between the same two lines" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003840_detc2010-28961-Figure6-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003840_detc2010-28961-Figure6-1.png", + "caption": "Fig. 6 An eight-bar linkage with prismatic joint", + "texts": [ + " The approach is simple and straightforward. It is based only on the loop equations of the eight-bar linkages, which reveals the mathematical fundamentals for the formation of branches, sub-branches and other mobility issues of the entire linkage. The proposed method has the following merits, features, and suggestions. (1) This method is algebraic and its applicability is not restricted by the type of joints or the physical appearance of the linkage [9]. For example, the method is valid for the linkage type in Fig. 6, which contains prismatic joint. The method may be used even in spherical eight-bar linkages. (2) The use of the discriminant method can be explained and understood easily with the concept of joint rotation space. (3) Any multiloop linkage can be regarded as one consisting of certain basic chains or modules, which are the building block for more complex linkages. In this paper, these basic modules include four-bar, five-bar, six-bar, and sevenbar chains. It suggests that if the mobility of the basic modules becomes known, the mobility of complex linkages consisting of these basic modules can be rectified by considering the interaction between the basic modules [24]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003854_ssp.147-149.776-Figure4-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003854_ssp.147-149.776-Figure4-1.png", + "caption": "Fig. 4. Defining superficies of contact for a single layer", + "texts": [ + " The diagram representation of this method is presented in Fig. 2. Sections obtained with this method were then registered with Canon EOS 5D in 2496x1664 pixels resolution. A sample picture has been presented in Fig. 3a. Each image has been put through edge detection filter (Fig. 3b). Next, a circle of implant-sized perimeter was inscribed into the resulting outline (Fig. 3c). The above operations allowed for precise identification of the contact area between the implant and the surrounding tissue (Fig. 4). The difference in areas between the outline and the inscribed circle constitutes the amount of clearance gap between the implant and the tissue. To define the superficies of the clearance area, the computer image analysis software MicroMeter was used. The macroscopic testing showed that the weakest osteointegration was achieved in the case of solid alloys. All implants were loosely embedded and easily removed from the bone after post-processing. Simultaneously conducted microscopic testing of the implant surface also confirmed a lack of clearly defined places of bone adherence to the implant surface" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002285_1.2839012-Figure3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002285_1.2839012-Figure3-1.png", + "caption": "Fig. 3 Auxiliary reference frames for gear i \u201ei=1,2\u2026", + "texts": [ + "org/ on 01/28/201 M0 = \u2212 1 0 0 a0 0 \u2212 u0 \u2212 0 0 0 \u2212 0 u0 0 0 0 0 1 5 where u0 = cos 0, 0 = sin 0 6 The point of contact between a right-hand flank of gear 1 and a right-hand flank of gear 2 is bound to lie on a straight line that is tangent to the base cylinders of the two gears at points P1,1 and P2,1 in this two-index notation, the first index refers to the gear and the second index to the tooth flank: 1 for a right-hand flank and \u22121 for a left-hand flank . The line segment P1,1P2,1 is the path of contact for right-hand flanks. To determine points P1,1 and P2,1, together with their mutual distance 1, two auxiliary reference frames, V1,1 and V2,1, are introduced. The origin Bi,1 of Vi,1 i=1,2 is on the axis of gear i, at the transverse section for gear i that contains point Pi,1. The z axis of Vi,1 has the same orientation and direction as the z axis of Wi, whereas the x axis of Vi,1 is oriented from Bi,1 to Pi,1 see Fig. 3 . If i,1 is the angle of the rotation about the z axis of Wi that would make the axes of Wi parallel to the axes of Vi,1, the 4 4 matrix Mi,1 for the transformation of coordinates from Vi,1 to Wi is given by Mi,1 = ci,1 \u2212 si,1 0 0 si,1 ci,1 0 0 0 0 1 bi,1 0 0 0 1 7 where ci,1 = cos i,1, si,1 = sin i,1 8 and bi,1 is the z coordinate of point Bi,1 in reference frame Wi. The homogeneous components in Vi,1 of the unit vector ei,1 of the contact path P1,1P2,1, directed from Pi,1 to the other extremity of the contact path, is provided by ei,1 Vi,1 = 0 \u2212 ui i 0 T 9 where ui is defined by Eq" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003644_ijmmm.2009.026900-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003644_ijmmm.2009.026900-Figure1-1.png", + "caption": "Figure 1 (a) set-up showing the location of spray nozzle relative to the exit-interface surface and workpiece (b) details of the MQL equipment (see online version for colours)", + "texts": [ + " The initial lubricant flow was 15 ml/h and the air flow was 9 m3/h. A flow of 30 ml/h was tested that performed similarly for the same air flow, and thus, the 15 ml/h flow was chosen to carry out the definitive experiments because it is a low oil consumption condition. The MQL equipment permits fine separate lubricant/air regulation, by a needle-type gauge, atomising at an air flow of 6.0 kgf/cm2 constant pressure. The spray nozzle was placed about 25 mm from the tool turned towards the exit-interface surface as shown in Figure 1(a), where the lubricant dosage and air flow rate adjustment are carried out. Figure 1(b) shows the details of the MQL equipment and its parts are enumerated in order to be of easy understanding about description and functioning. The following results refer to the summary of the best cutting conditions (vc = 500 m/min, f = 0.10 mm/rev and ap = 0.35 mm) observed in Inconel 718 machining. No publications were found related to nickel-based alloy machining using the MQL technology, especially at high cutting speed, indicating that it has not yet been sufficiently researched. Figures 2 and 3 show the influence of geometry on the mean Ra roughness (\u03bcm) values with ceramic tools (CC650 and CC670) and PCBN tools (CB7050) in the dry cutting and using the MQL technique" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000597_cbo9780511547126.019-Figure17.3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000597_cbo9780511547126.019-Figure17.3-1.png", + "caption": "Figure 17.3.1: Normal sections of pinion and gear rack-cutters: (a) mismatched profiles; (b) profiles of pinion rack-cutter in coordinate systems Sa and Sb; (c) profiles of gear rack-cutter in coordinate systems Se and Sk.", + "texts": [ + " Henceforth, we consider the normal and transverse sections of the rack-cutter tooth surface. The normal section a\u2013a of the rack-cutter is obtained by a plane that is perpendicular to plane and whose orientation is determined by angle \u03b2 [Fig. 17.2.1(b)]. The transverse section of the rack-cutter is determined as a section by a plane that has the orientation of b\u2013b [Fig. 17.2.1(b)]. It was mentioned above that two mismatched rack-cutters are applied for separate generation of the pinion and the gear of the new version of helical gears. Figure 17.3.1(a) shows the profiles of the normal sections of the rack-cutters. Figures 17.3.1(b) and 17.3.1(c) show the profiles of the pinion and gear rack-cutters, respectively. Dimensions s1 and s2 are related by module m and parameter s12 as follows: s1 + s2 = \u03c0m (17.3.1) s12 = s1 s2 . (17.3.2) Here, s12 is chosen in the process of optimization, relates pinion and gear tooth thicknesses, and can be varied in the design to modify the relative rigidity. In a conventional case of design, we have s12 = 1. Cambridge Books Online \u00a9 Cambridge University Press, 2009available at https://www.cambridge.org/core/t rms. https://doi.org/10.1017/CBO9780511547126.019 Downloaded from https://www.cambridge.org/core. University of Warwick, on 18 Sep 2018 at 01:19:11, subject to the Cambridge Core terms of use, P1: JXR CB672-17 CB672/Litvin CB672/Litvin-v2.cls February 27, 2004 0:58 The profiles of the rack-cutters are parabolic curves that are in internal tangency. Points Q and Q\u2217 [Fig. 17.3.1(a)] are the points of tangency of the normal profiles of the driving and coast sides of the teeth, respectively. The common normal to the profiles passes through point P that belongs to the instantaneous axis of rotation P1\u2013P2 Figure 17.3.2: Parabolic profiles of rack-cutter in normal section. Cambridge Books Online \u00a9 Cambridge University Press, 2009available at https://www.cambridge.org/core/t rms. https://doi.org/10.1017/CBO9780511547126.019 Downloaded from https://www.cambridge.org/core. University of Warwick, on 18 Sep 2018 at 01:19:11, subject to the Cambridge Core terms of use, P1: JXR CB672-17 CB672/Litvin CB672/Litvin-v2.cls February 27, 2004 0:58 [Fig. 17.2.1(a)]. A parabolic profile of a rack-cutter is represented in parametric form in an auxiliary coordinate system Si (xi , yi ) as follows (Fig. 17.3.2): xi = ui , yi = ai u 2 i (17.3.3) where ai is the parabola coefficient. The origin of Si coincides with Q. The surface of the rack-cutter is designated by c and is derived as follows: (i) The mismatched profiles of pinion and gear rack-cutters are represented in Figure 17.3.1(a). The pressure angles are \u03b1d for the driving profile and \u03b1c for the coast profile. The locations of points Q and Q\u2217 are designated by |QP | = ld and |Q\u2217 P | = lc where ld and lc are defined as ld = \u03c0m 1 + s12 \u00b7 sin \u03b1d cos \u03b1d cos \u03b1c sin(\u03b1d + \u03b1c ) (17.3.4) lc = \u03c0m 1 + s12 \u00b7 sin \u03b1c cos \u03b1c cos \u03b1d sin(\u03b1d + \u03b1c ) . (17.3.5) (ii) Coordinate systems Sa and Sb are located in the plane of the normal section of the rack-cutter [Fig. 17.3.1(b)]. The normal profile is represented in Sb by the matrix equation rb(uc ) = Mba ra (uc ) = Mba [uc acu 2 c 0 1]T. (17.3.6) (iii) The rack-cutter surface c is represented in coordinate system Sc (Fig. 17.3.3) wherein the normal profile performs translational motion along c\u2013c. Then we obtain that surface c is determined by vector function rc (uc , \u03b8c ) = Mcb(\u03b8c )rb(uc ) = Mcb(\u03b8c )Mba ra (uc ). (17.3.7) \u03b8 Figure 17.3.3: For derivation of pinion rack-cutter. Cambridge Books Online \u00a9 Cambridge University Press, 2009available at https://www.cambridge.org/core/t rms. https://doi.org/10.1017/CBO9780511547126.019 Downloaded from https://www.cambridge.org/core. University of Warwick, on 18 Sep 2018 at 01:19:11, subject to the Cambridge Core terms of use, P1: JXR CB672-17 CB672/Litvin CB672/Litvin-v2.cls February 27, 2004 0:58 We apply coordinate systems Se and Sk [Fig. 17.3.1(c)] and coordinate system St which is similar to system Sc (Fig. 17.3.3). The coordinate transformation from Sk to St is similar to the transformation from Sb to Sc . The gear rack-cutter surface is represented by the following matrix equation: rt (ut , \u03b8t ) = Mtk(\u03b8t )Mkere (ut ). (17.3.8) The idea of mismatched rack-cutters may be extended to the design of modified involute helical gears as follows: (i) The rack-cutter for the pinion is applied as a parabolic one, but the rack-cutter for the gear is a conventional one and has straight-line profiles in the normal section" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000508_a:1008709005382-Figure6-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000508_a:1008709005382-Figure6-1.png", + "caption": "Figure 6. Absorption spectra of eriochrome doped film in copper solution (a) with (-\u2217-\u2217-\u2217-) and (b) without ( ) possible interfering metal ions (Fe+2, Mg+2, Al+3, Ni+2, Cr+2, Zn+2).", + "texts": [ + " There was no restoration of the colour complex, which indicated the complete removal of Cu++ from the matrix. The doped films were again equilibrated in copper solutions and detection was found to be reproducible (Fig. 5). The interferences of other metal ions, namely Fe+2, Mg+2, Al+3, Ni+2, Cr+2, Zn+2 were observed with the detection of Cu++ in solution. The concentration of each metal ion was set at five times the concentration of the Cu++ in any mixture. None of these ions appeared to affect the maximum absorption peak of copper at 565 nm (Fig. 6). This suggests that eriochrome cyanine in the sol-gel thin film is selective for Cu++. However this dye in solution is additionally selective for Al+++. This change in selectivity of eriochrome in the sol-gel matrix particularly in monoliths was studied in depth by Sommerdijk et al. [25]. They attributed this behaviour to the thermodynamic effect of complexation within the sol-gel matrix, which results in a modified translational entropy of solvent molecules arising through interaction with pore walls" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001198_esda2006-95466-Figure7-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001198_esda2006-95466-Figure7-1.png", + "caption": "Figure 7. RADIAL CLEARANCE", + "texts": [ + " So, the volumetric flow for unit length (qp), due to the pressure drop is given by: qp = h3 12\u00b5 d p dx (3) The volumetric flow due to the entrained flow has a linear distribution from zero to u, where u is the relative velocity of the tooth tip with respect to the case and defined by the expression: Qu = buh 2 (4) 5 nloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 02/07/2016 Term The height of the clearance between the tooth tip and the case, depends on the position of the gear shaft and on the case wear. So, the radial backlash will be different for each tooth along the gear case. This variation is depicted in figure 7 and is expressed by the relation: hre,i = hrc \u2212AB (5) For low value of the eccentricity e [6]: AB \u223c= AC = ecos(\u03d5i \u2212\u03d5e) (6) Therefore equation (5) becomes: hre,i = hrc \u2212 ecos(\u03d5i \u2212\u03d5e) (7) The height of the clearance between the tooth tip and the case depends also on the wear profile of the case. Therefore the clearance height hi of the tooth i, is given by: hi = hs(\u03d5i)+hre,i (8) where hs(\u03d5i) is the radial height of the wear profile for the tooth tip i. Copyright c\u00a9 2006 by ASME s of Use: http://www" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002819_6.2007-3813-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002819_6.2007-3813-Figure2-1.png", + "caption": "Figure 2. Blower Configuration.", + "texts": [ + " A detailed set of flow measurements comparing various mechanical and jet actuators for subsonic projectiles was not found in the literature, and it is hoped that the present study can provide a useful reference for others in this field. IV. Experimental Apparatus For the experiments described in this paper, a compressible flow wind tunnel at GTRI was modified to include a special test section. This test section reduced the wind tunnel cross section to a 16.51 cm by 12.7 cm rectangular section with a working length of 39.37 cm. With the two centrifugal blowers shown in Figure 2 operating in series in suction mode, it was possible to achieve Mt = 0.75. The test section shown in Figure 3 was designed to accommodate a wall mounted model of a projectile which was equipped with 60 pressure taps. It was necessary to wall mount the model in order to accommodate this large number of pressure taps. As the model itself was wall mounted, it was necessary to treat the boundary layer in order to improve the fidelity of the results. Given the two options of sucking off the boundary layer or energizing the boundary layer, the latter was chosen as the pressure inside to tunnel section was already well below atmospheric and it was difficult to provide a large suction pressure at that mass flow rate" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001164_1-84628-559-3_20-Figure5-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001164_1-84628-559-3_20-Figure5-1.png", + "caption": "Fig. 5. The 3-DOF rotational precision positioning stage", + "texts": [ + " Since one actuating unit is capable of acutaing the stage in only one direction of motion, totally, six actuating units marked from (a) to (f) are required to performe the 3-DOF of positioning. For examples, the actuating units set on the top side of xy plane and marked with (a) and (b) are used for actuating the stage to rotate with respect to z-axis; simlarily, the actuating units set on the xz plane and marked with (f) and (e) are used for actuating the stage to rotate with respect to y-axis as shown in Fig. 4 which is the sectional view A-A in Fig. 3. Figure 5 shows the perspecitve schematic drawing of the semi-shperical 3-DOF positionig stage. The friction adjusting mechanism set on the top surface of the stage is used for providing a suitable holding fricitonal force for the rotaional stage. Figure 6 shows a photograph of the modualrized springmounted PZT actuator, in which the PZT actuator has the dimension of 5 5 10 mm (Tokin). The stiffness of the spring is 0.023 N/mm. Six actuating units are symmetrically mounted to the positioning stage having a radius on the bottom side" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002567_6.2007-2277-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002567_6.2007-2277-Figure1-1.png", + "caption": "Fig. 1 A rotating tapered beam with (a) a constant width and a linearly varying depth for which the variations of the cross-sectional area and the second moment of area along the length are respectively linear and cubic (n = 1) and (b) a linearly varying width and depth for which the variations of the cross-sectional area and the second moment of area along the length are respectively second and fourth order (n=2).", + "texts": [ + " The dynamic stiffness matrix is then applied with particular reference to the Wittrick-William algorithm17, yielding natural frequencies and mode shapes of some examples. The numerical results are discussed and compared with published ones where possible. When presenting results, particular emphasis is placed on the inaccuracies of results that may occur as a result of using the Bernoulli-Euler theory as opposed to the Timoshenko theory. II. Theory Two types of rotating tapered beams are considered in this paper. They are shown in Fig. 1 in a right-handed Cartesian co-ordinate system with the Y-axis coinciding with the axis of the beam. The Z-axis is taken to be parallel, but not coincidental with the axis of rotation. It is assumed that the tapered beam is rotating at a constant angular velocity \u2126 with an arbitrary hub radius rH as shown. The tapered beam of rectangular cross section shown in Fig. 1(a) displays a linear variation of depth and a constant width of the cross-section along the length whereas the one in Fig. 1(b) shows a linear variation of both width and depth. Clearly for the former the variations of cross-sectional area and second moment of area are linear and cubic whereas those for the latter are of second and fourth order, respectively. If L is the length, c is the taper ratio, A(y) and I(y) are the area and second moment of area of the cross-section at a distance y , and \u03c1 and E are the density and Young\u2019s modulus of the (isotropic) beam material respectively, then the variations of the mass per unit length m(y), the flexural rigidity EI(y) and shear rigidity kGA(y) for both types of tapered beam can be expressed by using the following formulas. n L y cmyAym )1()()( 0 \u2212== \u03c1 (1) T American Institute of Aeronautics and Astronautics 3 2 0 )1()( +\u2212= n L y cEIyEI (2) n L y ckGAykGA )1()( 0 \u2212= (3) where 0m , 0EI and kGA0 are the mass per unit length, flexural rigidity and shear rigidity at the left-hand end of the beam, respectively. The integer n = 1 for the first type, see Fig. 1(a), and n = 2 for the second type, see Fig. 1(b), of tapered beams described by Eqs. (1)-(3). A large number of cross-sections can be constructed18-19 by using these two values of n, covering many practical cases. However, the rectangular cross-section is shown in Fig. 1 only for convenience. (It is evident that for such a cross-section, if one of the dimensions, say, the depth, is varied linearly and the other, say width, is kept constant, the value of n will be equal to 1 whereas if both the dimensions are varied linearly, n will take the value 2.) Note that since the density \u03c1, Young\u2019s modulus E, shear modulus G and section shape factor k are all constants, the variation of the cross-sectional area A(y) and second moment of area I(y) follow the same rules as Eqs", + " Any other form of coupling through Coriolis forces is also neglected. The governing differential equations of motion of the rotating tapered beam in free vibration are derived for the general case, see equations (1)-(3), by applying Hamilton\u2019s principle which requires the expressions for potential (or strain) and kinetic energies of the beam as fundamental prerequisites. The potential or strain energy U of the beam is given by20 [ ]\u222b \u2212\u2032+\u2032+\u2032= L dywkAGwFEIU 0 222 })()()({ 2 1 \u03b8\u03b8 (4) where w is the transverse displacement (in the Z-direction, see Fig. 1), a prime denotes differentiation with respect to y, and )(yF is the centrifugal force at a distance y from the root cross-section of the beam, arising from the rotating action. The centrifugal force F(y) can be expressed as8 0 2 )()()( FdyyrymyF H L y ++\u222b= \u2126 (5) where F0 is an outboard force at the right hand end of the beam, as shown in Fig.1 . (In order to make the theory sufficiently general, this outboard force F0, which is zero when the right hand end is completely free, is necessary4,8,15 e.g. when assembling dissimilar rotating tapered beams rigidly joined together.) The kinetic energy T of the beam is given by20 \u222b \u222b++= dywAdyIT L L 0 0 2222 )( 2 1 && \u03c1\u03b8\u03b8\u2126\u03c1 (6) where an over dot represents differentiation with respect to time t. Hamilton\u2019s principle states 0)( 2 1 =\u222b \u2212 dtUT t t \u03b4 (7) where t1 and t2 are the time intervals in the dynamic trajectory, and \u03b4 is the usual variational operator" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000215_icsmc.1995.538489-Figure9-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000215_icsmc.1995.538489-Figure9-1.png", + "caption": "Figure 9: Simulation results of the SICQP method for Example 1.", + "texts": [ + " 7, it is obvious that the PI method will produce discontinuous solutions of 81 when the End-effector is passing through the singular point. The discontinuity of 81 will cause abrupt change of the End-effector. As for (22). In accordance with W.1 and W.2 selecte h in the SRI the results of the SRI method, the solution curve of 04 revealrr a lump within the restricted region. (See Fig. 8 small lump can also be found in the solution curve o 6'2. Therefore, the End-effector will deviate when it is passing through the singular point. Observing Fig. 9, all of the undesirable phenomena shown in the results of the PI and SRI methods are disappeared if the SICQP method is adopted. The superiority of the SICQP method can ale0 be observed by showing Fig. 10. It indicates that position and orientation tracking errors of the SICQP method are nearly zero; while those of the SRI method show a certain amount. 3 \" 5 Summary and Conclusions In this paper, various singularities of the 5-DOF GRYPHON manipulator are analyzed. They are interior singularity and boundary singularity" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003405_8.6202-Figure14-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003405_8.6202-Figure14-1.png", + "caption": "Fig. 14 Tubular 7178-T6 test specimen for hydrostatic tests of reinforced pressure vessels", + "texts": [ + " This prestressed condition may be developed by one of two techniques: 1) pressurize the wound composite case (as shown in Fig. 13) so as accurately to strain the aluminum a predetermined amount; and 2) accurately pretension the circumferential reinforcement during winding. These prestressing techniques can be applied to aluminum-Fiberglas and aluminum-steel wire composites. Table 4 gives a summary of the hydrostatic testing of fiberglas reinforced 7178-T6 tubular specimens of the type shown in Fig. 14. Ref. 5 describes, in greater detail, some of the test results included in Table 4. The test shown in Table 4 demonstrate the significant in crease in burst strength to density ratio possible with this technique. In particular, the small scatter in strength for tests 5 through 9, in which the same glass-resin combination was used, should be noted. Design, stress analysis, and the actual winding operation are relatively simple with either steel or Fiberglas compared to that required in an all filament wound vessel, because only the cylindrical portion is wound and the winding angle is nearly 90\u00b0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003791_biorob.2010.5628033-Figure3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003791_biorob.2010.5628033-Figure3-1.png", + "caption": "Fig. 3 Estimation of gait phase by observing value of the DC current of the treadmill motor", + "texts": [ + " After acquiring enough strength of affected leg, the velocity difference is gradually set smaller to conduct natural-like gait training, that is, to train the gait movement on the flat ground. At every phase, the real-time biofeedback of stance phase balance is used. The developed robotic treadmill is composed of following three factors. A. Gait phase detection A patient\u2019s gait phase, which includes the time balance of the stance and swing legs, is one of the most useful indexes to evaluate gait. As shown in Fig. 3, a novel algorithm estimating the gait phase on a split belt treadmill by observing only the DC motor current was previously proposed [17]. The accuracy of the developed gait phase detection system is useful, because it is the same as that of the foot switch, which is one of the most common measurement tools in clinical rehabilitation [18]. The advantages of our system are that it does not require a significant amount of preparation and it does not impose a burden on either the patient or the therapist, because the patient does not have to wear a device" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000343_cdc.1991.261538-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000343_cdc.1991.261538-Figure1-1.png", + "caption": "Figure 1: LVLH frame", + "texts": [ + " Section 2 pmvides backgmund about the station, its dynamics, and its dishubance envir~lrment. Section 3 describes control system designs and performance tdeof\u20acs. and section 4 describes mbusmess analysis with respeu to major mass property variations, using the structured singular value analysis m e ~ o l o g y . Background NomiaalOricncationS: As the space station orbits the earth, the attitude controller attempts to keep the x axis pointed along the flight path. the y axis aligned with the normal to the orbit plane. and the z axis pointed towards the surface of the earth (see Figure 1). Ihe mtating reference frame centered at the space station and oriented in these directions is referred to as Verti_cal Local Horizontal (LVLH). This orientation facilitates the operatim of antemw. star senson, and other fixed expehentat packages. Howcvcr, the solar panels must be kept pointed towards the sun for maximal efficiency, so they m mounted on mtating joints to keep tMU attitude inertiauy 6x4. CH3076-7/91/0000-2206$01 .OO 0 1991 IEEE 2206 Dynamics: given by: The space station rigid body attitude dynamics and kinematics are J + i, X J \"= 3%'% X J 5 + TE + (1) & + @ - a x ~ = O i = x , y , z where the parameters and variables are: J Body momentaf-inextia matrix E = k5e" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003881_j.precisioneng.2010.10.003-Figure5-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003881_j.precisioneng.2010.10.003-Figure5-1.png", + "caption": "Fig. 5. Principle of torque detection.", + "texts": [ + " Measurement of cutting torque in small end milling 4.1. Principle of measurement Depending on the transmission torque, a countertorque acts on a fixed part. Therefore, the transmission torque can be calculated by measuring the magnitude of countertorque. In the case of the WTD speed-increasing spindle, the ring and sun rollers are used as an input and output, respectively, and the carrier is fixed. We devised a mechanism in which the beam is fixed between the carrier and the stopper, and a piezofilm is placed on the beam, as shown in Fig. 5. In a previous study [6], a strain gage was used to detect the strain; however, it was changed to a piezofilm to improve sensitivity. The magnitude of the strain changes according to the cutting torque applied on the sun roller during end milling. The piezofilm used here is 16 mm \u00d7 41 mm and 0.028 mm in thickness. The developed speed-increasing spindle with measurement equipment was set in a vertical MC as shown in Fig. 6. The output time response at the piezofilm during idling is shown in Fig. 7(a), where the input rotary speed of the tool spindle of the MC is 2000 min\u22121 (the rotary speed of the tool is 10,000 min\u22121)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001393_kem.291-292.483-Figure5-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001393_kem.291-292.483-Figure5-1.png", + "caption": "Fig. 5 Schematic process of internal gear machining by NC method", + "texts": [ + " The fly-blade is design based on the mathematical modeling of the tooth profile of internal gear in Equation (5). The movement between the work piece and the fly-blade is the same like the meshing movement between the internal gear and the planet worm-gear. The teeth profiles of the internal gear are manufactured automatically and continuously without using index mechanism. The manufactured internal gear is shown in Fig. 4. The schematic of NC machining the tooth profile of internal gear is shown in Fig. 5. The Internal toroidal tooth profile is generated by the compound movement of the rotating of the ball end milling cutter which accomplishes the machining of tooth flank and the spatial movement along the trajectory of the milling cutter. The ball end milling cutter is designed in identical shape with the meshing roller. After the finish of one tooth profile, the index movement of the work piece of internal gear is completed by the rotating of the work table at angular velocity of t\u03c9 which depends on the teeth numbers of the internal gear" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002313_1.2827448-Figure4-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002313_1.2827448-Figure4-1.png", + "caption": "Fig. 4 First self-sustained mode", + "texts": [ + " Relative requencies of self-sustained vibrations diminish with the increase f rotor revolution velocity. The first, lowest, self-sustained mode exists in the whole range f rotor revolutions. Within the frequency range of 250\u2013330 Hz, t manifests itself in the shape of a straight conical precession, ith the amplitude at the shaft ends 170\u2013180 m and the presnce of a node between bearings, Figs. 3 and 4. With increasing otation frequency, there takes place a displacement of the modal ode toward the compressor\u2019s bearing and an increase of the shaft xis curvature Fig. 4 . At the rotor rotation frequency 396 Hz, the second selfustained mode of the rotor gets excited jumpwise, and beatings, haracteristic for a double-frequency oscillatory process, are atched. Initially, the shape of this mode is close to cylindrical, igs. 3 and 5. At the end of the investigated rotation frequency ange, at the frequency value 650\u2013670 Hz, the second mode hape becomes deflected with the node on the turbine shaft end Fig. 5 . The computed values of amplitudes of self-sustained odes comply satisfactorily with those obtained by experiment" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001868_iecon.2006.347655-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001868_iecon.2006.347655-Figure2-1.png", + "caption": "Fig. 2. Rotary inverted pendulum.", + "texts": [ + " (step 3) Obtain W1,i and W2,i approximated by \u039b1\u03c9(j\u03c9j) and \u039b2\u03c9(j\u03c9j) for j = 1, \u00b7 \u00b7 \u00b7, n as minimum phase and stable transfer matrices. (step 4) Design C\u221e,i again for the augmented plant Gi =W2,iPiW1,i. (step 5) Calculate the resulting controller Ci = W1,iC\u221e,iW2,i and the corresponding robust stability margin \u00b2i. (step 6) Evaluate \u00b2i\u2212 \u00b2i\u22121. If \u00b2i\u2212 \u00b2i\u22121 is enough small, then let W1 = W1,i and W2 = W2,i and terminate the procedure. Otherwise, let i = i + 1 and return to step2. In this paper, we address a rotary inverted pendulum depicted in Fig.2 as a controlled plant. The plant param- eters estimated through some identification tests and the description of symbols contained in Fig.2 are listed in Tab.I. By Lagrange formulation, we get the motion equations of the inverted pendulum as follows: \u03c4=J1\u03b8\u03081+Z1 cos(\u03b81\u2212\u03b82)\u03b8\u03082+Z1 sin(\u03b81\u2212\u03b82)\u03b8\u030722 +\u00b5p(\u03b8\u03071\u2212\u03b8\u03072)+\u00b5a\u03b8\u03071\u2212Z2g sin \u03b81, (8) 0=J2\u03b8\u03082+Z1 cos(\u03b81\u2212\u03b82)\u03b8\u03081\u2212Z1 sin(\u03b81\u2212\u03b82)\u03b8\u030721 \u2212\u00b5p(\u03b8\u03071\u2212\u03b8\u03072)\u2212Z3g sin \u03b82, (9) where J1 = Ja +mas 2 a +mpl 2 a, J2 = Jp +mps 2 p Z1 = mplasp, Z2 = masa +mpla, Z3 = mpsp Eqs.(8) and (9) represent a nonlinear model of the rotary inverted pendulum. Moreover, let the scheduling parameter p \u2261 1 \u2212 cos \u03b81 and introduce some approximation, we get an LPV model of the rotary inverted pendulum as \u03c4 = J1\u03b8\u03081 + Z1(1\u2212 p)\u03b8\u03082 + (\u00b5a + \u00b5p)\u03b8\u03071 \u2212\u00b5p\u03b8\u03072 \u2212 Z2g(1\u2212 p 3 )\u03b81 (10) 0 = J2\u03b8\u03082 + Z1(1\u2212 p)\u03b8\u03081 \u2212 \u00b5p\u03b8\u03071 + \u00b5p\u03b8\u03072 \u2212 Z3g\u03b82 (11) We will design the scheduling controller based on this LPV model and evaluate the control performance in time domain on the nonlinear model " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003893_iceee.2010.5661528-Figure4.2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003893_iceee.2010.5661528-Figure4.2-1.png", + "caption": "Figure 4.2. Distribution of measuring point concerning vibration and noise", + "texts": [ + " EXPERIMENT VALIDATION The PULSE vibration and noise system, which is produced by B&K Company from Denmark, is used to measure the noise of 5kW AFSMPMSM. Power source of this experiment is sine generators. The vibration and noise of motor rotating in noload is measured in anechoic chamber. Fig.4.1 shows the equipment and the anechoic chamber. And the measuring results are distracted according to the international standard of measurement noise, and distribution of measuring point concerning vibration and noise is showed in Fig.4.2. Fig.4.3 show time region wave of noise at same moment. Fig.4.4 shows electric current time region wave. Finally, the average sound pressure level after experiment is pTL =58.12dB. The noise of 22kW and 220kW AFSMPMSM are also calculated by using this FEM, respectively. And the average sound pressure level of comparison between calculating results and testing results are shown in Tabl . V. CONCLUSION The three dimensional sound field of AFSMPMSM is built and the FEM is also used to calculate the noise of the AFSMPMSM and the calculation errors of 5kW, 22kW, and 220kW AFSMPMSM are 2" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003928_s10409-010-0364-1-Figure3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003928_s10409-010-0364-1-Figure3-1.png", + "caption": "Fig. 3 The mechanical forces acting on a filament when cell swims in viscous fluid", + "texts": [ + " 2b, sense of flagellar rotation is defined as CCW or CW from the viewpoint of an observer looking along the helical axis of the flagellum into its point of attachment to the cell. In Fig. 2c, the torque applied by the motor to the flagellum combines with viscous resistance to rotation to give a twisting moment, or torsion, which is LH if the motor is rotating CCW and RH if the motor is rotating CW. 2.1 Experiment of bacterial filament in vivo The experiment was to investigate the natural flagellar filament attached to an energized cell body [17]. During cell swimming, Macnab observed normal to curly form transition on both wide and mutant bacteria filament. Figure 3 shows the motion of a filament of radius \u03b7 and wavelength (pitch) \u03bb moving at a constant velocity through a viscous medium. Let VN and VT be the velocities normal and tangential to an element of the filament of length \u03b4s. The medium offers a resistance to reactions \u03b4N = VNCN\u03b4s, \u03b4T = VTCT\u03b4s, where CN and CT are the frictional constants in the appropriate directions; \u03b8 is the pitch angle of the helix. Now we define \u03b4F = (\u03b4N sin \u03b8 \u2212 \u03b4T cos \u03b8)\u03b4s, (1) \u03b4C = (\u03b4N cos \u03b8 + \u03b4T sin \u03b8)\u03b4s\u03b7. Each element of a filament moving in the x direction (Fig. 3) has two velocity components: u is in the axial direction and \u03b7\u03c9 is perpendicular to it. The helix is assumed to be rotating at an angular frequency of \u03c9, thus tan \u03b8 = 2\u03c0\u03b7 \u03bb , VT = u cos \u03b8 + \u03b7 sin \u03b8, (2) VN = \u2212u sin \u03b8 + \u03b7 cos \u03b8. The total propulsive force and the torque exerting on the helix axis are, respectively F = m\u03bb sec \u03b8(CNVN sin \u03b8 \u2212 CTVT cos \u03b8), (3) C = m\u03bb sec \u03b8(VNCN cos \u03b8 + VTCT sin \u03b8), where m is the wave number along the filament, \u03b4s =\u03b4x sec \u03b8 . Substituting sin \u03b8 = 2\u03c0\u03b7/ \u221a 4\u03c02\u03b72 + \u03bb2 and cos \u03b8 = \u03bb/ \u221a 4\u03c02\u03b72 + \u03bb2 into Eq" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002686_eej.20406-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002686_eej.20406-Figure1-1.png", + "caption": "Fig. 1. Definition of coordinates.", + "texts": [ + " When the authors performed experiments on their estimation method using a 6-pole, 400-W, 1750 r/min IPMSM, they confirmed that normal operation occurred without incident over a range of 80 r/min to 1800 r/min under a rated load of 0% to 100%. Moreover, good transient characteristics were obtained. Note that because this estimation method uses the induced voltage of the motor, stable estimation at low speed is problematic. The authors have previously reported in Ref. 9 on the estimation of salient position when a motor has been stopped. 2.1 Motor coordinate system Figure 1 shows the coordinate system for the IPMSM. The \u03b1\u2013\u03b2 coordinate axes represent the two-phase orthogonal coordinate axes for the stator. The d axis follows the direction of the magnetic pole of the rotor, and the angle from the \u03b1 axis (electrical angle) is designated \u03b8. Furthermore, \u03b3 follows the direction of the estimated magnetic pole for the controller. The angle from the \u03b1 axis is designated \u03b8M, and the error with the d axis is \u2206\u03b8 (= \u03b8 \u2013 \u03b8M). 2.2 The machine model using the extended induced voltage The machine model on the rotating coordinates is in general given by the equation Here, vd and vq represent the d and q axis components of the rotor voltage; id and iq, the d and q axis components of the rotor; \u03c9, the rotor angular velocity (electrical angle); R, the winding resistance; Ld and Lq, the d and q axis inductance; KE, the induced voltage constant; and p, the differential operator (d/dt)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001476_s11661-006-1006-x-Figure6-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001476_s11661-006-1006-x-Figure6-1.png", + "caption": "Fig. 6\u2014Schematic of the interpenetration of two sets of close-packed planes in the and phases. The relationship of g and the spacing of the Moire planes is indicated.", + "texts": [ + " (3) The invariant line will generally be stepped on an atomic scale; this is not inconsistent with Aaronson et al.\u2019s concept of structural ledges[10] as part of the equilibrium structure of the irrational habit-plane interface. (4) In essentially all cases of diffusional precipitation, growth ledges are required for the advance of habit-plane (and other) interfaces. As discussed in Section III\u2013A, these ledges must reproduce the equilibrium facet structure at its new, advanced, position. The g criterion is useful in predicting the heights of ledges that will meet this requirement (as indicated in Figure 6). (5) On occasion, growth ledges may be required to carry elastically accommodated (uncompensated) displacements, both parallel to and perpendicular to the habit plane. Elastic interactions among growth ledges are then to be expected. (6) Martensitic interfaces, as modeled by Pond et al.,[19] are thought to comprise sets of glissile disconnections, whose synchronous motion is responsible for the growth of the new phase. This is a strong constraint, which overrides the usual requirement that the transformation interface structure be one of minimum energy for a given OR", + " However, within this constraint, one expects that any facets formed on these interfaces will still be in local minimum-energy configurations. The geometry of one of the well-characterized Nie\u2013Muddle[29] interfaces was recently modeled by Reynolds et al.[31] The energetics of this interface is the subject of a separate investigation by Reynolds and Farkas, using embedded-atom methods.[32] This facet is one for which close-packed planes in the two phases meet edge to edge; it is also one for which an invariant line exists, as indicated schematically in Figure 6. The density of NCS is low, however (of the order of 1 near-coincident site per 20 m atoms), and this raises questions about the significance of coincidence-site density as opposed to plane-edge matching. Moire planes are also noted in the figure, indicating that the interface will encounter the edge-to-edge matching condition in a periodic fashion as it is displaced normal to itself. Since one of the requirements of growth ledges is that the facet-plane structure be displaced and 862\u2014VOLUME 37A, MARCH 2006 METALLURGICAL AND MATERIALS TRANSACTIONS A reproduced, one expects that the growth-ledge height will be the Moire plane spacing (1/| g |), or some multiple of it.*[22] *This is not entirely general, as may be seen by allowing the angle between the matching planes to approach zero in Figure 6. The Moire plane spacing then becomes large, and it is likely that the interfacial ledges will adopt a smaller height and tolerate an uncompensated elastic strain. It is clear that these massive-transformation interfaces must sample a range of possible facet orientations as they move into the parent phase and that, for any imposed OR, there is a strong possibility of an encounter with a (local) minimum-energy planar configuration. On the strength of recent experimental and modeling investigations, an edgeto-edge plane-matching condition appears to be a necessary and sufficient condition for the formation of a facet" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001793_gt2006-90280-Figure4-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001793_gt2006-90280-Figure4-1.png", + "caption": "Fig. 4 Mass, spring, and damper elements for tilting pad", + "texts": [ + " However, pivot stiffness can be of the same order of magnitude as the oil film stiffness, and thus can play an important role in the bearing \u00a9 National Research Council of Canada rl=/data/conferences/gt2006/71120/ on 02/13/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use D dynamic properties. Figure 3 shows pivot stiffness for a typical 100 mm diameter bearing calculated using the formulae given by Kirk and Reedy [16]. Thus, each pad can be represented by the mass, spring, and damper elements, as shown in Figure 4. The shaft experiences the combined action of all the elements of the system represented by equivalent stiffness and damping coefficients, which can be evaluated from the following formulae: ) = (10) (11) where (12) i,j denote directions \u03be or \u03b7 Vector summations of the and for all the pads give the bearing dynamic coefficients. )( , i jik )( , i jic In order to evaluate the frequency effects on the bearing dynamic properties calculations, an experimental investigation has been carried out for two operating conditions, which are typical for high speed rotating machinery" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003695_17543371jset49-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003695_17543371jset49-Figure1-1.png", + "caption": "Fig. 1 Forces acting on the athlete at the beginning of the drive phase", + "texts": [ + " An allowance of 7 per cent is added for wave-making resistance [25]. The drag is calculated in a quasi-steady (time-independent) way at each time step of the integration. The aerodynamic drag of each component is calculated on the basis of the apparent wind speed experienced by the component and using data from the work of Hoerner [26]. FFoot is the X component of the sum of the force FI due to the change in the momentum of the athlete and the reaction of the handle force through the feet, as can be seen in Fig. 1. The derivation of FI as a function of rowing technique is given in \u2018the inertial Proc. IMechE Vol. 224 Part P: J. Sports Engineering and Technology JSET49 by guest on March 9, 2015pip.sagepub.comDownloaded from forces and system mass\u2019. FX Gate is the component of the gate force acting in the fore\u2013aft direction, which is determined from the handle force by applying the no-slip assumption. The blade is assumed to rotate about a point which remains stationary in the axis of the boat\u2019s forward movement" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003847_cefc.2010.5481064-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003847_cefc.2010.5481064-Figure2-1.png", + "caption": "Fig. 2 6-pole magnetized PCMM", + "texts": [ + " In Fig. 1, the cross-sectional view of a pole-changing memory motor with 32 tangentially magnetized magnets is shown. On the rotor side there are four magnets per pole, all of them being magnetized in the same direction. PM along with iron segments build the rotor wreath which is mechanically fixed to a nonmagnetic shaft. After the stator winding is reconnected into six-pole configuration, a short pulse of stator current changes the rotor eight-pole magnetization into a six-pole one, as shown in Fig. 2. Since the number of magnets per pole is not any more an integer (32/6=.333...), same magnets can remain demagnetized. Issues such as magnetizing direction and quantity are important in evaluating the performance of the memory motor. Such characteristics depend upon the characteristic of material and, therefore, require a numerical evaluation. Whereas in other kinds of machines a rough estimation of hysteresis and magnetizing characteristics can be accepted, their importance in a memory motor justifies a greater effort in calculating them more precisely" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000282_tmag.2003.810347-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000282_tmag.2003.810347-Figure2-1.png", + "caption": "Fig. 2. Distribution of stress.", + "texts": [ + " In addition, the influence that it has on the torque characteristic is investigated. First of all, the portion of the stator that stress concentrates is clarified by the structural analysis. Next, the magnetic field of the claw-poled permanent magnet stepping motors is analyzed. The analysis model used in the structural analysis is shown in Fig. 1. The analysis model is for one pole of the stator, and a force as shown in Fig. 1 is applied. The stress distribution of a central section of a tooth is shown in Fig. 2. From Fig. 2, the pulled stress at magnet side and compressed stress at coil side is added. The part of the largest stress is a part of \u201cA\u201d at the coil side as shown in Fig. 2. Therefore, it becomes possible by taking into consideration that the magnetic characteristic at the Manuscript received June 18, 2002. Y. Okada is with Motor Technology Center, Motor Company, Matsushita Electric Industrial Co., Ltd, Osaka 574-0044, Japan (e-mail: yukihiro@mot.mei.co.jp). Y. Kawase and Y. Hisamatsu are with the Department of Information Science, Gifu University, Gifu 501-1193, Japan (e-mail: kawase@info.gifu-u.ac.jp). Digital Object Identifier 10.1109/TMAG.2003.810347 inside of the claw pole deteriorates to carry out magnetic field analysis with higher accuracy" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000345_iros.2000.895270-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000345_iros.2000.895270-Figure1-1.png", + "caption": "Figure 1: Capturing a spinning object.", + "texts": [ + " To capture a spinning object without the slip of the hands on the object, the motion of the hands must be synchronized with the spin of the object when the hands make contact. In addition, the spin has to be braked smoothly. The controller is established so that the two hands make contact with the object safely and grasp the object exerting the desired forces/moments. Finally, experimental results illustrate the effectiveness of our method. 2 Capturing Procedure In this paper, we propose the procedure of capturing a spinning cylindrical object using two flexible manipulators as shown in Figure 1. The procedure is established so that the manipulators follow the spin, grasp the object, and brake the spin automatically and smoothly. The procedure to capture the object is composed of the following five phases. phase 1: The hands are located near both sides of the object. tance to the object is always equal. phase 2: The hands move so that their relative dis- phase 3: As the hands approach and make contact with the object, they are synchronizing spin. phase 4: The internal forces/moments exerted to the object are controlled while the spin of the object are braked" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002979_978-3-540-70534-5_7-Figure16.22-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002979_978-3-540-70534-5_7-Figure16.22-1.png", + "caption": "Figure 16.22: Bug1 and Bug2 examples [Ng, Br\u00e4unl 2007]", + "texts": [ + "23 shows two more examples that further demonstrate the DistBug algorithm. In Figure 16.23, left, the goal is inside the E-shaped obstacle and cannot be reached. The robot first drives straight toward the goal, hits the obstacle, and records the hit point, then starts boundary following. After completion of a full circle around the obstacle, the robot returns to the hit point, which is its termination condition for an unreachable goal. To point out the differences between the two algorithms, we show the execution of the algorithms Bug1 (Figure 16.22, left) and Bug2 (Figure 16.22, right) in the same environment as Figure 16.21, right. Figure 16.23, right, shows a more complex example. After the hit point has been reached, the robot surrounds almost the whole obstacle until it finds the entry to the maze-like structure. It continues boundary following until the goal is directly reachable from the leave point. 263 16.10 Dijkstra\u2019s Algorithm Reference [Dijkstra 1959] Description Algorithm for computing all shortest paths from a given starting node in a fully connected graph" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001330_detc2005-84694-Figure6-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001330_detc2005-84694-Figure6-1.png", + "caption": "Figure 6 Front, cross-section, and bottom views of the deformed shape of the tire at time 0.601 sec. at N.m120=T .", + "texts": [ + " A force F is applied in the negative Y-direction in order to cause the tire to come into contact with the pavement. A pressure P is used to inflate the tire. The pavement is moved with a linear velocity V, while an opposing torque T is applied on the tire\u2019s wheel. F, P, T and V are ramped-up from a value of 0 at time 0 to their nominal value at time 0.3 sec. Then, they are maintained constant at the nominal value for the rest of the simulation. P is ramped up to a nominal value of 0.9 MPa. V is ramped up to a nominal value of 7.1 m/sec. F is ramped up to a nominal value of 1200 N. Figure 6 shows the deformed shape of the tire at time 0.6015 sec., when near-steady-state is reached, for a nominal T of 120 N.m. Figure 7 shows the time-history of the friction torque between the tire and the pavement for three values of the opposing torque T (0, 60, and 120 N.m). As expected the friction torque is almost equal to the opposing torque. The residual oscillations of the friction torque at steady-state are mostly due to the finite circumferential discretization of the tire. Figure 8 shows the time-history of the motion of the tire\u2019s rim center for the three values of T" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001293_rob.20099-Figure13-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001293_rob.20099-Figure13-1.png", + "caption": "Figure 13. Implementation of step 2 of corrective phase, minimization of rotational errors. a Rotation angles are determined. b Offsets are projected. c Vehicle is rotated about center of detector for locked LOS.", + "texts": [ + " Step 2: There exist two types of errors in the system; rotational and translational. The guidance algorithm addresses each type of error in separate substeps, respectively. The rotational errors are addressed first. The transformation, C4, applied to the vehicle e.g., aFc , is with respect to coordinate frame dFc, where M4 = TdTr 20 and C4 = KpM4 + Kd dM4 dt . 21 In order to keep the first LOS determined in step 1 locked to the center of its corresponding PSD, dFc is translated virtually not physically to the center of the specific PSD for which the LOS is locked, dFc * Figure 13 a . The transformation is then applied and the frame dFc is translated back virtually. In Eq. 22 , Tr represents the rotation applied to the vehicle, Tr = cos z cos y cos z sin y sin x \u2212 sin z cos x cos z sin y cos x + sin z sin x 0 sin z cos y sin z sin y sin x + cos z cos x sin z sin y cos x \u2212 cos z sin x 0 \u2212 sin y cos y sin x cos y cos x 0 0 0 0 1 , 22 where x, y, and z are the respective rotation angles with respect to dFc * Figure 13 a . The angular displacement values are determined as follows by utilizing all three PSD offsets: Step 2a: If e1 is emin, the distance from the y projection of e1, e1y to the center of aFc , f2, and the distance from the z-projection of e1, e1z to the center of aFc , f3, are calculated Figure 13 b . They are subsequently utilized, along with h1, the distance from the center of aFc to the center of Fd1, in determining the rotations about the desired y and z axes of dFc as y = cos\u22121 e1z 2 \u2212 f 3 2 \u2212 h1 2 \u2212 2h1f3 23 and z = cos\u22121 e1y 2 \u2212 f 2 2 \u2212 h1 2 \u2212 2h1f2 . 24 The minimum offset among the two remaining PSDs is used next in order to determine the last rotation angle about the desired x axis of dFc: x = cos\u22121 emin 2 \u2212 f 1 2 \u2212 h2 \u2212 2hf1 , 25 where f1 is the distance from the z or y projection e3 or e2 determined in Fd3 or Fd2, respectively, to the center of aFc depending on whether the minimum offset emin is e3 or e2 Figure 12 , e" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000396_2001-01-1445-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000396_2001-01-1445-Figure2-1.png", + "caption": "Fig. 2 Measurement points in RNCA", + "texts": [ + " In the end of this paper, the validity of this procedure is evaluated experimentally with new rubber mounts modified the stiffness of the initial mounts, and the utility of this procedure is verified. In this section, an in-operation test is performed to identify the problem frequency band of road noise, and the multiple inputs single output analysis technique is applied to the target vehicle to specify the relationship between road noise and accelerations on the suspension system. The in-operation test is carried out on a road noise evaluation course with a speed of 60km/h for the target vehicle. A schematic diagram of the measurement points is shown in Fig. 2. Accelerations are measured at a total of six points, i.e., four points (steering knuckle, shock absorber, lower arm and cross member) on the target front suspension system and two points (steering knuckle and shock absorber) on the rear suspension system, in order to consider the influence. The road noise is measured at the driver\u2019s ear position. The 1/3 octave band level of the road noise measured in this test is shown in Fig. 3. The conditions of the 160 and 315Hz bands are much worse than the other bands" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002456_1.25389-Figure14-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002456_1.25389-Figure14-1.png", + "caption": "Fig. 14 Re 1:12 105. Phase portrait onto \u2019\u2013 plane (channel 1\u2013 channel 2).", + "texts": [ + " The inner region of the attractor is thus more visited than what happened in Fig. 11. It should be noticed that the smaller orbits arising in the central region have almost the same shape as the larger ones, a typical characteristic of fractals. D ow nl oa de d by U N IV E R SI D A D D E S E V IL L A o n Fe br ua ry 2 3, 2 01 5 | h ttp :// ar c. ai aa .o rg | D O I: 1 0. 25 14 /1 .2 53 89 A further increase of the Reynolds number to Re 1:12 105 carries along amore complex structure of the attractor projected onto the plane \u2019\u2013 . In Fig. 14, two big trajectories seems to be associated with three small ones. This is due to the presence of four frequencies in the spectra, which are not reported but their values are plotted in Fig. 28. Because their ratio is still an apparently irrational number, the motion is believed to be quasi-periodic. More insights can be obtained from Poincar\u00e9 sections. Here we perform a cut in the plane 0; _ > 0 and plot\u2019 vs its time derivative _\u2019. Figure 15 shows the presence of a region frequently visited by the attractor and another more rarely visited" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000285_robot.1998.677403-Figure4-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000285_robot.1998.677403-Figure4-1.png", + "caption": "Fig. 4. The frames used an mathematical models", + "texts": [ + " It is impossible to obtain a general solution of its motion equations in a convenient analytical form. Therefore, its mathematical model is based on simplified and partially linearized equations. Solutions of the equations developed have a simple analytical form and accuracy acceptable for processing the measurement data and, motion prediction. To form the equations of mathematical model the right-hand Cartesian frame Oxyz is introduced. The point 0 is the middle of the straight line segment connecting the points A1 and A2 of the threads attachment to a fixed beam (Fig. 4). The axis O x is directed along the vector AzAl, the axis Oz is directed vertically downwards. Below, point coordinates and vector components are referred to the frame Oxyz. In this frame A1 = ( a , 0 , 01, Az = ( -a , 0 , 0 ) , and at rest the rod endpoints (i.e. B1, B2, see Fig. 4) have coordinates: B1 = (b , 0 , h ) , BZ = (4, 0 , h) . Here a = 0.34 m, b = 0.28 m and h = 2.2 m. ___t For the rod arbitrary position the coordinates of the points B1 and Bz are assumed to be B1 = ( b + \u20ac1, vi, h + Ci), Bz = ( - b + \u20ac2, qz, h + CZ). The quantities ti, qi, Ci (i = 1 , 2 ) are not independent variables. Three constraints are imposed on them which express invariability of length of the segments A1 B1, A2B2 and BI B2. The variables \u20ac 1 , ql and QZ can be assumed to be Lagrangian coordinates of the mechanical system under consideration" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001293_rob.20099-Figure9-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001293_rob.20099-Figure9-1.png", + "caption": "Figure 9. Utilization of LOS concept: a ideal vehicle relocation and b actual vehicle relocation, respectively.", + "texts": [ + " When an autonomous vehicle moves to a desired pose in 3D space, the actual transformation, Ta, differs from the desired transformation, Td, due to systematic and random errors. The first goal of the proposed guidance algorithm is, thus, to use the initial sensory offsets, immediately after the vehicle\u2019s longrange motion and prior to any short-range corrective actions, to estimate this actual transformation. Let us assume that for the generalized 6-dof docking problem considered herein, the vehicle\u2019s shape is a cube with array type detectors on at least three of its orthogonal faces, Figure 9. Three spatial LOSs are sufficient to specify any 3D relocation of the vehicle, Table I. Two actions are carried out concurrently: While the galvanometer mirrors are in the process of align- ing the three LOS, the vehicle moves to its desired pose, Td xd ,yd ,zd , d , d , d , with respect to the world coordinate frame, Fw, but only achieves an actual pose defined by Ta xa ,ya ,za , a , a , a Figure 10 a . It is assumed that the LOS can be aligned in a significantly shorter amount of time than that required to move the vehicle itself " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001857_ipemc.2006.4778106-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001857_ipemc.2006.4778106-Figure1-1.png", + "caption": "Figure 1.vector control phaser diagram", + "texts": [ + " The mismatch produces a coupling between the flux and torque producing channels in the machine and degrades the performance of the controller. In order to exact study of effect of machine parameters variation on drive outputs, the mathematical analyses are employed. For doing this, the machine equations which were derived in previous section are used. The motor electromagnetic torque from the equation (14) is equal to: Trtee iKT \u03bb= (16) Replacing r\u03bb and teK from equations (9) and (15) in equation (16), respectively we get: Tf r m r m it e ii pT L L L K T + = 1 1 (17) Where )2)( 3 2( p Kit = . Referring to the figure 1, following equations are achievable: f T T i i =\u03b8tan (18) TsT ii \u03b8sin= (19) Tsf ii \u03b8cos= (20) And replacing equations (18) to (20) in equation (17) produces: TTs r m r m it e i pT L L L K T \u03b8\u03b8 cossin 1 1 2 + = (21) Also the slip speed derived from equation (10) as: T r r f T r r r m T r m r T r m sl T pT i i T pT pT idsL i T Li T L \u03b8 \u03bb \u03c9 tan 11 1 + = + = + == (22) Rearranging the above equation for T\u03b8tan we can get: r rsl T pT T + = 1 tan \u03c9\u03b8 (23) From which sine and cosine of the torque angle are defined as: 2) 1 (1 )1()(sin r rsl rrsl T pT T pTT + + += \u03c9 \u03c9\u03b8 (24) 2) 1 (1 1cos r rsl T pT T + + = \u03c9 \u03b8 (25) Replacing T\u03b8sin and T\u03b8cos from (24) and (25), respectively in equation (21) we have: 2 22 2 ]) 1 (1][)1([ 1 ssl r rsl rr m it e i pT TpTR L K T \u03c9\u03c9 + ++ = (26) Where rrr RLT = " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001123_978-3-540-30585-9_52-Figure3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001123_978-3-540-30585-9_52-Figure3-1.png", + "caption": "Fig. 3. Nominal ( )2L j\u03c9 and bounds in automatic loop-shaping", + "texts": [ + " 2 shows the results of the loop-shaping for 1L to calculate the controller 1g with automatic loop-shaping using GA. The loop transmission 1L for the nominal case is calculated from the A, B, C, and D matrices and is satisfied with the margin specifications. The result of the manual loop-shaping shows that its 0[dB] frequency is approximately 10 [rad/s]. Next, the second procedure is that 2g is designed to satisfy inequality (8) where ( )im \u03c9 =3[dB]. The calculated bounds and the nominal second loop 2L are shown in Fig. 3. The designed controller 2g is ( ) ( )( ) 2 2 2 2 2 1 12.754 23.0830.873 0.8141 2 120.601 120.601 199.749 5161.035 12700.019 196.309 14544.600 s s g s s s s s s s + + = \u2212 + \u00d7 + + + = \u2212 + + (19) The optimization procedures for loop roll angle \u03a6 and loop yaw rate r based on GA allow obtaining controllers satisfying given specifications respectively. A gust input of 1[m/s] is shown in Fig 4-(a). And its results are shown in Fig. 4 from all 128 plant parameter variations. Each response is within the time response specifications" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003682_fie.2009.5350878-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003682_fie.2009.5350878-Figure2-1.png", + "caption": "FIGURE 2 HCI PARADIGM VS. SID PARADIGM.", + "texts": [ + " \u2022 Application Domains - including design, workspaces, education, e-commerce, entertainment, digital democracy, digital cities, policy and business. Figure 1 shows some SID characteristics and their granularities. By considering this framework as a focus for teaching engineering design, undergraduate students look for social 2 Prof. Toyoaki Nishida in Trento at 6th SID workshop, 2007. 978-1-4244-4714-5/09/$25.00 \u00a92009 IEEE October 18 - 21, 2009, San Antonio, TX 39th ASEE/IEEE Frontiers in Education Conference M1J-3 aspects and characteristics and keep their thinking at a systemic level [16]. Figure 2 shows the SID paradigm compared with the Human Computer Interaction (HCI) paradigm. In a HCI-based system the interaction is between a human and a machine/computer/system. With a SID-based system humans interact through it and by interacting they gain social and human capital, in particular social intelligence. In Figure 1 we see that for example by interacting through a SID-based System a group of humans can enhance different granularities of social intelligence such as their conviviality, reputation, awareness, etc" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003062_9780470447734.ch1-Figure1.1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003062_9780470447734.ch1-Figure1.1-1.png", + "caption": "FIGURE 1.1 (a) How the lattice vectors a1 and a2 of a graphene sheet can be used to describe a \u2018\u2018roll-up\u2019\u2019 vector to form a single-walled carbon nanotube; (b) space-filled model of a \u2018\u2018zigzag\u2019\u2019 (n,0)-SWCNT.", + "texts": [ + " However, work from 1976 by Oberlin and Endo [3] and in 1978 byWiles and Abrahamson [4] (subsequently republished in 1999 [5]) present arguably the earliest and clearest characterization of what would later be recognized as multiwalled CNTs. For a more detailed discussion of who should be credited with the discovery of CNTs, the interested reader is directed to the comprehensive editorial by Monthioux and Kuznetsov [6]. Structurally, CNTs can be approximated as rolled-up sheets of graphite. CNTs are formed in two principal types: single-walled carbon nanotubes (SWCNTs), which consist of a single tube of graphite, as shown in Figure 1.1, and multiwalled carbon nanotubes (MWCNTs), which consist of several concentric tubes of graphite fitted one inside the other. The diameters of CNTs can range from just a few nanometers in the case of SWCNTs to several tens of nanometers for MWCNTs. The lengths of the tubes are usually in the micrometer range. Conceptually, the way in which the graphite sheet is rolled up to form each nanotube affects the electronic properties of that CNT. In general, any lattice point in the graphite sheet can be described as a vector position (n,m) relative to any given origin. The graphite sheet can then be rolled into a tube such that the lattice point chosen is coincident with the origin (Figure 1.1). The orientation of this roll-up vector relative to the graphite sheet determines whether the tube forms a chiral, armchair, or zigzag SWCNT, terms that describe the manner in which the fused rings of a graphite sheet are arranged at the termini of an idealized tube.Alternatively, SWCNTs are more precisely described in the literature by the roll-up vector coordinates as [n,m]-SWCNTs. It has been shown that when |n m|\u00bc 3q, where q is an integer, the CNT is metallic or semimetallic and the remaining CNTs are semiconducting [7,8]", + " Therefore, statistically, one-third of SWCNTs are metallic depending on the method and conditions used during their production [7] and can possess high conductivity, greater than that ofmetallic copper, due to the ballistic (unscattered) nature of electron transport along a SWCNT [9]. If one considers the structure of a perfect crystal of graphite, two crystallographic faces can be identified, as shown in Figure 1.2. One crystal face consists of a plane containing all the carbon atoms of one graphite sheet, which we call the basal plane; the other crystal face is a plane perpendicular to the basal plane, which we call the edge plane. By analogy to the structure of graphite, two regions on a CNT can be identified (and are labeled in Figure 1.1) as (1) basal-plane-like regions comprising smooth, continuous tube walls and (2) edge-plane-like regions where the rolled-up graphite sheets terminate, typically located at the tube ends and around holes and defect sites along tube walls. As we discuss in Section 1.2, it is these edge-plane-like defects that are crucial to an understanding of some of the surface chemistry and the electrochemical behavior of CNT (MWCNT, in particular)\u2013based analytical and bioanalytical systems. In the case of MWCNTs, a number of morphological variations are also possible, depending on the conditions and chosen method of CNT formation [e" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003319_epepemc.2010.5606812-Figure3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003319_epepemc.2010.5606812-Figure3-1.png", + "caption": "Fig. 3. Simplified common mode equivalent circuit of a DFIG", + "texts": [ + " Only when the both voltage vectors are located in the same group, either even or odd, or the same zero vector is selected, the CMV is zero. The common mode voltage is applied to a complex circuit, e.g. in [2], which mainly consists of parasitic capacitances between the motor windings, laminations and the frame. Due to the isolation between the rotor windings and the rotor lamination in a DFIG, this equivalent circuit is more complex than the one of a squirrel-cage induction machine. As shown in Fig. 3 the CMV is mapped through the parasitic capacitances to the bearing, thereby the ratio is called bearing voltage ratio and can be expressed as BVR = bearing voltage common mode voltage = Ubrg Ucmv . (3) The bearing is modeled as a changeable capacitance. For normal operation the lubricant film is an isolator between the roller elements and the race ways. But if the voltage over the bearing is to high, the lubricant film breaks down and the bearing acts as a short circuit. This leads to a very short, sharp current impulse flowing through the bearing with a high energy density, because of the discharge of the nearby capacitances" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003450_mace.2010.5536045-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003450_mace.2010.5536045-Figure2-1.png", + "caption": "Figure 2. Forces and moments acting on a roller", + "texts": [ + " The roller waviness can be given by: [ ]1 1 cosj nj b bj n W C n t\u03c9 \u03be \u221e = = + (5) [ ]2 1 cos ( )j nj b bj n W C n t\u03c9 \u03c0 \u03be \u221e = = + + (6) Where, C is amplitude of race waviness, b\u03c9 is rotating speed of roller. The contact deformation between the jth roller and outer ring, inner ring can be drawn as: ' 1 1 1 1jk jk j jP W\u03b4 \u03b4= \u2212 \u2212 (7) ' 2 2 2 2jk jk j jP W\u03b4 \u03b4= \u2212 \u2212 (8) Where, 1 jk\u03b4 2 jk\u03b4 is the contact deformation between the jth roller and rings without considering waviness of bearing, which can be obtained by [11]. Fig. 2 gives the forces acting on a roller, where yjM , zjM are the inertia moments of the roller, djF is fluid drag force and 1 jT , 2 jT are traction forces. 1 jQ , 2 jQ are Hertzian contact forces. 1 jP , 2 jP are film stress of outer ring and inner ring. cjF , cjf are stress and friction force of cage. These forces can be obtained from the equations in [11]. An accurate traction model is used, in which non-Newton fluid is considered, the traction equation is given by[13]: ( ) CsA Bs e D\u03bc \u2212= + + (9) Where, s is slide to roll ratio of roller, the coefficients A,B,C and D are determined for a wide range from experiment data" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001979_50037-5-Figure7.49-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001979_50037-5-Figure7.49-1.png", + "caption": "Figure 7.49 Configuration of the contact between cylinder and two blocks.", + "texts": [ + " Pressing [B] OK implements the selection. From ANSYS Main Menu select General Postproc \u2192 Plot Results \u2192 Contour Plot \u2192 Nodal Solu. In appearing frame (see Figure 7.40), the following are selected as items to be contoured: [A] Stress and [B] von Mises (SEQV). Pressing [C] OK implements selections made. Figure 7.48 shows stress contours on the pin resulting from pulling out the arm. 7.2.2 Concave contact between cylinder and two blocks Configuration of the contact between cylinder and two blocks is shown in Figure 7.49. This is a typical contact problem, which in engineering applications is represented by a cylindrical rolling contact bearing. Also, the characteristic feature of the contact is that, nominally, surface contact takes place between elements. In reality, this is never the case due to surface roughness and unavoidable machining errors and dimensional tolerance. There is no geometrical interference when the cylinder and two blocks are assembled. This is a 3D analysis and advantage could be taken of the inherent symmetry of the model" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003735_s0025654410020020-Figure7-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003735_s0025654410020020-Figure7-1.png", + "caption": "Fig. 7.", + "texts": [ + "14) in a majority of problems of dynamics. The second-order model (2.15) is required for more precise qualitative analysis, for example, for determining the boundaries of the stagnation region or the motion full stop time. MECHANICS OF SOLIDS Vol. 45 No. 2 2010 As an example of systems with combined dry friction whose dynamics can be described by using the above models, consider the system consisting of a spring-fixed cylindrical rod vertically standing on an infinite band moving at a constant velocity v (Fig. 7). The dynamics of this system was studied in detail in [8] under the assumption that the rod end resting on the band has a hemispherical shape and hence the distribution of normal contact stresses obeys the Hertz law (2.5) (a \u201cpoint contact\u201d). Since the distinction of the Coulomb law in classical from that in generalized differential form is most significant for sufficiently extended contact site, it is assumed that the rod end resting on the band has a plane circular shape. In this case, the normal contact stress distribution is described by the Galin law (2", + "15) is required for a more precise qualitative analysis, for example, for determining the boundaries of stagnation regions or the motion full stop time. It is shown that the appearance of additional polynomial terms in the friction models (2.3), (2.14), and (2.15) permits explaining the possibility of the appearance of torsional self-excited vibrations in systems with combined dry friction consisting of an elastically fixed cylindrical rod one of whose ends rests on an infinite band moving at a constant velocity (Fig. 7). 1. P. Contensou, \u201cCouplage Entre Frottenment de Glissement et Frottenment de Pivotement Dans la The\u0301orie de la Toupie,\u201d in Kreiselprobleme Gyrodynamics (IUTAM Symp. Celerina, Berlin etc., Springer; Mir, Moscow, 1967), pp. 201\u2013216 (60\u201377). 2. V. Ph. Zhuravlev, \u201cThe Model of Dry Friction in the Problem of the Rolling of Rigid Bodies,\u201d Prikl. Mat. Mekh. 62 (5), 762\u2013767 (1998) [J. Appl. Math. Mech. (Engl. Transl.) 62 (5), 705\u2013710 (1998)]. 3. V. Ph. Zhuravlev and A. A. Kireenkov, \u201cPade\u0301 Expansions in the Two-Dimensional Model of Coulomb Friction,\u201d Izv" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003699_s1068798x10020085-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003699_s1068798x10020085-Figure2-1.png", + "caption": "Fig. 2. Slip line fields when \u03b3d \u2260 0 and \u03b1 \u2260 0.", + "texts": [ + " Cross sections close to surfaces that have already been machined and are about to be machined are characterized by a plane stress state. This leads to broadening of the chip (Figs. 1a and 1b). For blocked cutting (channel cutting in turning), the whole chip formation zone is characterized by a plane strain state (Fig. 1c). The slip line field in the plastic chip formation region for the case where the dynamic front (\u03b3d) and rear (\u03b1) cutter angles in each cross section are zero was considered in [2]. The slip field in the plastic region (\u03b3d \u2260 0; \u03b1 \u2260 0) was considered in [3] (Fig. 2). Formulas for \u03b1 and \u03b2 slip lines starting at the front surface were obtained: (1) where \u03bcfo is the frictional coefficient at the tip of the front surface [2]. The boundary \u03b1 line OAN (Fig. 2) passes through the cutter tip and is described by the equation (2) where for the \u03b1 lines zd 2lf\u03bcfo lf yd\u2013 \u03bcfolf \u03b3dtan\u2013 lf \u03b3dtan\u2013 yd \u03b3dtan \u03bcfolf\u2013+( ) \u03b3d 2tan 1+( ) yd \u03b3dtan( ) 2 yd+( )\u2013ln\u2013 1 \u03b3dtan+\u2013( ) 2 CI;+= for the \u03b2 lines zd 2lf\u03bcfo yd lf\u2013 \u03bcfolf \u03b3dtan yd \u03b3dtan lf \u03b3dtan\u2013 \u03bcfolf\u2013+ +( ) \u03b3d 2tan 1+( ) yd yd \u03b3dtan( ) 2 +\u2013ln\u2013 1 \u03b3dtan+( ) 2 CII,+= zd 2\u03bcfolfc\u03b1 1 \u03b3dtan( ) 2 +[ ]\u2013 1 \u03b3dtan( ) 2 \u2013[ ]yd+ 1 \u03b3dtan+( ) 2 ,= c\u03b1 lf\u2013 yd \u03bcfolf \u03b3dtan lf \u03b3tan yd \u03b3dtan\u2013 \u03bcfolf+ + + + lf 1\u2013 \u03bcfo \u03b3dtan \u03b3dtan \u03bcfo+ + +( ) \u239d \u23a0 \u239b \u239e ", + " 2 2010 PETRUSHIN, PROSKOKOV The boundary \u03b2 line EAA' is perpendicular to the zd line and leaves point E, with the coordinates yd = lpl cos \u03b3d, zd = \u2013lpl sin \u03b3d. It is described by the equation (3) where The coordinates of point A are obtained by solving Eqs. (2) and (3). For the boundary \u03b1 and \u03b2 lines at the rear cutter sur face, when \u03b1 \u2260 0, we may write the following formulas [3]: (4) (5) where \u03bcro is the frictional coefficient at the tip of the rear sur face [2]. The position of point B is determined by solution of Eqs. (4) and (5). The boundary CA'K (Fig. 2) is parallel to line OAN at a distance equal to the slip band thick ness; AA' = OC = OB. The method of determining the vertex coordinates for curvilinear plastic triangle KLM was presented in [3]. Thus, we may calculate the coordinates of any point on the boundaries of the plastic deformation zone by means of Eqs. (2)\u2013(5). In Fig. 3, we plot the boundaries of the plasticity zone for two values of the dynamic front angle. The initial data for the calculation are taken from the experiments in [4]", + " Decrease in \u03b3d is accompanied by expansion of the zone of secondary plastic deforma tion and broadening of the bands in the shear region. If the front cutter surface is nonplane (with a hard ening facet, a chip coiling groove, etc.), it is very diffi cult to obtain an analytical solution for the boundary slip lines. We must resort to numerical calculation, with step by step construction of an orthogonal slip line grid in the plastic zone [1]. The stress\u2013strain state in the cutting zone may be calculated by means of Fig. 2. Precise theoretical determination of the stress\u2013strain state is possible for a rigid\u2013plastic model of the blank, without harden ing. In that case, the resulting slip line field in the plastic region is clearly related to the stress state there. Thus, the change in the mean stress \u03c3meK along the slip lines is proportional to the corresponding angle of rotation [1, 5] (6) where L and K are points of the slip line; \u03c9LK is the angle of slip line rotation on passing from point L to point K; k is the maximum tangential stress in plastic deformation", + "ln= for the \u03b1 lines yd 2\u03bcrod\u03b1 zd+( ) 4hpl \u03b1cos( )3 2zd \u03b1cos( )2\u2013 3hpl \u03b1cos\u2013 hpl \u03b1sin\u2013+ 2 \u03b1 \u03b1cossin 1+ ;= for the \u03b2 lines yd 2\u03bcrohnd\u03b2 1 \u03b1tan( ) 2 +[ ]\u2013 1 \u03b1tan( ) 2 \u2013[ ]zd+ 1 \u03b1tan+( ) 2 ,= d\u03b1 hpl \u03b1cos( )2 hr \u03b1cos\u2013 \u03bcrohr \u03b1sin hr \u03b1sin\u2013 hpl \u03b1 \u03b1cossin \u03bcrohr \u03b1cos\u2013+ + hr \u03b1cos\u2013 zd \u03b1cos \u03bcrohr \u03b1sin hr \u03b1sin\u2013 zd \u03b1sin \u03bcrohr \u03b1cos\u2013+ + + \u239d \u23a0 \u239b \u239e ;ln= d\u03b2 hr\u2013 zd \u03bcrohr \u03b1tan hr \u03b1tan zd \u03b1tan\u2013 \u03bcr0hr+ + + + hr 1\u2013 \u03bcro \u03b1tan \u03b1tan \u03bcro+ + +( ) \u239d \u23a0 \u239b \u239e ;ln= \u03c3me K \u03c3me L 2k\u03c9LK,\u00b1= 2/ 3 RUSSIAN ENGINEERING RESEARCH Vol. 30 No. 2 2010 THEORY OF CONSTRAINED CUTTING 135 slip line, we may calculate the stress components for two dimensional plasticity theory [1] (7) where \u03c9 is the angle between the tangent to the slip line and the yd axis at the given point. We now determine the stress at the left boundary LKCBD (Fig. 2). At point L on the machined surface, = 0. (Correspondingly, is the compressive primary stress.). From the plasticity condition \u03c31 \u2013 \u03c32 = \u00b12k for this point, we find that \u03c3zL = \u20132k. At this point, the mean stress \u03c3meL = (\u03c31 + \u03c32)/2 = \u2013k. The inclina tion of the tangent to the \u03b1 slip lines is \u03c9L = \u03c0/4. On passing along the slip line from point L to point K, according to Eq. (6) Analogously, for points C, B, and D, the boundary \u03b1 slip line is \u03c3yd \u03c3me k 2\u03c9;sin+= \u03c3zd \u03c3me k 2\u03c9;sin\u2013= \u03c4ydzd k 2\u03c9,cos\u2013= \u23ad \u23aa \u23ac \u23aa \u23ab \u03c3ydL \u03c3zdL \u03c3me K k 1 \u03c0 2 2\u03d5K+ +\u239d \u23a0 \u239b \u239e .\u2013= \u03c3me C \u03c3me K 2k\u03c9KC+ k 1 \u03c0 2 2\u03d5K\u2013 2\u03d5C+ +\u239d \u23a0 \u239b \u239e ;\u2013= = \u03c3me B \u03c3me C k2\u03c9CB\u2013 \u03c3me C k2 yd B /s( );arcsin\u2013= = \u03c3me B \u03c3me C k2\u03c9CB\u2013 \u03c3me C k2 yd B /s( );arcsin\u2013= = \u03c3me D \u03c3me B k2\u03c7D,+= where \u03c7D is the inclination of the \u03b1 slip line at point D, which may be determined by taking the derivative of Eq. (4); s is the shear band thickness [2]. Passing along the \u03b2 line successively from point K to points N and M (Fig. 2), from point N along the \u03b1 line to point A, and then along the \u03b2 line from point A to point E, we may calculate the mean stress at the points on the right boundary of the plasticity zone. From Eq. (7), we find the stress components. The stress components at the left and right boundaries of the plastic zone are shown in Figs. 4 and 5 for the cor responding conditions in Fig. 3. Sharp change in the stress\u2013strain state of the blank is observed in the region of the plastic triangle KLM and at the cutter tip. We now calculate the contact stress at the front and read cutter surfaces. This is important in determining the cutting force and the heat released in friction. To this end, we divide the plastic contact section OE (Fig. 2) into n equal sections of length ydj = (lpl/n)j (j = 1, 2, \u2026, n). The coordinates of the resulting points are substituted into Eq. (1) for the \u03b1 slip lines, and then the constants of integration CIj are successively determined CI j yd j 2lf\u03bcfo lf ydj\u2013 \u03bcfolf \u03b3dtan\u2013 lf \u03b3dtan\u2013 yd j \u03b3dtan \u03bcfolf\u2013+( ) \u03b3d 2tan 1+( )ln\u2013 yd j \u03b3d 2tan\u2013 yd j+ 1 \u03b3dtan+\u2013( ) 2 .\u2013= Substituting the results into Eq. (1), we obtain n equations for the \u03b1 slip lines leaving the section of the front cutter surface adjacent to plastic contact of the secondary plastic deformation zone" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002844_ijtc2008-71265-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002844_ijtc2008-71265-Figure2-1.png", + "caption": "Fig. 2. Problem model", + "texts": [ + " 1 is considered as an equivalent to the Boussinesq problem for the elastic half-space. This loading assures the layer equilibrium. 1 Copyright \u00a9 2008 by ASME s of Use: http://www.asme.org/about-asme/terms-of-use Downlo The geometric and loading symmetry make the displacements on z direction vanish in the layer mid-plane, as well as the shear stress rz\u03c4 . This allows modeling the problem by a half-thickness (t) layer supported without friction by a rigid flat substrate and loaded by a normal force F as in Fig. 2. The boundary conditions associated with this model are: 0=z\u03c3 and 0=rz\u03c4 when 0=z (7) 0=w and 0=rz\u03c4 when tz = (8) In order to find the displacements and stresses in the layer of Fig. 2, an idea used by Timoshenko and Goodier [6] when solving Boussinesq\u2019s problem is applied. They used as the starting point the solution for a point force acting in an elastic space (the first type of simple solution). Then, they added supplementary stresses (the second type of simple solution) to make the shear stress vanish in the plane 0=z . Following this procedure, the starting point in the layer problem is Boussinesq solution for the half-space. Then supplementary displacements su and sw are added to Bu and Bw , respectively, such that the boundary conditions (7, 8) be satisfied" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001327_1.2080521-Figure3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001327_1.2080521-Figure3-1.png", + "caption": "FIGURE 3 \u2014 The configuration of sample and incident light in a measurement of the angular dependence of retardation.", + "texts": [ + ") as shown in Fig. 2. Soller slits of 0.48\u00b0 and 0.45\u00b0 vertical divergences were used for the source and detector, respectively. A divergent slit (1 \u00d7 10 mm) was used to control the exposure area of incident X-ray, while another divergent slit (0.1 \u00d7 10 mm) was used to limit the output signal from a sample. Retardation of samples at \u03bb = 589 nm was measured by using the parallel-Nicole rotation method (KOBRA-WR, Oji Scientific Instruments). The dependence of retardation on incident angles was investigated. Figure 3 illustrates the configuration of the sample and incident light in the retardation measurement. Samples were rotated about their in-plane slow and fast axes to measure the angular dependence of retardation. According to the theory of form birefringence media,6 the refractive indices of a layered stack of different materials are shown in Fig. 4 and are described as (1) where \u03b1 denotes the type of materials, n\u03b1 is the refractive index of the material and f\u03b1 is the volume fraction for material \u03b1. The first equation is for in-plane refractive indices and the second is for a refractive index in the thickness direction" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002130_1.3601533-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002130_1.3601533-Figure1-1.png", + "caption": "Fig. 1 Coordinates of tilting-pad journal bearing with three shoes", + "texts": [ + " Among these parameters are (lie mass and inertia properties of the pads and rotor, pad preload, pad mounting characteristics such as stiffness and damping, pivot characteristics such as friction moment and damping, and many others. The existence of parametric stability-boundary information would greatly alleviate the need for tilting-pad design information on the feasibility or preliminary design level. In addition to this basic need, there also exists a need for an economical, reliable, and easy-to-use analytical method for generating information on the detailed design level. The primary configuration under study in this paper is shown in Fig. 1. Although this figure shows a tilting-pad journal 1 Also Professor, Columbia University, N e w York, N. Y . Contributed by the Lubrication Division of THE AMERICAN SOCIETY OF MECHANICAL ENGINEERS and presented at the A S L E - A S M E Lubrication Conference, Chicago, 111., October 17-19, 1967. Manuscr ipt received at A S M E Headquarters, August 7, 1907. Paper No . 67\u2014Lub-8. -Nomenc la ture - Au,. = coefficient of L a G u e r r e polynomial C = ground-in clearance C'j = c o e f f i c i e n t of f po ly - nomial [see equat ion (32)] C\" = rad ius of p i v o t cii cle minus radius of shaft C\" = distance from origin of fixed coordinate system to pivot minus radius of shaft, d = radial distance from face of shoe to pivot point DT = liondimensional time differential e' = pivot circle eccentricity Fti fy \u2014 f o r c i n g f u n c t i o n [see equations (2) and (3) h = local film thickness Hu = liondimensional SF;j i = V ~ l I p = shaft polar moment of inertia I p ' = shoe moment of inertia about pitch axis VIS , Illy , 1 / = shoe moment of inertia about roll axis 11 = shaft transverse moment of inertia L \u2014 axial width of shoe Lk = A'th LaGuerre polynomial m \u2014 mass of shaft m' = mass of shoe m j = forcing f u n c t i o n s [see equations (6), (7), and (8)] (Continued on next page) 1 6 2 / J A N U A R Y 1 9 6 8 Transections of the AS M E Copyright \u00a9 1968 by ASME Downloaded From: http://tribology", + " It is assumed that the deviations from equilibrium are - N o m e n c l a t u r e - ma 1, IllaZ = Mi = n p = N = P = p = Pa = R = )SXi(T) + f I ' s x i ( T - r ) ^ 4 , , r J i ( \u00ab r ) e \u2014 dr. J o k (16) Figs. 5, 6, and 7 show a typical set of responses for the bearing in Fig. 1, and Table 1 gives the corresponding LaGuerre coefficients. In this case, the pivot circle eccentricity is zero and the symmetry of the situation can be exploited, thereby reducing the required number of separate responses to be computed. Treatment of the Dynamical Equations Three methods of solving the dynamical equations (11) have been explored. They are referred to as the \"Linearized Orbit\" method, the \"Lumped Parameter\" method, and the \"Characteristic Equation\" method. (a) The Linearized Orbit Method" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000240_6.2000-4164-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000240_6.2000-4164-Figure1-1.png", + "caption": "Fig. 1 Missile coordinate systems", + "texts": [ + " The initial guesses of the adjoint variables and the state variables required for solving the two point boundary value problem (TPBVP) are obtained from the analytic solutions solved by neglecting the missile aerodynamics. In addition, the mechanism of the optimal turn laws for the agile missile during the post launch phase will also be highlighted by the solutions obtained by neglecting the missile aerodynamics. A rigid body model for missile flight over flat earth is used to represent the missile trajectory dynamics in the present study. The equations of motion for the mathematical model of the missile flight in pitch plane are as follows5 (see Fig. 1): 9 = 0) M D ow nl oa de d by A U B U R N U N IV E R SI T Y o n Ju ly 3 1, 2 01 7 | h ttp :// ar c. ai aa .o rg | D O I: 1 0. 25 14 /6 .2 00 0- 41 64 u N Av, = \u2014 \u2014 sm 0 - \u2014 sin 9 - \u2014 cos 9 m m m u n N . Av = \u2014 cos<9 +\u2014cos#-\u2014 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, vv 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, Q the attitude of the missile body axis in degree, / the thruster location w", + " the center of gravity in m, 7 the moment of inertia in kgm , g the acceleration of gravity in m/s2. The parameter ./V is the aerodynamic normal force perpendicular to the missile body axis in Newton, A the aerodynamic axial force along the missile body axis in Newton, and MA is the aerodynamic pitching moment in Ntm. Furthermore, u is the side jet thrust normal to the missile body axis, the trajectory control variable in the present problem. The magnitude of the side jet thrust is constant but the thrust direction is to be determined. Figure 1 illustrates the missile coordinate system and the variable of interest. The missile equations of motion are expressed in terms of the Cartesian earth fixed frame oxEyE where the coordinate XE is in the local horizon plane. The form of the missile aerodynamic forces, moment and the definition of other related variables are given by the following expressions: = QSCN(a) (2) MA = where CN is the normal force coefficient, CA the axial force coefficient, XCG the center of gravity, and xcp the center of pressure due to the angle of attack" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001785_piee.1971.0221-FigureI-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001785_piee.1971.0221-FigureI-1.png", + "caption": "Fig. IB G'", + "texts": [], + "surrounding_texts": [ + "Existing formulas for the mutual inductance between parallel circuits with earth return, which are required for the calculation of interference, are either inadequate or difficult to apply in certain cases. The original formulas given by Carson have been modified and extended to give a more useful presentation in terms of the rectangular co-ordinates of the conductors, and Tables calculated by a computer are given for the ready application of the formulas.\nList of principal symbols xng \u2014 distance between conductors N and G parallel to\nthe ground plane m ys = distance of conductors N and G, respectively, from\nthe ground plane dng = actual distance between N and G\nDng = distance between N and image of G in the ground plane\n\\(y8\n6 = cot-1// /\n00\np\nfrequency angular frequency earth resistivity\n= (47rlO-7a;//>)1/2\nr X = Y=\nM =\nM\u00b0 =\nm = h =\n= ccDn\n(yg + yn) mutual inductance per unit length between parallel\nconductors with earth return mutual inductance per unit length with perfectly\nconducting earth component of mutual inductance due to the finite\nresistivity of the earth distance pertaining to the depth of penetration of\ncurrent in the earth = 659 (p//)1/2 metres\nIntroduction Experience over many years has established that the\nmethods used by power and telecommunication authorities for the prediction of levels of interference from fault currents in overhead power-transmission lines to adjacent telecommunication circuits provide ample accuracy for practical purposes. Where important discrepancies have occurred between prediction and subsequent measurement, they have been accounted for by insufficient foreknowledge of the magnitudes of the parameters concerned in the calculations, notably regarding the resistivity and possible stratification of the earth in the vicinity of the lines.\nOn the other hand, in the course of an investigation into the effects of interference to telecommunication circuits from 50 Hz load currents in 3-phase overhead transmission lines where small differences in induction effects relating to the individual phase conductors become important, it was found that the methods currently available1\"4 were unsuitable for the calculation of this type of interference, either because they they did not take adequate account of the height of the conductors above ground, or because the formulas were inconvenient for practical application. Accordingly more appropriate methods were developed, and this work has been completed with the asistance of a computer program and calculations made by the Headquarters Computer Branch\nPaper 6482 P, received 13th January 1971 Dr. Rosen is a part-time consultant to the Transmission Development & Construction Division, Central Electricity Generating Board, Burymead House, Portsmouth Road, Guildford, Surrey, England\nPROC. IEE, Vol. 118, No. 9, SEPTEMBER 1971\nof the Central Electricity Generating Board. The application to interference from 50Hz load currents in 3-phase circuits will form the subject of another paper. The formulas and Tables are somewhat extensive, and they may have useful applications to similar interference problems over a wide range of frequencies where small variations in the disposition of the conductors are important, e.g. induction from 2-wire power circuits and induction into open-wire telephone loops; it was accordingly thought desirable to present the formulas and Tables in the present paper. This is a condensed version of a CEGB report5 in which the derivation of the formulas is given in more detail than space permits here, and the Tables are more extensive with closer intervals between the arguments.\nCarson's formulas for mutual and self inductance\n2.1 Preliminary considerations Carson6 gives expressions for the mutual impedance\nbetween two parallel circuits with earth return, and for the self impedance of a circuit witn earth return, in terms of the corresponding impedances with perfectly conducting earth, plus functions of two dimensionless parameters r and 6; r pertains to the attenuation of electromagnetic waves propagated through the earth, and 6 is an angle defined by the geometry of the system. Equivalent expressions are given by Pollaczek.7\nReferring to Fig. 1A, which represents a cross-section through the parallel conductors G and N, assumed to be horizontal, and the surface of the ground assumed to be hori-\nzontal and plane, the X axis of rectangular co-ordinates is the straight line on the Earth's surface perpendicular to G and N, and the Y axis is the straight line through N perpendicular to the X axis. N is considered as fixed in position with coordinates (0, yn) metres, while G is a variable, with co-ordinates (xng, yg) metres.\n1227", + "In general, the surface of the ground is not a smooth plane, and by convention hypothetical supports are assumed to be in the plane of the cross-section. The X axis is then taken as the straight line through the feet of the supports, and the surface of the ground is taken as the plane parallel to the conductors and containing the X axis.\nThe quantities r, X and Y may be considered as 'electrical lengths', as distinguished from the corresponding physical lengths, Dng, xng and yg + yn.\nConductors on sloping ground\nWhere the feet of the supports are on different levels, as in Fig. 1B (exaggerated for clarity), the co-ordinates of N and G on the Y axis, namely yn, yg, are not the same as the respective heights hn, hg, and xng is not the same as the distance between the feet of the supports; however, in practice, the differences are usually trivial. If a conductor lies below the X axis, as in the case of a buried cable, the corresponding Y ordinate has a negative sign. (Note that the actual values for ym yg, hn, hg, used in calculations for aerial lines are less than the dimensions at the supports, as allowance has to be made for the sag of the conductors.)\nFrom Figs. 1A and 1B,\ndng = distance between iV and G,\n= W* + &g ~ yn)1)\"2 metres . . . . (1)\n= (1 + v2yi2xng metres (2)\nDng = distance between N and G', where G' is the image of G reflected in the conventional ground plane,\n= {xlg + (yg + y n ) 1 } \" 1 m e t r e s . . . . ( 3 )\n= ( 1 + \u00ab 2 ) 1 / 2 x n g m e t r e s ( 4 )\nand 6 = s\\n~l(xng/Dng) = cot~lu radians . . (5)\nwhere u = \\(yg + yn)/xng\\ (6)\n\u00ab = \\(yg - yn)lxng\\ (7)\nThe parameter r is defined by\nr = ccDng (8)\nwhere a = (47rlO~7aj/p)l/2 per metre (9)\n= 2-810. lO-^Z/p)1 ' 2 per metre . . . (10)\n(X) = 277/\n/ = frequency, Hz\np = earth resistivity, Qm\nThe term a is the modulus of the propagation coefficient per metre for plane electromagnetic waves propagated through the earth, and is the square root of the a used by Carson.\nIn practice, r and 6 are less convenient to use as parameters than those based on the rectangular co-ordinates, as shown in Fig. 2, namely,\nX = axng = r sin 0 (11)\nY= ,) and (x2, y2) respectively, and the parallel conductor be N, located at (0, yn).\nIf M(xltyi) is the mutual inductance per unit length between circuits 1-E and N-E, and M(x2, y2) the corresponding mutual inductance between circuits 2-E and N-E, then the mutual inductance per unit length between the circuits 1, 2 and N-E, 8M is\n8M = M(x{, yx) - M(x2, y2) (28)\nIf conductors 1 and 2 are electrically close to N, 8M can be calculated directly from the values of M(xu y{) and M(x2, y2). However, when the distance between 1 and 2 is small compared with their separation from N, the values of Af(jc1,.j>i) and M(x2, y^) approach each other, so that calculation to a large number of significant figures is necessary to determine 8M in this way. In this case, 8M can be obtained more easily in terms of the partial differential coefficients of M with respect to x and y, and the difference terms 8x and 8y where\n8x = Xj \u2014 X2 8y = y{ - y2 (29)\nWhen M(xu y{) and M(x2, y2) are expanded by means of Taylor's theorem,5 the resulting expression for 8M contains only odd powers of 8x and 8y, and only odd partial differential coefficients of M with respect to x and y, provided that the differential coefficients refer to the point whose coordinates are (x, y), where\nX = y = + (30)\nIn most cases, the terms involving the cubes and higher powers of 8x and 8y, and the third and higher differential coefficients can be neglected, whence\n= 8X\u2014 + oy Tix\n(31)\nWe may write\n8M = 8M\u00b0 + 8m (32)\nwhere\n8M\u00b0 = component for earth of perfect conductivity\na W \u00b0 , a= 8x -^\u2014 + 8y 7)x\n(33)\nand 8m = component due to the finite resistivity of the earth\n~bm\nFrom eqn. 17, one obtains5\n200(M-V)\n(34)\nx(l + \u00ab2)(1 + v2)\n{\u2014 (M + v)8x + (1 \u2014 uv)8y}\nmicrohenrys per kilometere . . . (35)\nwhere u = (y + yn)jx \\ v = {y-yn)lxj\n1229" + ] + }, + { + "image_filename": "designv11_69_0003918_pecon.2010.5697626-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003918_pecon.2010.5697626-Figure2-1.png", + "caption": "Fig. 2 Effect of Voltage Vector on Stator Flux and Torque", + "texts": [ + "()0()( 0 s t ssss dtiRV \u03bb\u03bb +\u2212= \u222b From Fig.1, it is observed that the derivative of stator flux reacts instantly to changes in the stator voltage, the respective two space vectors are separated in the circuit by the stator resistance Rs only. However, the vector derivative of rotor flux \u2018\u0440\u03bbr\u2019 is separated from that of stator flux by stator and rotor leakage inductances. Therefore the reaction of rotor flux vector to stator voltage is somewhat sluggish in comparison with that of stator flux vector. In Fig.2, at a certain instant t, the initial vectors of stator and rotor are denoted by \u03bbs(t), \u03bbr(t) respectively. After a time \u0394t , the new stator voltage vector \u03bbs(t+\u0394t) differs from \u03bbs in both magnitude and direction. As time \u0394t is very small the changes in rotor flux has been negligible. The stator flux has increased and torque angle \u03b8sr has been reduced by \u0394 \u03b8sr . Thus, appropriate selection of inverter states allows adjustments of both the strength of magnetic field of motor and developed torque. The basic premises and principles of the direct torque control methods can be formulated as follows" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000170_02ye9034-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000170_02ye9034-Figure1-1.png", + "caption": "Fig. 1. The sketch of heat flux distribution at the weldpool surface.", + "texts": [ + " Thus, the partial arc heat flux is assumed to distribute in a Gaussian model along the deformed weldpool surface near points O1 and O2 q l ( s ) = qlmexp - , near point O1, (1) q2(s ) = q2mexp - 2 ~ ' near point 02, (2) where qlm and q2m are the maximum heat flux at points Ol and 02, al and ~2 are the distribution parameters, and s is the length of curve segment between O1 or 02 to some point on the deformed weldpool surface. The longer the distance from the wire tip to some point on the weldpool surface, the less the heat flux at this point and the larger the value of distribution parameter. Thus, there are the following assumptions : qlmLl = q2mL2, (3) L1 L2 (4) ~1 ~2 r 1 + r 2 = d , ( 5 ) where d is the total diameter of hot spot in GAM welding, which is based on the principle that the experimental measurements and the predicted results are in agreement. And the meanings of other symbols are shown in fig. 1. As shown in fig. 2, under the action of arc pressure Pa, droplet impact Pd, gravitational force pg~ of liquid metal and surface tension )' , the shape of the weldpool surface is deformed. In the coordinate system of 01-xyCP, the governing equation of the weldpool surface deformation is P a - pg~ + Pd + 2 = - )' ( i + ~ ) ~ x x - 2~x~y~xy (1 + q52x + q~2)3/2 (6) where p is liquid density, g is gravity acceleration, ~ is Lagrange muhipler, 9 is configuration function of weldpool surface, q~ with subscripts represents partial derivative of q~ to subscripts, for example, ~x = 3 q~/3 x", + " According to the coordination on the weldpool surface and the coordination of the wire tip W(0, 0, Zo), the points of O1 and 02 can be determined. In the same way the points of 03 and 04 on the weldpool surface curve at the x = 0 section can also be determined (fig. l ( b ) ) . Then 0 1 ( X l , 0, H - q0(xl , 0 ) ) , 02(x2, O, H - ~ ( x 2 , 0 ) ) , 03(0 , Yl, H - qO(0, y ~ ) ) , 04 (0 , - y t , H - ~ ( 0 , - Y l ) ) , where H is the workpieee thickness. The point O' corresponding to the maximum depression is with the following coordinate: (0 , 0, H - ~ ( 0 , 0 ) ) . According to the geometric relation shown in fig. 1 ( b ) , eq. (5) may be written as 2\"/-6atxl 2\"f'60\"2 I x21 + = d . (7) 4x . 0) - 0 ) ] 2 4x . E (x2, 0 ) - 0 ) ] 2 The values of al and a2 will be achieved based on the simultaneous solution of eqs. (4 ) and (7) . As shown in fig. 1 ( b ) , if the points O1, O2, 03 and 04 are not in a same horizontal plane, the average coordinate value in z-direction of the four points is taken as the z-coordinate value of the distribution origin of arc heat flux, ( i ~ ( X l , 0 ) + ( ~ ( X 2 , 0 ) + ( ~ ( 0 , Yl) + tP( 0, - Yl) (8) z = H - 4 That is, along the deformed weldpool surface all the points whose coordination values in z-direc- tion arc the value of z ~ are connected and form a closed curve as shown in fig. l ( b ) . A series of the distribution origins of partial arc heat flux are all on the closed curve" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000672_0025-5416(73)90005-0-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000672_0025-5416(73)90005-0-Figure1-1.png", + "caption": "Fig. 1. Wedge disclination loop in an elastic half-plane.", + "texts": [ + " It was found that if a loop is pinned down beneath a free surface, it can assume an orientation of lowest strain energy. The aim of this paper is to extend the above investigation to a more complicated case, namely, a wedge disclination loop. The term \"wedge disclination loop\" was first used by Li and Gilman 5 in their study of deformation configurations in polymers. It is used to denote a disclination loop where the rotation axis lies in the plane of the loop. Hence, it is understood that loops of this type bear some twist characteristics. In Fig. 1 we demonstrate the elastic half-plane and the geometry of the loop under consideration. The center of the wedge loop with radius a and angle of rotation co is situated at a distance z0 below the free surface. The orientation of the loop is defined 164 by the angle ~ m a d e by the n o r m a l h to the loop and the z axis. The axis of ro t a t i on is a s sumed to be parallel to the y axis. The shear m o d u l u s and Po i s son ' s ra t io of the elastic m e d i u m are deno t ed by # and v, respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002076_2008-01-1044-Figure4-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002076_2008-01-1044-Figure4-1.png", + "caption": "Figure 4: Connecting rod free body diagram", + "texts": [], + "surrounding_texts": [ + "Computational models are very important for every aspect of engineering design, so the optimal operation of the piston, and the internal combustion engine as a whole, greatly depends on them. The piston, during its operation, apart from the axial reciprocating motion experiences small transverse oscillations. The identification of the magnitudes of these oscillations and the ability to control them is very crucial, as the piston performance depends on them. In the process of this identification computational models are utilized. Piston computational models were developed in the early 1980\u2019s. Li et al. [12] developed an automotive piston lubrication model to study the effects of piston pin location, piston-to-cylinder clearances and lubricant viscosities on piston dynamics and friction assuming a rigid piston. Li [11] considered the elastic deformation of the piston skirt; integrating this with hydrodynamic lubrication has formed the elastohydrodynamic lubrication analysis which is considered by most of the recent efforts [2, 3, 4, 5, 8, 9, 13, 22, 24, 25]. All these have contributed in better understanding the piston secondary dynamics problem. Zhu et al. [24, 25] were the first to consider the elastic deformation of the cylinder bore. In more recent years Duyar et al. [4] used the mass-conserving Reynolds equation to solve for the hydrodynamic pressures developed on the skirt. This method predicted lower hydrodynamic pressures which allowed for higher transverse motion of the piston and consequently higher contact forces. All these models however assumed that the effects of piston motion along the wrist-pin and land interactions with the cylinder bore are negligible. The authors have recently introduced [15] a new piston dynamics model that considers translation along the wrist-pin and rigid second land interactions with the cylinder bore. They modeled a conventional gasoline piston and demonstrated that motion along the wrist-pin becomes important in predicting piston wear, especially when the cylinder bore deformation and temperature distributions are asymmetric. In this work, the authors developed the above model further to consider an elastic second land. In the current model it is assumed that pressures due to lubrication or scuffing have no effect on the second land deformation, however, it deforms due to combustive, inertial and thermal loads. Only the second land was chosen to be modeled as it is, traditionally, the land with the larger diameter, thus it is expected to have the most dominant interactions with the cylinder bore. Also the piston rings are assumed to have negligible effects on the piston dynamics. The authors utilize a new generation gasoline piston with uneven thrust sides to demonstrate the differences on piston motion and piston wear between different modeling approaches." + ] + }, + { + "image_filename": "designv11_69_0001538_20060906-3-it-2910.00121-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001538_20060906-3-it-2910.00121-Figure2-1.png", + "caption": "Fig. 2 Schematic diagram of the left side of ARAS", + "texts": [ + " The net body motion is the composite effect of all wheels and can be obtained by combining (3) into a single matrix equation as \u239f \u239f \u239f \u23a0 \u239e \u239c \u239c \u239c \u239d \u239b =\u239f\u239f \u23a0 \u239e \u239c\u239c \u239d \u239b \u239f \u239f \u239f \u23a0 \u239e \u239c \u239c \u239c \u239d \u239b \u03b7 \u03d5 \u03d5 & & & & & M c c r r 6 6 u J u I I (4) where the composite identity matrix on the left is nn\u00d76 , ( )Tcn2c1cc uuuu &L&&& = and ( )Tcn2c1cc \u03d5\u03d5\u03d5\u03d5 &L&&& = are 1n3 \u00d7 vectors of composite wheel velocities at the contact points, and \u03b7& is the 1\u00d7\u03bd vector of the joint variables which has both active (actuated) and passive joints. Note that in general some wheels share common suspension links and joints so that \u2211 = \u03bd\u2264\u03bd n 1i i . The composite Jacobian matrix of the rover J has a dimension of )6(6 \u03bd+\u00d7 nn . 2.1 Example The articulated rover with active suspension (ARAS) to be considered here is similar to the JPL Sample Return Rover shown in Fig.1. The schematic diagram of ARAS to be analyzed is shown in Fig. 2. The rover has four wheels with each independently actuated and rotation angles subscripted with a clockwise direction so that 1\u03b8 , 4\u03b8 are for the left side and 32, \u03b8\u03b8 are for the right side. At either side of the rover, two legs are connected via an adjustable hip joint. In Fig. 2 the hip angles on the left and right sides are denoted as 12\u03c3 and 22\u03c3 , respectively. These joints are actuated and used for balancing the rover. The two hips are connected to the body via a differential which has an angle \u03c1 on the left side and \u03c1\u2212 on the right side. On a flat surface \u03c1 is zero but becomes non-zero when one side moves up or down with respect to the other side. The differential joint \u03c1 is passive (unactuated) and provides for the compliance with the terrain. The wheels are steerable with steering angles denoted by i\u03c8 " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001308_ls.11-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001308_ls.11-Figure1-1.png", + "caption": "Figure 1. Schematic of double-decker high-precision bearing.", + "texts": [ + "8 This work is taken up to analyze theoretically the axial deflection of DDHPBs under the influence of axial loads and its comparison with that of the conventional deep-groove and angular contact ball bearings of the same bore and outer diameter. Two different approaches have been adopted, namely the equivalent diameter approach and the summation approach. This leads to an evaluation of the performance characteristics of DDHPBs and equivalent angular contact and deep-groove ball bearings. The basic design of the DDHPB is based on the use of two series of rolling-elements one riding on another and separated by an intermediate rotating race as shown in the Figure 1. The intermediate rotating race acts as an inner race of the second series of rolling-elements and outer race of first series of rolling-elements. The outer race of the DDHPB is mounted in the housing like the outer race of a conventional bearing. The inner race of the first series of rolling elements of a DDHPB rotates at the shaft speed, and the intermediate race rotates approximately at one-fourth of the shaft speed due to frictional torque acting on it from both series of rolling elements.1,2 The decrease in operating speed of the intermediate race and uniform distribution of load on both series of rolling elements improves the performance characteristics of the DDHPB and reduces the energy loss in it as compared to that of conventional bear- Copyright \u00a9 2006 John Wiley & Sons, Ltd" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003616_detc2009-87503-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003616_detc2009-87503-Figure1-1.png", + "caption": "Figure 1. Section 472-2 Paper Boat sketch with high level of skill and detail", + "texts": [ + " Level 3 sketches display product form and may contain shading and brief annotations. Level 4 sketches show product form with annotations, illustrations of features and detail, and may include dimensions. Level 5 sketches are the most detailed sketches that are two or three dimensional drawings to display total product form with all dimensions and annotations. Three sketches are shown here to demonstrate the variety in student sketching skills. They are all from the 472-2 section and the Paper Boat assignment. The sketch in Figure 1 portrays an entire artifact (F1) with specific features. This sketch is one of the rare McGown level- 4 sketches, which is a high skilled level sketch. The entire artifact sketch (F1) signals overall thought on the entire system and how parts will interact with one another and how it will be used. The sketches in Figure 2 show more typical paper boat depictions. The sketch on the left of Figure 2 is a McGown level-1 sketch and the one on the right is a level-2 sketch. Copyright \u00a9 2009 by ASME Downloaded From: http://proceedings" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003596_amc.2010.5464064-Figure3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003596_amc.2010.5464064-Figure3-1.png", + "caption": "Fig. 3. ZMP and IZMP", + "texts": [ + " We represent the condition of the internal force as the condition of the operation. This condition is derived with the position and posture of the robot and its manipulators, status of the objects, and friction coefficient. The robot implements torque control to generate the internal force. We use IZMP(Imaginary Zero Moment Point[10]) of the target object to consider the rolling of an object. IZMP is not proper ZMP since ZMP is approximated that the robot\u2019s foot is fixed on the floor. But, in this assumption, the IZMP can locate outside of the support polygon. Fig. 3 shows single mass robot model. IZMP is the intersection of the floor and the extended line of the total force of the gravity and the inertia 978-1-4244-6669-6/10/$26.00 \u00a92010 IEEE 601 force(Fig. 3). Usually, ZMP indicates the center of rotation of the object. When the IZMP is in the support polygon of the object, it is the same position of ZMP. And the moment around it becomes zero. However, if the IZMP locate the outside of the support polygon, the moment around ZMP occurs since ZMP cannot be in the outside of the support polygon by the definition. So, the IZMP can indicate the stability of the object. In other words, if the IZMP doesn\u2019t exist in the support polygon, the object rolls around ZMP" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000282_tmag.2003.810347-Figure4-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000282_tmag.2003.810347-Figure4-1.png", + "caption": "Fig. 4. Analysis model (except the air).", + "texts": [ + " The calculated distribution of magnetization vector in the permanent magnet is shown in Fig. 3. 0018-9464/03$17.00 \u00a9 2003 IEEE A cogging torque analysis is performed for the model whose magnetic characteristic in the root of the claw pole has been deteriorated. The influence, which the deterioration of the magnetic characteristic of the claw pole has on the cogging torque characteristic or the distribution of magnetic flux density, has been discussed. The analysis model except the air region is shown in Fig. 4. From Fig. 4, it is formed by resin between the permanent magnet and the shaft, it is assumed that the magnetic flux did not distribute at the shaft, and the rotor portion of the model is considered the only permanent magnet. Three-dimensional meshes are shown in Fig. 5. The magnetization (B-H) curves before and after annealing used in this analysis are shown in Fig. 6. As the measurement of the magnetic characteristics, after manufacturing the stator, the B-H curve is directly measured by the coil winding the stator", + " That is, by taking into consideration the deterioration of magnetic characteristic in the root of the claw pole, the magnetic flux density of the inside of root of the claw pole becomes small, and the amplitude of cogging torque also becomes small. The discretization data and CPU time for the cogging torque analysis is shown in Table I. A holding torque is calculated when the direct current is passed in both coils. We discuss the influence which the deterioration of magnetic characteristic of the claw pole has on the holding torque characteristic. The direct current of 235 AT is supplied to each coil in the same direction. In addition, this analysis is carried out using the analyzed model shown in Fig. 4, meshes shown in Fig. 5, the B-H curve shown in Fig. 6, the magnetization of permanent magnet shown in Fig. 3, and the analysis condition shown in Fig. 7. The contours of the magnetic flux density in the root of the claw pole on mechanical angle 10 are shown in Fig. 11. From Fig. 11, the same tendency in as the case of cogging torque is shown. The holding torque characteristic in Type 0 is shown in Fig. 12. This figure shows that one cycle is 30 . Although not illustrated, the holding torque characteristics of Types 1\u20134 are also the same waveform" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001979_50037-5-Figure7.92-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001979_50037-5-Figure7.92-1.png", + "caption": "Figure 7.92 Configuration of the contact.", + "texts": [ + " From ANSYS Main Menu select General Postproc \u2192 Plot Results \u2192 Contour Plot \u2192 Nodal Solu. The frame shown in Figure 7.90 appears. In the frame shown in Figure 7.90, the following selections are made: [A] Contact and [B] Pressure. These are items to be contoured. Pressing [C] OK implements selections made. In response to that an image of the cylinder with pressure contours is produced as shown in Figure 7.91. 7.2.3 Wheel-on-rail line contact Configuration of the contact to be analyzed is shown in Figure 7.92. This contact problem, which in practice is represented by a wheel-on-rail configuration, is well known in engineering. Also, the characteristic feature of the contact is that, nominally, contact between elements takes place along line. In reality, this is never the case due to unavoidable elastic deformations and surface roughness. As a consequence of that, surface contact is established between elements. This is a 3D analysis and advantage could be taken of the inherent symmetry of the model. Therefore, the analysis will be carried out on a quarter-symmetry model only" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001967_978-1-4302-0088-8-Figure1-5-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001967_978-1-4302-0088-8-Figure1-5-1.png", + "caption": "Figure 1-5. A CubeBot", + "texts": [ + " At the back of the book, I\u2019ve included a quick reference for PBASIC as well as a Javelin Stamp version of the examples. You can download the latest version of the programmer from www.parallax.com. Figure 1-4 shows a picture of a sample program loaded into the BASIC Stamp Windows editor. If you do not have a robot and would like one in a kit, several fine specimens can be found at www.lynxmotion.com, www.parallax.com, or www.prestonresearch.com. For most of the examples in this book, I use a differential drive robot or CubeBot (as shown in Figure 1-5). C H A P T E R 1 \u25a0 A P R I M E R 5 Hopefully you have everything you need to get started. I\u2019m now going to start by explaining the thought process behind robotics programming (in other words, getting your robot to do stuff). Then I\u2019ll talk about some concepts in Java you should know about before embarking on a robotics project. Finally, I\u2019ll show you an example of how to begin modeling your software in a way that\u2019s both easy to use and effective. What does your robot do? This is the number one question people ask when I tell them I have a robot" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002675_001872086801000408-FigureI-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002675_001872086801000408-FigureI-1.png", + "caption": "Fig. I . Scllemotic represeiitafion of cnnicrn nioiuiled oii bipodso ns to nrc over visiidfiefd.", + "texts": [], + "surrounding_texts": [ + "There are two distinct phases in the work of producing induced stereoscopic motion by moving pictures. One is the production of the appropriate film. The other is the projection of the pictures. Each of the two phases has its own particular difficulties and techniques and these in a way require separate discussion. However, the difficulties of one phase tie in with the difficulties of the other phase. Therefore, for this review of the work it seems best to discuss both phases together in our attempts to produce IShl. Two attempts were made to use film taken from aircraft. In the first of these attempts film already in existence was tried. The angle of the camera was in the neighborhood of 4S0 below the horizontal. Two copies of a single film were run backward and forward in two Eastman Analyst projectors. A difference in the presence and absence of a stereoscopic effect could be detected, but there was little ISM as such. Several problems were evident. The changes in stereo-base were too slow; the changes of camera points could not be stabilized much less duplicated; there were other depth cues in the photographs which confused the issue; resolution in the film was poor. In a second attempt, a flight especially organized for obtaining film was made. An attempt was made to obtain vertical views. However, a number of practical difficulties made the film unsuitable for producing ISM. The camera mount was unstable. Targets moved across the field of view so rapidly that only a small stereobase could be obtained before the target moved off the field of view. Finally, the resolution of the film again was poor. These attempts were made at different times during the course of the work and convinced us that it was more expedient to solve problems in the laboratory before trying again to use actual aerial photographs. It became obvious very early that a rear view screen was necessary with the stereoscope; this caused no difficulty. I t also became obvious that a closely matched pair of projectors which could be stopped and reversed quickly without changing light intensity of the projected images, and which could be synchronized, was essential. This was a serious difficulty, which was finally overcome by the use of University of Arizona funds, for which we are grateful. There are roughly two methods both for producing the film and for projecting it. One can be called, in general terms, the continuous method and the other the cycling method. The same method does not necessarily have to be used both in producing and in projecting the film. However, in our thinking, at first, the two did go together. The continuous method of producing the film is what we call the single pass method because it is quite obvious that a stereopair cannot be produced by two passes of a plane over terrain. In the laboratory two passes can be made only under controlled conditions in which the second pass is exactly the same as the first. In all attempts to date to develop a single pass manner of producing film, we have used cameras in fixed position, i.e. either one camera pointed downwardly or two cameras pointed fore and aft about 12O from vertical. The reasons for this policy of fixed position cameras, and it has been a policy, are two: first, there is the dificulty of turning a camera on a moving platform to keep it precisely pointed to one target on the ground; second, in surveillance work one may not know in advance what one is searching to find and so where to point the camera. \\Vith fixed cameras on the other hand the films are not difficult to take, at least in the laboratory. The cameras are mounted in a fixed position and the targets are carried beneath them on a belt. Difficulties arise, however, in projecting the continuous roll of film. The two projectors have to seesaw the film backwards and forwards to get the ISM phenomenon. This creates several difficulties. Two compensatory actions must be provided. First, the movement of objects across the field of view of the projected image must be counteracted by turning both projectors to hold images of objects at n constant position in the stereoscopic view. The eyes can converge and diverge somewhat to compensate for varying lateral displacements of objects in the stereo fields, but such movements are relatively slow and of limited extent. Second, there is enough play between the film perforations and the framing teeth of the projectors to allow a shift of the image on the screen at each reversal and this shift is sufficient to upset the stereoscopic effect. A complex bit of instrumentaiion (which rather impressed our friends) was put together by Slingerland (1967) to provide for these compensations, and to control the timing of these compensations and the reversals of the two projectors. This instrumentation was reasonably successful and gave good results for short periods of time. However, more critical troubles arose, The film would not stand up under continuous reversing and repeatedly tore. Components in the projectors and timers burned out under continuous reversing. There were also other problems connected with the projection of continuous reels. The outcome was that with some reluctance because of giving up the favorable features of using continuous reels, we turned to the cycling method. The cycling method turned out better. A bipod which carries the camera stands astride a single fixed field of view (Figure 1). The camera system is driven in an oscillating arc over the field of view. The camera is always pointed to the same object in the center of the target field. Several variations in this method were tried, with the most successful being as follows. The camera oscillated to one side to produce the film for one projector and then to the other side to produce the film for the other projector. For comparison films with fixed stereo-base, the camera is put in the extreme position, right or left, and run in that position for an appropriate time. The finished film is cut and spliced into short loops. The essential feature of the cycling method of projecting the film is the use of small loops of film. Many problems arising in the continuous method simply vanished with the cycling method. TWO disadvantages remain in the method. A pair of film loops cover only one field of view, and they have to be changed in both projectors before another field of view can be examined. Second, the film must be spliced and the splice, while not serious, is some source of distraction to the viewer. Better splicing methods may be of help here. THE EXPERIMENTAL TEST The test to be reported was made with film produced and projected by the cyclical method. I t is a laboratory test, and more work needs to be done before it can be used in the field, if indeed it can be. The most promise we see at this time is a film produced from a single pass over terrain but using some cycling technique and then projected by the cycling method. An experiment is getting under way along this line. Filtii Pmfrrction - In the experiment reported here the bipod supported the camera 48\u201c above the target field. The throw was 6\u201d to each side of center giving an extreme apex angle of a little m&e than 14\u201d. The cycling rate was about 1/2 per second. All photographs used in the experiment were taken with an H16S Bolex 16mm motion picture camera using a Switar lA.4, 25mm lens and fine grain Kodak XT Panchromatic film (ASA20). Negative film rather than reversal film was used in order to have available exact duplicate copies during the test runs. The camera was driven at 24 frames per second by a Bolex Electric Motor, which provided constant speed and remote control during movement of the camera. Lighting was diffuse so as to eliminate shadows and the extra cues that would be provided by shadows. The background of the targets was a layer of pea gravel, which provided a randomly mottled surface. The targets were mounted on small squares of cardboard to keep them upright, and the gravel was used to weight down and to cover the cardboard. At no time were the cardboards visible and thus they offered no cues to target location. There are a number of variables that contribute to the ease, or difficulty, of target location. Since the purpose of the experiment was not to examine these variables, but rather to compare ISM to fixed base stereoscopy, no completely systematic program of target construction and use was followed. Four kinds of targets were made using cut nails as the central support. These \\\\.\u2018ere gravel-coated nails, vermiculite-coated nails, uncoated nails, and horizontal thread supported on two uncoated nails. These targets were varied in height from 1 to 2 inches, and the gmvelcoated targets were varied in width from 1/S to 1 inch. On a judgmental basis 9 targets were selected to provide a range of difficulty. The gravel targets were judged to be easier to locate because they were bulky; the uncoated nail and the horizontal thread were judged to be more difficult because they were slender; the final difficulty of the vermiculite targets was not prejudged but they were called \u201cmedium.\u201d The data confirni that there was such a range and that the difficulty of the two kinds of targets was as judged. Myers (1965) had found that targets were more easily located if they were in the central region of the field of view than if they were in a more peripheral region. For this reason it was planned to counterbalance the location of each target by placing it once in a central region and once in a peripheral region. A rectangular grid was used to locate targets. On such a grid and with different directions, possibly having different weights, precise balancing is impractical but a rough balancing was carried out. As between conditions of search (ISM versus fixed stereo-base) the target difficulty due to location was completely balanced because the same targets at the same location were used for both conditions. The location grid was six squares wide and four squares high. This provided 24 target locations, only 16 of which were used in the experiment. The procedure for taking the film was as follows. A target was positioned in the field of view by the use of a removable grid, which was withdrawn before taking photographs. The bipod and camera were started in their oscillating motion between the midpoint and one extreme end. During this action the camera was started and stopped to get slightly more than enough frames for three complete cycles of oscillation. The bipod was then set stationary at the extreme end and the same amount of film was run in that position for the comparison loop of fixed stereobase film. With the target and its background untouched, the same sequence was run on the other side and also a t the other extreme end. When processed the fiIm was cut and spliced to make loops about 150 frames in length. For any target the pair of ISM films of course had to be exactly the same length to maintain the syncrony in projection." + ] + }, + { + "image_filename": "designv11_69_0003425_0926-6577(64)90108-1-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003425_0926-6577(64)90108-1-Figure2-1.png", + "caption": "Fig. 2. C h a m b e r to record spec t ropho tomet r i c changes in lobster hea r t (explana t ion in the tex t ) .", + "texts": [ + " In addition, the absence of hemoglobin and myoglobin in this animal makes it an ideal material for the study of cytochrome changes, as has been shown for many marine invertebrates by TAPPEL 2. BALTSCHEFFSKY a (Expt. H-6). The rest of the spectrum is very similar to that of the amphibian heart I and of mammalian heart mitochondria 3. The kinetics of the oxidation-reduction changes of cytochromes as a consequence of modifications in mechanical activity were studied with a special chamber holding the whole heart (Fig. 2). For this purpose the heart strips could not be used as in the case of amphibian heart because the integrity of the intracardiac ganglia in the heart of crustacea is necessary for it to be stimulated electrically. The holder consisted of a stretcher with two hinged stainless-steel rings to maintain the dorsal and ventral heart walls at a constant distance. The stretcher was put inside the heart with the two rings together through the caudal opening, Biochim. Biophys. 3cta, 88 (1964) 648-65 \u00b0 SHORT COMMUNICATIONS 649 s t i mu l a t i on f requency " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000694_b:tels.0000029042.75697.f0-Figure17-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000694_b:tels.0000029042.75697.f0-Figure17-1.png", + "caption": "Figure 17. The topology of the 4 aircraft DDF network when post processing the flight data.", + "texts": [ + " The aircraft pose was obtained by fusing the logged GPS/IMU information, and tracking information was obtained using vision sensor which logged frames at 50 Hz. Artificial vision targets (0.9 m \u00d7 0.9 m white plastic sheets) were deployed at surveyed locations in order to allow the camera to track known objects. The data sets were used initially with all four aircraft acting independently and not sharing any information. The entire process was then repeated with the aircraft configured in a DDF network in the topology illustrated in figure 17. All data association was done using the information gate [Fernandez, 10], which is the information form of the state space innovation gate. Results of running in these two configurations are presented in section 7.1. When the platforms operated independently, they each generated target tracks using only the information from their locally attached sensor. The maps of 50 targets generated by each aircraft under these conditions are plotted in figures 18 and 19. Each of these maps is completely independent as they were generated without any sharing of information at all" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002319_20080706-5-kr-1001.02701-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002319_20080706-5-kr-1001.02701-Figure1-1.png", + "caption": "Fig. 1. Vehicle reference frames", + "texts": [ + " The distance from the ocean bottom is assumed to be known (for example measured with a Doppler velocity log (DVL) in its usual bottom-looking configuration). An Extended Kalman Filter is used to estimate the remaining state of the system. In the next section the vehicle model equations are stated. In the third section a general formulation for the model predictive control is described. In the fourth section a state dependent nonlinear weight is defined, and its benefits are explained. In the fifth section the simulation results are presented and then conclusions are given in the sixth section. Figure 1 shows a sketch of the vehicle with the reference frames. The vehicle dynamics is given in the body-fixed frame defined by Cxz. The surge and heave speed are defined by u and w, respectively. The pitch angle rate is denoted q. In the earth-fixed reference frame OXZ we describe the motion of the vehicle as shown in (2). The attitude of the vehicle is defined by \u03b8c. The vehicle dynamics is described, according to Sutton and Bitmead (1998), as MI ( u\u0307(t) w\u0307(t) q\u0307(t) ) = mq(t) ( 0 \u22121 0 1 0 0 0 0 0 )( u(t) w(t) q(t) ) +u(t)Dh ( u(t) w(t) q(t) ) +\u0393g(t) + Ucw(t) + Ucp(t) (1) x\u0307c(t) z\u0307c(t) \u03b8\u0307c(t) = ( cos(\u03b8c(t)) sin(\u03b8c(t)) 0 \u2212 sin(\u03b8c(t)) cos(\u03b8c(t)) 0 0 0 1 )( u(t) w(t) q(t) ) (2) where MI is the inertia matrix including the hydrodynamic added mass, m is the vehicle mass, Dh is the damping matrix, and the buoyancy term \u0393g(t) is zero because the vehicle is assumed to be neutrally buoyant" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001961_s1387-2656(06)12004-9-Figure3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001961_s1387-2656(06)12004-9-Figure3-1.png", + "caption": "Fig. 3. Microfluidic chip devices of a protein-incorporated planar lipid bilayer [42]. The channel width and depth are 500 and 100 mm, respectively. The aperture size is 50\u2013200 mm. The bottom channel is filled with buffer. The surface of the buffer stays at the front of the aperture because of the surface tension. The lipid solution is introduced to the upper channel and flushed with air and a thin layer of lipid monolayer is formed over the buffer surface at the aperture. The bilayer is then formed by passing either the same or different lipid solution through the upper channel at low flow rates of 0.1 mL/min.", + "texts": [ + " Reflectivity [38], laser-intracavity absorption [39], steady-state and time-resolved fluorescence [17], holography [40], and interferometry [41] have all been utilised for simultaneous characterisation of BLMs with electrical measurements. Formation of a planar lipid bilayer on a microchip device has also been developed for automated multichannel studies of membrane proteins. The bilayer membrane is formed at the etched aperture with a diameter of 50\u2013200 nm on a silicon substrate using either a microfluidic system or a liquid droplet injection system [42]. It consists of two fluidic channels on both sides of the silicon substrate and they are connected with apertures (Fig. 3). A lipid bilayer is formed at the aperture by flowing lipid solution and buffer alternatively. After the bilayer is formed, membrane proteins can be incorporated through for example proteoliposome fusion, and the activity can be monitored with the integrated electrode. In this microchip set-up, the thickness of the initial layer of phospholipids on the aperture is crucial for the bilayer formation. A high amount of lipid solution, which results in a thick layer of lipid on the aperture, should be avoided" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002428_01495730701738280-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002428_01495730701738280-Figure1-1.png", + "caption": "Figure 1 Geometry and coordinate system.", + "texts": [ + " On the basis of these results, the location and direction of initial crack growth are predicted numerically for the case of high carbon-chromium bearing steel (AISI52100). The initial crack growth is assumed to be determined by applying the maximum energy release rate criterion to randomly oriented cracks emanating from an inclusion. The effects of frictional coefficient, slide/roll ratio and depth of inclusion on the location and direction of the initial crack growth are considered. And stress intensity factors or energy release rate for the crack are investigated quantitatively. PROBLEM FORMULATION Figure 1 shows the geometry and coordinate system for this problem. The rolling contact is simulated as a contact pressure P1 x\u0303 and frictional load fP1 x\u0303 moving with constant velocity V over the surface of the half-space. Here, f is the frictional coefficient. A frictional heat input Q1 x\u0303 in the contact region is included to incorporate the thermal loading. The coordinates x\u0303 y\u0303 are fixed to the moving roller, the polar coordinates r\u0303 are used as a origin at a center of an inclusion, and the coordinates \u03031 \u03031 are fixed to the crack l\u03031 emanating from the inclusion. An rigid circular inclusion with radius a\u0303 is located at the depth d\u0303 and the distance e\u0303 from the center axis y\u0303 of the moving roller. In order to predict the start location and direction of the initial crack growth, the angle = 0 and the angle 1 between the crack direction \u03031 and r\u0303 axis are arbitrarily arranged respectively as shown in Figure 1. In the present analysis, the following dimensionless parameters are used. x = x\u0303/c y = y\u0303/c e = e\u0303/c d = d\u0303/c a = a\u0303/c r = r\u0303/c 1 = \u03031/c 1 = \u03031/c l1 = l\u03031/c Sr = Vs/V Pe = Vc/ t P x = P1 x\u0303 /P0 (1) H0 = 2 tG0 t 1+ /Kt 1\u2212 where, Sr is slide/roll ratio, Pe is Peclet number, Vs is sliding velocity under rolling, t is the thermal diffusivity, Kt is the thermal conductivity, t is the coefficient of thermal expansion, G0 is the shear modulus, is Poisson\u2019s ratio and P0 is the maximum Hertzian contact pressure" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001652_icar.2005.1507413-Figure3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001652_icar.2005.1507413-Figure3-1.png", + "caption": "Fig. 3. Transfigured mechanism by shifting the endplate and rotating the endplate", + "texts": [ + " 2020-7803-9177-2/05/$20.00/\u00a92005 IEEE We divide this mechanism into normal and transfigured types. When the endplate is horizontal at initial position and all the links are equal, we call it the normal mechanism (Figure 2). Link lengths of the normal mechanism remain fixed at all positions. The link manipulations are done using the actuators. The transfigured mechanism can be achieved in two different ways: One is offsetting the endplate by some distance and adjusting the links to hold this position (Fig.3 (a)). This transformation is called \u201dparallel transformation\u201d The other is adjusting the links upon rotating the endplate about one of the principle axes (Fig.3 (b)). This trasnformation is called \u201drotating transformation\u201d Thus, the links of the transfigured mechanisms are not equal and vary depending upon position or angle by which the end plate is transfigured. All these are adjusted passively. III. WORKSPACE ANALYSIS Workspace of parallel mechanism depends upon the mechanical parameters such as link size, size of endplate, size of base plate and stroke of the actuator. For our analysis we have set these parameters as shown in Table I and Fig.4 . Workspace can be divided into rotational and translational workspace" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000147_thc-2003-11202-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000147_thc-2003-11202-Figure1-1.png", + "caption": "Fig. 1. Geometric definitions for the assumed human arm structure, consisting of three segments.", + "texts": [ + " The upper extremity was modelled in terms of an inverse dynamics equation for a three segment planar manipulator [20,21]. The aim of the presented study is providing an alternative upper extremity clinical evaluation method which could be used on patients suffering from neuromuscular disorders usually following a stroke. Passive moment patterns obtained from such subjects are expected to show noticeable differences from the healthy ones. In this experimental work the human arm was described as a three degree of freedom kinematic and dynamic structure (Fig. 1). The segment lengths are denoted with a i, their centers of mass with li while qi indicate the positive angle directions with respect to the zero position (dashed line). Positive angle values are denoted with the arrow. The masses and inertias are presented with the m i and Ii variables. The centers of gravity were expressed as a distal distance from the joint marked with the same index. As in every other manipulator system, the dynamic behavior, as a relationship between applied driving torques \u03c4(u), environment forces h and joint motion trajectories q\u0308, q\u0307, q, of mechanical joints can be described as [20]: B(q)q\u0308 + C(q, q\u0307)(q\u0307) + G(q) + Fv q\u0307 + Feq + Fdsgn(q\u0307) = \u03c4(u) \u2212 JT (q)h (1) Here q, q\u0307 and q\u0308 represent the three component joint angle, angular velocity and angular acceleration vectors" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001979_50037-5-Figure7.2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001979_50037-5-Figure7.2-1.png", + "caption": "Figure 7.2 A quarter-symmetry model.", + "texts": [ + " Stresses resulting from shearing of the pin and bending of the pin will be neglected purposefully. The dimensions of the model are as follows: pin radius = 1 cm, pin length = 3 cm; arm width = 4 cm, arm length = 12 cm, arm thickness = 2 cm; and hole in the arm: radius = 0.99 cm, depth = 2 cm (through thickness hole). Both the elements are made of steel with Young\u2019s modulus = 2.1 \u00d7 109 N/m2, Poisson\u2019s ratio = 0.3 and are assumed to be elastic. In order to analyze the contact between the pin and the hole, a quarter-symmetry model is appropriate. It is shown in Figure 7.2. In order to create a model shown in Figure 7.2, two 3D primitives are used, namely block and cylinder. The model is constructed using graphical user interface (GUI) only. It is convenient for carrying out Boolean operations on volumes to have them numbered. This can be done by selecting from Utility Menu \u2192 PlotCtrls \u2192 Numbering and checking appropriate box to activate VOLU (volume numbers) option. From ANSYS Main Menu select Preprocessor \u2192 Modelling \u2192 Create \u2192 Volumes \u2192 Block \u2192 By Dimensions. In response, a frame shown in Figure 7.3 appears", + " In order to move the arm (vol. 6) in required position, coordinates shown in Figure 7.9 should be used. Clicking [A] OK button implements the move action. From Utility Menu select Plot \u2192 Replot to view the arm positioned in required location. Finally from Utility Menu select PlotCtrls \u2192 View Settings \u2192 Viewing Direction. The frame shown in Figure 7.10 appears. By selecting coordinates X, Y, and Z as shown in Figure 7.10 and activating [A] Plot \u2192 Replot command (Utility Menu), a quarter-symmetry model, as shown in Figure 7.2, is finally created. The next step in the analysis is to define the properties of the material used to make the pin and the arm. From ANSYS Main Menu select Preferences. The frame shown in Figure 7.11 is produced. From the Preferences list [A] Structural option was selected as shown in Figure 7.11. From ANSYS Main Menu select Preprocessor \u2192 Material Props \u2192 Material Models. Double click Structural \u2192 Linear \u2192 Elastic \u2192 Isotropic. The frame shown in Figure 7.12 appears. Enter [A] EX = 2.1 \u00d7 109 for Young\u2019s modulus and [B] PRXY = 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002482_s1068798x08070058-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002482_s1068798x08070058-Figure1-1.png", + "caption": "Fig. 1. Radial ball bearing (a) and the relation between the radial gaps \u2206ci and elastic decrease \u03b4ci in the separation in ball\u2013ring contacts under a load Pr (b) and in the contact of the central ball with the bearing rings under a load Po (c).", + "texts": [ + " The load distribution between balls in a motionless ball bearing, in the presence of an initial radial gap ( e \u2260 0), was established in [4]. In the present work, we adopt the methodology of [4]. On that basis, we analyze the load distribution between the balls in the bearing when the initial radial gap e between the balls and rings increases on account of wear of the internal and external bearing rings. The increase in the radial gap must change the load P i on the balls and, in certain conditions, change the number of contacting balls. The external radial load P r is applied to the internal ball-bearing ring (Fig. 1). The ring turns relative to the motionless external bearing ring at small angular velocity \u03c9 . As a result, the centrifugal force and gyroscopic torques on the bearing parts may be neglected. The distribution of the load P r between the balls is considered in a static formulation, with a symmetric ball distribution relative to the line of load application. One ball in the loaded zone is on the line of action of this force. Initially, the radial gap in the bearing is e . (Other ball configurations relative to P r were considered in [5]", + " (1) for P i , we obtain a relation between the load on ball i and the total distance \u03b4 c i traveled by this ball toward the bearing rings (2) where For the central ball (3) where \u03b4 co = \u03b4 1o + \u03b4 2o is the total distance traveled by the central ball at the points of contact with the ring; \u03b4 1o and Pi C\u03b4\u03b4ci 3/2,= C\u03b4 1.887/ n1\u03b4 \u03b72 \u03c11\u2211( )1/3 n2\u03b4 \u03b72 \u03c12\u2211( )1/3 +[ ] 3/2 .= Po C\u03b4\u03b4co 3/2,= RUSSIAN ENGINEERING RESEARCH Vol. 28 No. 7 2008 WEAR CALCULATIONS FOR RADIAL BALL BEARINGS 645 \u03b42o are the distances traveled by the central ball toward the internal and external rings. For ball i (Fig. 1b), \u03b4ci may be expressed in terms of the radial gap e and the total wear of the rings at their points of contact with the central ball (hco) and ball i (hci) (4) where hco = h1o + h2o; hci = h1i + h2i; h1o and h2o denote the wear of the internal and external bearing rings at their points of contact with the central ball; h1i and h2i denote the wear of the internal and external bearing rings at their points of contact with ball i; \u03b3 = 360\u00b0/z is the angle between two adjacent balls; z is the number of balls in the bearing", + " (10) may be written in the form (11) where The geometric model of the contacting bearing components must ensure equal reduced curvatures of the bodies at their contact points and take account of the specified channel curvature in the plane perpendicular to the plane of bearing rotation. Modification in shape of the channel surface changes the size of the contact J G p HB -------\u239d \u23a0 \u239b \u239e m ,= G k \u03be f( ) \u03b5 ----------\u239d \u23a0 \u239b \u239e m Rmax r\u03c81/\u03c7------------\u239d \u23a0 \u239b \u239e c .= area, the contact pressure, the position of the instantaneous axis of ball rotation relative to the channel, the slip at contact points, and ultimately the wear of the bodies. To satisfy these conditions, the actual ball 3 and channel 2 (Fig. 1c) are replaced by elastic bodies [7]: respectively, a roller 1 (Fig. 2) with variable radius of curvature Rx along the Ox axis; and a cylinder 2 with variable radius of curvature \u03c1. Here Rx = R2xR3x/(R2x + R3x), where R2x and R3x are the radii of curvature of the channel and ball in the plane of rotation parallel to, and at a distance x from, the plane passing through the Oz axis. The axis of cylinder 2 is perpendicular to the axis of roller 1; its generatrix is specified by the variation of the gap zx between ball 3 and channel 2 in the xOz plane (Fig. 1c), with no load. Each profile of the roller race will correspond to a different law z = \u03d5(x). Estimates of the variation in the radius Rx of body 1 within the contact area show that the current radius of curvature Rx of body 1 is close to its value Ro at the point x = 0, since the contact area is small relative to the ball radius [7]. This simplifies the geometric model and permits the following assumptions within the contact area: Rx \u2248 Ro and zx = z1x + z2x \u2248 z2x, where Ro is the reduced radius of curvature of cylinder 1 in the yOz plane; z1x, z2x are the gaps between the corresponding bodies and the x axis", + " 7 2008 WEAR CALCULATIONS FOR RADIAL BALL BEARINGS 647 area and takes the following form, according to the Boussinesq solution (14) where \u03b7 is the reduced elastic modulus of the contacting materials; pxy is the contact pressure at the point (x, y, 0), which is at a distance r from the point where the elastic strain is determined; ao and bo are the semiaxes of the contact-area ellipse. The gaps between the ball and various channel profiles were presented in [10]. For a channel of circular profile (R2 \u2248 1.03R3), the gap between bodies 1 and 2 at a distance x from the Oz axis in the xOz plane (Fig. 2) is (15) where R2ch and R3 are the radii of the channel and the wall (Fig. 1c). During a single contact with the ball at point E (Fig. 3), which is a distance x from the plane of rotation yOz, the channel wear is as follows, according to Eq. (11) (16) where Lx = is the frictional path at the selected point of the channel. The duration of a single ball\u2013channel contact at this point is t = 2bx/V, where bx is half the length of an element of contact area in cross section x; V is the velocity of the ball center O relative to the channel in the Oy direction. The contact surface of the ball moves relative to point E of the channel in the y direction, at constant slip velocity The frictional path at this point is where xc is the coordinate of the instantaneous axis of ball rotation relative to the channel; R3 is the ball radius", + " If some limiting permissible wear [h] is specified for the channel, the working life of the ball\u2013channel contact\u2014in the form of the number N of corresponding contact cycles\u2014is determined from the equation = [h] \u21d2 N. Taking account of the wear of the internal and external bearing rings, the working life N of the ball bearing with initial radial gap e may be found from the equation (24) where N is the number of rotations of the internal bearing ring before the total wear of the rings is equal to the limiting permissible wear [h], while the radial gap in the bearing reaches the limiting permissible value [e]; here the calculation points of the channels are A for the outer ring and B for the inner ring (Fig. 1): h1, h2 denote the total wear of the internal and external bearings in a single rotation of the inner ring characterized by j2 contacts with the balls under a load Po at point A of the outer ring, and by j1 contacts under loads Po, P1, P2 at point B of the inner ring, in accordance with Eqs. (8) and (9). The wear of bearing rings 1 and 2 during a single contact with ball 3 at contact points A and B when x = 0 (Fig. 1c) may be calculated approximately, disregarding the change in working-surface geometry of the balls, from the formulas (25) where p1o, p2o and L1o, L2o are the maximum Hertz contact pressures and frictional paths during a single contact of the balls with the external and internal bearing rings at points A and B, respectively, when x = 0 (Fig. 1c); the coefficient C(m) takes account of the elliptical pressure distribution along the frictional paths L1o and L2o; in the range 1 \u2264 m \u2264 6, we find that 0.75 \u2264 C(m) \u2264 0.45. In this case, the working life N of the ball bearing with an initial radial gap e is also determined by Eq. (24), where the calculation of the total wear h1 and h2 of the bearing rings takes account of Eqs. (8), (9), and (25) and the variation in the load Po, P1, P2 at the balls on account of wear of the bearing rings. \u2206h2oi 1= N\u2211 \u2206h1 \u2206h2+( ) i 1= N \u2211 h[ ] e[ ] e\u2013 N ,\u21d2= = \u2206 \u2206 \u2206h1o G p1o/HB( )mL1oC m( );= \u2206h2o G p2o/HB( )mL2oC m( ),= \u2206 \u2206 Consider the working life of a radial ball bearing of type 210 (State Standard GOST 8338-75) in terms of the limiting permissible wear" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003534_s0006297909040051-Figure4-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003534_s0006297909040051-Figure4-1.png", + "caption": "Fig. 4. Possible transfer mechanism of laccase forms in solution: NI, native intermediate; RE, resting form of the enzyme [32].", + "texts": [ + " Spectroscopic and crystallographic data on copper containing oxidases suggest a scheme of molecular dioxy gen reduction to water via a four electron mechanism [31]. According to this hypothetical scheme, completely reduced copper ions interact with the dioxygen molecule yielding \u201cperoxide intermediates\u201d, which then form a \u201cnative intermediate\u201d with all copper ions in the oxidized form. This \u201cnative intermediate\u201d can be further reduced and, in contrast, can transfer into the \u201cresting\u201d enzyme form as a result of protonation [32]. A possible mecha nism of this transfer obviously accompanied by change in the catalytic activity of laccase is presented in Fig. 4, but there is a lack of exact data in the literature. Studies of phase transitions of laccase allow evaluation of changes in the intermediate enzyme forms, including transfers relat ed with protonation of copper ions; this was one of the goals of the present work. To evaluate the dynamics of enzyme activity in the course of phase transitions, we studied the temperature effect of laccase activity from 4 to 20\u00b0C. Changes in activ ity of the native and frozen laccase preparations were evaluated during 4 h after cooling to 4\u00b0C and thawing, respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003435_12.841847-Figure5-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003435_12.841847-Figure5-1.png", + "caption": "Figure 5. Photograph of the measurement setup with the flow-through chamber as central element.", + "texts": [ + " The signal of the pyroelectric detector was analyzed with a lock-in amplifier (SR830, Stanford Research Systems, Inc., Sunnyvale, CA, USA) at a modulation frequency of 110 Hz. Proc. of SPIE Vol. 7560 75600E-4 Downloaded From: http://proceedings.spiedigitallibrary.org/ on 05/17/2015 Terms of Use: http://spiedl.org/terms The flow control and the read-out of the lock-in amplifier was performed by a PC with an instrumental control software (LabVIEW 8.6, National Instruments Corp., Austin, TX, USA). Figure 5 shows a photograph of the measurement setup consisting of the adjustable fiber coupler, the flow-through measurement chamber with the inserted fiber sensor and the pyroelectric detector at the end of the fiber. The design of the fiber sensor was accompanied by biotoxicity tests for the identification of other possible sensor materials. Therefore, in a so-called XTT test the material was eluated in a growing medium and mouse fibroblasts were brought in contact with the eluat. The number of viable cells could be evaluated due to the metabolic conversion of the added tetrazolium salt XTT" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001365_j.jmaa.2006.04.035-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001365_j.jmaa.2006.04.035-Figure1-1.png", + "caption": "Fig. 1. Geometry of a cylindrical journal bearing.", + "texts": [ + " From a mathematical point of view the problem is written in the following form which takes a compressibility operator B into account:\u23a7\u23aa\u23a8\u23aa\u23a9 Find u \u2208 L2(\u03a9) with u+ \u2208 H 1(\u03a9) such that \u2202 \u2202x ( h3 12\u03b7 \u2202u+ \u2202x ) + \u2202 \u2202y ( h3 12\u03b7 \u2202u+ \u2202y ) = \u2202 \u2202x ( hUB(u) 2 ) , (1.2) where B(.) is a known operator associated to a pressure compressibility relation: B(.) : R \u2192 R such that 0 < B \u2032(r) 1 \u03b2 , B(0) = 1, (1.3) for a positive constant \u03b2 (for which we will give a physical meaning in Section 4) and u+ = sup(u,0). Equation (1.2) has to be supplemented by some boundary conditions depending of the physical device considered. It is possible to use various boundary conditions, see [8,12,13]. In the following, we will consider the particular case of a journal bearing (Fig. 1) so that: \u03a9 := [0,1]2 and an input flow \u03b80(y) given on \u03930 = 0 \u00d7 [0,1] and the pressure is zero on \u2202\u03a9 \\ \u03930. The goal of this paper is to study existence and uniqueness properties for (1.2). The presence of the degenerate term u+ in the left-hand side prevents us to use classical techniques. The paper is organized as follows. In Section 2, existence of a weak solution of (1.2) is proved by using a generalization of Ky Fan\u2019s lemma. Condition of existence so obtained covered a wider range of parameters than the one issued from the Schauder fixed point approach in [8] and devoted to the particular case compressibility pressure law: B(u) = 1 + 1 \u03b2 u or the one issued from a 1- dimensional approach like in [1]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001967_978-1-4302-0088-8-Figure3-7-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001967_978-1-4302-0088-8-Figure3-7-1.png", + "caption": "Figure 3-7. The Lynxmotion Pan and Tilt Kit", + "texts": [ + "java: The extended version of TimedDiffDrive that gives speed control to any motion The next type of motion is still going to be done with servos, but this time it will move something on your robot rather than the robot itself. It will rely on the same principles discussed here but instead of creating methods like forward() or pivotRight(), it will create methods like lookUp() or lookRight() to move a camera. Sometimes you just want to move part of your robot. If you have a camera and are doing some things with machine vision (see Chapter 6), then you definitely want a pan and tilt camera system. The one I use is shown in Figure 3-7 and comprises a few brackets and two servos, which can be purchased from Lynxmotion for less than $35. I will take the same concepts used in our differential drive systems like grouping servos together in a class, and then we will use our MiniSSC class to control the servos. Figure 3-8 shows a diagram that summarizes the classes. 72 C H A P T E R 3 \u25a0 M O T I O N The objective in this section is to create a class to control a pan and tilt mechanism from a servo controller. The most important part of this class is the preconfigured constants\u2014for example, what pins connect the servos" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003122_978-90-481-3141-9_4-Figure3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003122_978-90-481-3141-9_4-Figure3-1.png", + "caption": "Fig. 3 ENF cylinder specimen", + "texts": [], + "surrounding_texts": [ + "Delamination specimens were machined from composite cylinders consisting of E-glass fibers in an epoxy resin. The internal diameter of the cylinders was 160 mm and the nominal wall thicknesses were 6 and 12 mm (12 and 24 layers). The lay-ups of the cylinders were \u0152\u02d9\u2122 6 and \u0152\u02d9\u2122 12, where \u2122 D 30\u0131; 55\u0131 and 85\u0131. A 58 mm long and 13 m thick, release agent coated aluminum film, was inserted at the mid-plane of the cylinders during the filament winding process to define a starter delamination crack, see Fig. 1. The film insert was wrapped around the circumference of the cylinder to enable machining of multiple test specimens from each cylinder. After filament winding, the cylinders were cured at 160\u0131 C for 3 h. The average fiber volume fraction was 0.61 for all cylinder lay-ups. Beam fracture specimens of a nominal length of 200 mm and a nominal width of 18 mm were cut from the cylinder wall for the subsequent delamination tests as schematically illustrated in Fig. 1. The beam axis was parallel to the cylinder axis producing straight beams with a curved cross-section. Figures 2\u20134 show the DCB, ENF and MMB cylinder specimens and loading principles. In order to accommodate the curved cross-section of the beam fracture specimens, contoured aluminum loading tabs were fitted to the DCB and MMB specimen and contoured loading pins and supports were attached to the ENF and MMB test fixtures as shown in Figs. 2\u20134. Further experimental details are provided in Refs. [10\u201312]. Glass/epoxy cylinders of lay-up \u0152\u02d955 n were manufactured for the external pressure tests using the filament winding process. Cylinder internal diameters were 55 and 175 mm, with wall thicknesses of 6.5 and 19 mm. The fiber volume fraction was 0.68. E-glass fibers in these cylinders were impregnated with the same epoxy resin as the cylinders used for delamination specimens. After winding, the cylinders were cured at 125\u0131 C for 7 h. In order to simulate fabrication defects in the form of delaminations that may arise during cure of thick cylinders due to exothermic heating, 50 mm square aluminum foil layers of 13 m thickness, coated with release agent on both sides, were introduced at different thickness locations during filament winding into some of the 55 mm diameter cylinders. The wall consisted of 12 layers and defects were placed between the third and fourth (referred to as 1/4 thickness), sixth and seventh (mid-thickness), and ninth and tenth layers (3/4 thickness) where ply #1 is the inner layer and ply #12 is the outer layer of the cylinder. Carbon fibre (T700) reinforced epoxy cylinders, (the same epoxy resin as for the glass reinforced cylinders), were filament wound with the same dimensions and cure cycle as for the glass/epoxy cylinders. Carbon (AS4)/PEEK (poly-ether-ether-ketone) cylinders were produced with the same dimensions by tape laying." + ] + }, + { + "image_filename": "designv11_69_0000523_iemdc.1997.604157-Figure3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000523_iemdc.1997.604157-Figure3-1.png", + "caption": "Fig. 3 shows the boundary conditions of the analysis model and the flux distribution when the displacement of moving part is 60[mm]. In the Fig. 3, the current angle is adjusted appropriately at the rated value in order to produce maximum thrust.", + "texts": [], + "surrounding_texts": [ + "I. INTRODUCTION\nPermanent magnet linear synchronous motors(PMLSMs) are proposed for many applications ranging from ground transportation to reciprocating servo system and conveyance system[l]-[3]. PMLSM's may be classified into the short primary (long PM poles) type and short secondary (short PM poles) type according to their structural features. Usually the short part acts as mover[2].\nIn PMLSM's, detent force is generated between the PM and the primary core. The ripple of detent force produces both vibration and noise, and deteriorates the control characteristics. The detent force characteristics are different to their structural features.\nThis paper analyzes the forces for those two types using finite element method (FEM). And, to reduce the detent force, the characteristics according to the magnet width are studied. The movement of moving part and the size of magnets are automatically varied in the FE analysis.\nFig. l(a) shows the basic structure of a short primary type PMLSM, and Fig. 1 (b) shows a short secondary type PMLSM. The parameters of the PMLSM are listed in Table I. A three phase winding is rolled, in distribution, on the primary core. In the analysis, the pole number of the short part is selected to four.\n(b) Short secondary type\nIn this paper, 2-D finite element method is used for the analysis of the magnetic field. The FE formulation is done for current source under the assumption that current is always appropriately controlled by using current-controlled inverter. The characteristic equation for 2-D FE analysis is given as (1).\nwhere, AI is the z 'component of the magnetic vector potential, Jo is the exciting current density and J,,, is the equivalent magnetizing current density of PM.\nSince the characteristics of NdFeB magnets are almost linear, with the recoil permeability very close to ,uo, the magnetizing current density of permanent magnet can be expressed as (2).\n(2) 1 J,,, = - V X M\nPO where, M is the intensity of magnetization.\nThe force is calculated using Maxwell's stress tensor. The integration surface of Maxwell's stress tensor is chosen to be at the center of air gap.\n0-7803-39464/97/SlO.O0 0 1997 IEEE. MC1-8.1", + "displacement\nwf : width ofPM ( STOP ] Fig. 2. Flowchart for automatic analysis o f PMLSM\nEven in the case of no exciting current in primary winding, force is generated between permanent magnets and teeth of core. This force is called detent force, and its magnitude varies according to relative position of permanent magnets and teeth of core. Therefore, it must be done to vary relative position of the primary and the secondary in order to analyze characteristics of PMLSM uprightly.\nFig. 2 shows the flowchart for a automatic analysis of a PMLSM. We consider the movement of a PMLSM with automatic moving mesh. Also, we make the width of PM change automatically in order to analyze the influence of the width of PM.\n111. ANALYSIS AND EXPERIMENTAL RESULTS\nA . Static Thrust and detent force\nFig. 4 shows the static thrust according to exciting current of a short secondary type PMLSM. In case of short primary type PMLSM, the static thrust is almost same as it. It is seen that the experimental results and the analysis results are in good agreement and that the static thrust is proportional to the exciting current.\nFig. 5 shows the analysis results of thrust under the assumption of sinusoidal control, and Fig. 6 shows the analysis results of the detent force according to mover's movement. The thrust of the short secondary type PMLSM pulsates with period of one slot pitch, and it can be known that the thrust pulsation is mainly due to the detent force generated between the PMs and the teeth of primary core. It can be known that the thrust of the short primary type LSM mainly pulsates with period of one pole pitch. Then, it can be seen that the thrust ripple of the short primary type PMLSM is mainly due to the influence of the primary's end section.\nFrom the comparison of Fig. 5 and Fig. 6, it can be seen that the aspect of thrust ripple is nearly identical with the detent force. Therefore, to reduce thrust ripple, detent force must be reduced.\nB. Normal force", + "-204 I 0 6 12 18 24 30 36\nDisplacement [mm] (a)\n2001 I\n- 2 0 0 4 , I , , r I , , J\nDisplacement [mm] (b)\n. Detent force according to displacement\nThe detent force can be reduced substantially by skewing either the primary teeth or the permanent magnets[5]. There are several disadvantages to skewing. It increases the complexity of motor construction, and reduces the thrust output. Another method for reduction of force ripple is to select magnet width appropriately[4].\nFig. 8 shows the variation of detent force ripple according to the width of magnets, and the detent force ripple is obtained from the difference between the maximum value and the minimum value of detent force. The detent force of the short secondary type PMLSM takes minimal value when the ratio of the magnet width to the slot pitch is (n+0.25) where n is an integer. But, in the short primary type PMLSM, the detent force ripple increases in proportion to the magnet width.\n(a)Short secondary type PMLSM (b) Short primary type PMLSM\nIV. CONCLUSIONS\nIn this paper, we analyzed and compared the thrust, detent force and normal force of a short primary type PMLSM and a short secondary type PMLSM.\nAs a result, for the short secondary type PMLSM, the detent force ripple can be reduced by changing the magnet width to the slot pitch. But, in the short primary type PMLSM, the detent force ripple is larger than that of short secondary type PMLSM.\nTo reduce the force ripple of short primary type PMLSM, further investigation should be required.\nREFERENCES\n[ I ] T. Mizuno, H. Yamada, \"Magnetic Circuit Analysis of a Linear Synchronous Motor with Permanent Magnets\". IEEE Trans. on Magnetics, Vol. 28, pp.3027-3029, 1992. [2] Guangyu Xiong, S . A. Nasar, \"Analysis of Fields and Forces in a Permanent Magnet Linear Synchronous Machine Based on the Concept of Magnetic Charge\", IEEE Trans. on Magnetics, Vol. 25, No. 3,\n[3] R. Akmese, J. F. Eastham, \"Design of Permanent Magnet Flat Linear Motors for Standstill Applications\", IEEE Trans. on Magnetics, Vol. 28,\n[4] Dal-Ho Im, Jung-Pyo Hong, Sang-Baeck Yoon, In- Soung Jung, \"The Optimum Design of Permanent Magnet Linear Synchronous Motor\", IEEE Proceeding on CEFC. pp. 166. 1996. [ 5 ] T. Li, G. Slemon, \"Reduction of Cogging Torque in Permanent Magnet Motors\", IEEE Trans. on Magnetics, Vol. 24, No. 6, pp.2901-2903, 1988.\npp.2713-2719, 1989.\nNO. 5, pp.3042-3044, 1992.\nMC1-8.3" + ] + }, + { + "image_filename": "designv11_69_0000171_icnn.1997.614157-Figure6-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000171_icnn.1997.614157-Figure6-1.png", + "caption": "Fig. 6 A usual jump by Kgroo;", + "texts": [ + " Either analytical or eqxrimental methcd can be used for the selection of the distinguishing measures. For instance, a simple neural net can learn that 0, has monotonic effect on D and A, with the data of a few similar jumps by a robot; or the theorems of parabolic lines can be used to determine the monotonicity as the trajectory of the robot mass center is parabolic for a jump. The 16 agents in Fig. 2 have all been s u m m y trained with three interactive layers (l3P neural net, associative memory, and fizzy rules). Fig. 5 shows an interpolation or extrapolation. Fig. 6 and 7 show the snapshots of two controlled usual jumps and two gymnastic jumps, respectively, by Kgroo and Rob00 (displayed by a graphical interface). ,.__ . .__ ' Hand D a,fe -3 landing .,' not propobonal ,:' back flip, vertical landing on the same spot ___.__......___________ ~ ... L-. Fig. 7 A gymnastic jump by Rdxx, 6. CONCLUSIONS A theory of multiagent neumfbzzy control has been analytically formulated. The new theory provides a natural explanation to both adaptive and explosive learning behaviors observed in locomotion learning by children and young animals" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001979_50037-5-Figure7.87-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001979_50037-5-Figure7.87-1.png", + "caption": "Figure 7.87 Contour plot of nodal solution (von Mises stress).", + "texts": [ + " Therefore, form ANSYS Main Menu select General Postproc \u2192 Read Results \u2192 By Load Step. The frame shown in Figure 7.85 is produced. The selection [A] Load step number = 1, shown in Figure 7.85, is implemented by clicking [B] OK button. From ANSYS Main Menu select General Postproc \u2192 Plot Results \u2192 Contour Plot \u2192 Nodal Solu. In the resulting frame (see Figure 7.86) the following selections are made: [A] Item to be contoured = Stress and [B] Item to be contoured = von Mises (SEQV). Pressing [C] OK implements selections. Contour plot of von Mises stress (nodal solution) is shown in Figure 7.87. Figure 7.87 shows von Mises stress contour for the whole assembly. If one is interested in observing contact pressure on the cylinder alone then a different presentation of solution results is required. A B C D Figure 7.88 Select Entities. From Utility Menu choose Select \u2192 Entities. The frame shown in Figure 7.88 appears. In the frame shown in Figure 7.88, the following selections are made: [A] Elements (first pull down menu); [B] By Elem Name (second pull down menu); and [C] Element Name = 174. The element with the number 174 was introduced automatically during the process of creation of contact pairs described earlier" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000188_ias.1997.626363-Figure3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000188_ias.1997.626363-Figure3-1.png", + "caption": "Fig 3. Architecture o f the RBFN", + "texts": [], + "surrounding_texts": [ + "In order to improve the iron-content controller performance the fixed statistical model used so far in the Smith-predictor controller to predict the Fe-content has been replaced by a. Radial Basis Function Network (RBFN) with on-line learning capabilities (see figure 2). The RBFN is based on multivariable interpolation theory. Its architecture is a feedforward multilayer type network. The Network architecture is composed of three layers, an input layer, one nonlinear hidden layer and a linear output layer. Each neuron of the hidden layer has a radial act.ivation function. The 21 39 Geswishen ~n D IPRESENTAIIEEE971GALMAlN7 DOC network output is simply the weighted sum (4,2) , and the weights 9, = 1, 9, = 1.5 , of the hidden units outputs (see figure A Radial Basis Function (RBF) is a function that maps the n-dimensional input domain J?\u201d onto the real axis. All input vectors which have the same distance to the center vector - c E IR\u201d yield the same output 4(x,c) = 4 (I(x - cll) = 4 ( r ) . ( 3 ) In equation ( 3 ) , 1 1 . 1 1 denotes the Euclidean norm. The input output relationship of the RBF model with the bias b is given by: (4) It has been shown previously [l] that RBF-networks are universal approximators. This means that any arbitrary nonlinear function will be described by a RBFnetwork arbitrarily precisely, if a sufficient number of RBFs is used. In order to be able to describe linear systems exactly without using an infinite number of RBFs, additional linear elements are included [21, i.e. equation . (4) is extended to A number of different basis functions is suggested in literature. In this paper the Gaussian radial basis function used: ~ . = e x p ( - l l ~ - g i l f / p ) I i = l , ..., m . (6) (5 =0.1, o2 =0.2, p = 1 and b=OS . The shape of surface y(x) is influenced by : the number and position of the centers, the spread of the RBFs (adjusted with and the magnitude of the weigths 9, 1 the factor p ) , undo,. For on-line identification applications using an RBF-network, some recursive rules are essential to update the centers and weights. The centers should suitably sample the network input domain and should be able to track the changing patterns of the data. The first phase of training the network is a clustering phase. In this phase, the incoming weights to the hidden layer learn to become the centers of clusters of input vectors. This clustering is done using a n-means clustering algorithm [ 3 ] . Compute the distances between the input vector and the current center vectors and find the minimum distance: .i(t)=l.(~)-.i(t-l>l i = l ... m Figure 4 shows as an example the output y of a RBF-Network with 2 inputs, x1 and x2, 2 hidden-units with the centers (2,4) und k=argb inQi ( t )> i = l ... m 1 ( 7 ) 2140 Geswlchert In D PRESENTAUEEE97iGALMAlN7 DOC Shift the center with the minimum distance in the direction of the current input vector: - c i (t) = C i (t - 1) i = I...m Ck (t) = Ck (t - I)+ a (Z(t)- C k (t - I)) ( 8 ) (18 The radii of the Gaussian functions at the cluster centers are fixed. The initial centers are chosen randomly. In order to provide a continual tracking capability, the time learning rate ct decreases to a constant learning rate a L . The second step of this hybrid algorithm is the algorithm for updating the RBFweights. Because of numerical problems the discrete square root filter in covariance form (DSFC) [ 4 ] was used instead of the conventional RLS-algorithm to minimize the cost function ( 9 ) with the error at time j With the input vector for the DSFCalgor i t hm 6 = [CD, ... CD, x, ... x,, I]' (11) - P = SS'' (12) - the covariance matrix _ - and the parameter vector - I 0 = [9 ..' 9,, 0 , ... 0,) D ] (13) one obtains: \"(k + 1)= y ( k + 1) - E , - \"(k + 1) &(k) - f ( k ) = s'@) E,@ + 1) @(k + 1) = @(k)+ y - ( k ) e(k + 1) ( 2 0 ) Using this hybrid algorithm a much faster convergence was achieved than by the back-propagation a:lgorithm [ 5 ] . The following additional actions are necessary before learning can start: 1.Select the input and output data: Due of the different positions of the sensing elements, every input has a different delay time to the output. Choose the appropriate values for one particular strip-segment. 2. Filter the measurement data 3. Freak value monitoring: Calculate on-line the variance and the mean value of every input und the output data and compare the current value with an upper and lower limit. If the current va:lue lies outside the permitted range, replace it by the mean value for a good prediction and stop learning." + ] + }, + { + "image_filename": "designv11_69_0003643_iciecs.2010.5678214-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003643_iciecs.2010.5678214-Figure1-1.png", + "caption": "Figure 1. Sketch of UAV configuration and control surface", + "texts": [ + " At the same time, as the natural frequency of the wing is very low, it will couple with the flight dynamics response and changing the dynamic response characteristics of UAV [1]. Undeformed rigid body model does not predict onset of flight dynamic instability seen in flexible model [2]. The order of high-fidelity flexible UAV model is up to hundreds [3], thus it is not suitable to be used in the conceptual design stage, in order to estimate the UAV scheme rapidly; the model must be simplified for the actual requirement. The solarpowered UAV discussed in this paper is shown in Fig. 1. There are two pairs of longitudinal control surfaces for the UAV: elevon (\u03b4ea) and elevator (\u03b4e). Since the stiffness of the fuselage and tail pole is much greater than the wing, so only consider the aeroelastic of the wing, the wing's torsional rigidity is also larger than the bending stiffness [4], therefore, in the stage of the conceptual analysis and control law design, it is enough to consider the wing bending only. The deformation characteristics of large aspect ratio UAV is large deformation and small strain [5], in the condition of large deformation, the change of moment of inertia will influent the flight dynamics greatly, so the change of moment of inertia while the deformation of wing must be considered [6]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001724_iros.1992.587376-Figure8-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001724_iros.1992.587376-Figure8-1.png", + "caption": "Fig. 8. Indirect transfer case for lemma 3.", + "texts": [ + " with radius pmin draw two circles CIhn and C2tangcritial to qi and qi+l respectively with centres 0\u2019 and O2 located on the same side where the angle formed by qi and qi+l is less than or equal to TC, as shown in Fig. 7b. The extension of qi+l is denoted as li+l . Lemma 2. If circle Clmin is on the same side of li+l as CZmin, and elrnin , direct transfer is possible. Draw an equilateral triangle with the line 0\u201902, connecting two centres of Clmin and CZhn , as the base of the triangle. One side passes through the point (xi+l, Y ~ + ~ ) , Fig.8. Lemma 3. If the circle elrnin is not completely on the Same side of li+i as C2min, and the length of the triangle is greater than or equal to 2pmin, an indirect transfer is possible. Otherwise, only reversal transfer is possible. Fig. 6. Traditional car parking problem It is notable that these lemmas should be used in the given order for deciding a transfer;proofs are not included in this paper. After the direct transfers or indirect transfers are determined corresponding path segments can then be constructed" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003761_15421400903112077-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003761_15421400903112077-Figure2-1.png", + "caption": "FIGURE 2 (a) Cross-section, (b) local layer structure, (c) fiber.", + "texts": [ + " We do not distinguish between the microscope fields n and p and the corresponding macroscopic ones representing their local averages. We denote n as the director and p as the polarization vector. Consider a cylindrical fiber with axis ex3, cross section R, and core X0 R x1x2 plane. The region X :\u00bc Rn X0 is an annular domain; that is R and X0 are open bounded simply connected sets with smooth boundaries C0:\u00bc @X0 and C1:\u00bc @R. Here X represents the cross-section of the smectic region, and n, p : X ! S 2 (See Figure 2). 2=[1136] P. Bauman and D. Phillips Let k be a unit layer normal at a given point x0 and u a unit vector perpendicular to k. Then the molecular orientation at x0 given through n(x0) and p(x0) is expressed in terms of three angles h, a, and / relative to k and u as depicted in Figure 3. Here 0< h< p is the tilt angle between n and k; jaj p is the angle obtained by rotating k n to p, about n using a right-hand rule; a measures the Stability of B7 Fibers 3/[1137] tilt of the molecular plane away from k n; / mod 2p is the azimuthal angle measuring the rotation of the molecule about k away from u toward k n", + " The free energy per unit length of the fiber, F , is determined by taking into account elastic energy (measured by layer strain and compression), polar divergence, electric self-interaction, and surface tension: F :\u00bc FSm \u00fe FF \u00fe FC \u00fe FP \u00fe FEl \u00fe FSu: The cross-section of the fiber\u2019s smectic portion is described by X and a function x (x1, x2) with x\u00bc 0 on the outer boundary and x\u00bc const. < 0 on the core boundary such that the cross-sections of the smectic layers foliating X correspond to level curves of x (see Fig. 2a). Variations in layer thickness from material dependent ground state values are energetically expensive. In contrast to this the layers themselves can undergo deformations with a (relatively) small cost in energy. A liquid crystal is in the smectic A phase if the director is parallel to the layer normal within the bulk and in the smectic C phase if the angle, h0, between the normal and n is such that 0< h0< p. The liquid crystals studied here are in the smectic C phase. The smectic C material has bulk layer thickness d0 and tilt angle h0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003170_icma.2009.5246362-Figure4-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003170_icma.2009.5246362-Figure4-1.png", + "caption": "Fig. 4 Elastic deformation", + "texts": [ + " I Reducer with one inner gear board angle is 180\u00b0 with each other; Before the dynam ic model is established, some assumptions are given as follows: (1) the inner gear is rigid; (2) the gap and damping of the rolling bearing is ignored, and the stiffness of the rolling bearing has a fixed value; (3) the stiffness and damping of the meshing gear pairs has a fixed value, which is computed according to the reference [6]. The inner planetary gear reducer with three axes is an over constraint mechanism, so the deformation equations should be set up. Fig. 4 is the sketch of the reducer, which is used to establish the elastic deformation vector equations in inputting shaft, supporting shaft 1, supporting shaft 2 and the meshing gear pair [7]. Index Terms - Inner planetary gear reducer with three axes, Mass unbalance response, Naturalfrequency The inner planetary gear reducer is a new kind of gear transmission. Its advantages are higher efficiency, greater ability to heavy load, smaller and simpler structure, larger transm ission ratio and so on. Meantime, it has some disadvantages such as acuter vibration and louder noise" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003592_1.4001258-Figure9-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003592_1.4001258-Figure9-1.png", + "caption": "Fig. 9 y+ contour of the exhaust gas boundary layer", + "texts": [ + "org/ on 0 help compare the respective parameters. Figure 8 shows the temperature contour of the exhaust pipe, where inlet 1 is furthest away from the outlet. Since the mass flow rate specified for each inlet is the total mass flow rate divided by six, the areas in red indicate that all the six exhaust inlets are open all the time. Since the stander wall function is used for modeling the turbulent boundary, it is needed to check the y+ value in the wall. The logarithmic law for mean velocity is known to be valid for 30 y+ 300. Figure 9 shows the y+ of the exhaust gas boundary is ranging from 30 to 191, which means the boundary layer mesh is appropriate for the simulation. Comparing the CFD and NTU theoretical calculation results in Table 3, the NTU method tends to overestimate the heat transfer rate due to the assumption that the exhaust gases from six inlets are not interactive with each other. Table 4 shows the difference between CFD and experimental results. The numerical simulation provides a good match with experimental data" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003197_j.ab.2009.03.034-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003197_j.ab.2009.03.034-Figure1-1.png", + "caption": "Fig. 1. Schematic of the electrochemical micro pH-stat built on the RSS platform. A 20-ll sample drop is surrounded and confined by a hydrophobic ring on a glass substrate containing an embedded Pt disc electrode and a junction. The dark arrows represent flow patterns generated in the drop by two tangential antiparallel air jets. A combination micro pH electrode is immersed to a depth of 1 mm from the top into the sample. The schematics are drawn approximately to scale.", + "texts": [ + " AChE, red cell cholinesterase \u2019s reagent), 5,50-dithio-bis-2hiocholine iodide; PBS, phos- Here we report on a novel approach for enzyme assays in microliter samples using electrochemical pH-stating. This approach takes advantage of the rotating sample system (RSS)1 platform that has been introduced in our laboratory as a convective platform for optical and electrochemical analyses in small sample drops [5\u20139]. The electrochemical micro pH-stat that we propose here incorporates the main concepts of this platform, as shown schematically in Fig. 1. A hydrophobic ring with proper inner diameter holds the 20-ll sample in place, confining it into a hemispherical shape by virtue of high surface tension at the air\u2013water interface. Vigorous rotation and mixing of the drop is generated by two antiparallel air jets tangential to the drop surface. Acid or base addition in our system is achieved by water electrolysis at a Pt mini-disc working electrode that is embedded in the stationary substrate. The magnitude of current is controlled by the actual difference between the real-time pH reading in the sample and the desired pH value set in the control system", + " All chemicals were obtained from Sigma Chemical (St. Louis, MO, USA) and Fisher Scientific (Pittsburgh, PA, USA). All aqueous solutions and subsequent dilutions were prepared with 18.2 MX-cm deionized water (Milli-QUV Plus, Millipore, Billerica, MA, USA). Substrate solutions were prepared freshly before the experiments. Serum samples stored in a refrigerator are stable for 7 days. Native and heat-inactivated fetal bovine serum (FBS) was obtained from HyClone Laboratories (Logan, UT, USA). The fabrication of the micro pH-stat cell (Fig. 1) was described in our previous work [10]. A 250-lm-diameter Pt mini-disc electrode (Alfa Aesar, Ward Hill, MA, USA) was employed as working electrode. Agarose (1% [w/w] type I) was prepared in heated 0.1 M KNO3 solution and filled into the junction hole embedded in the glass substrate. A hydrophobic ring of 4.2 mm inner diameter for sample positioning and confining it to a hemispherical shape (Fig. 1) was made of painted silicon elastomer (Dow Corning, Midland, MI, USA). A micro pH-electrode (MI 4154, Microelectrodes, Bedford, NH, USA) with a 1-mm-diameter tip was immersed at the vertical axis of the sampled drop to a depth of 1 mm for real-time pH monitoring. Determination of current efficiency in the electrochemical micro pHstat We measured current efficiency by adding 5 ll of HNO3 (2\u2013 8 mM) to lower the pH of 15 ll of FBS. A constant current was subsequently applied to neutralize the acid and bring the pH of the sample back to its original value" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003750_s10778-010-0361-x-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003750_s10778-010-0361-x-Figure1-1.png", + "caption": "Fig. 1", + "texts": [ + " This criterion is applied to rolling antivibration devices [2, 3]. In this connection, it is necessary to find the shape of the directrix of a cylindrical surface over which a heavy homogeneous cylinder rolls without slipping (pure rolling) in the least time. The requirement of no slipping is standard for vibration-protection problems. The present study continues the research reported in [4, 5] where the following differential equation of brachistochrone along which a cylinder rolls without slipping (Fig. 1) was for the first time derived by setting up and minimizing a functional T z x[ ( )]: ( )[ ( ) ] ( )z C z r z C 1 1 2 2 1 , z z x z L L ( ) , ( )0 0 . (1) 730 1063-7095/10/4606-0730 \u00a92010 Springer Science+Business Media, Inc. National University of Live and Environmental Sciences of Ukraine, 15 Geroev Oborony St., Kyiv, Ukraine 03041, e-mail: viktor.legeza@gmail.com. Translated from Prikladnaya Mekhanika, Vol. 46, No. 6, pp. 137\u2013143, June 2010. Original article submitted September 30, 2009. 1. Problem Formulation. Consider a homogeneous cylinder with mass m and radius r rolling without slipping over some cylindrical surface with directrix OKL from rest at a point O to a point L in a uniform gravity field (Fig. 1). The curve OKL lies in a vertical plane. The center of the cylinder is at a point M. We choose a rectangular Cartesian coordinate system OXZ (Fig. 1). Our task is to integrate Eq. (1) with boundary conditions and to identify the conditions under which the rolling of the cylinder is no longer pure. 2. Integrating the Brachistochrone Equation. The boundary conditions for Eq. (1) are defined by the coordinates of the points O( , )0 0 and L x z L L ( , ). An important fact is that if r 0, then Eq. (1) goes over into the well-known cycloid equation [1, 8], which is obvious because the cylinder degenerates into a material point. To integrate Eq. (1), we will use a special parametrization of the curve z z x ( ) different from that in [5]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002083_detc2007-34911-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002083_detc2007-34911-Figure2-1.png", + "caption": "Figure 2 In-plane flexible tire model.", + "texts": [ + " Furthermore, the modified elastic force description developed for a spatial curved beam element was recently proposed by the authors and the improvement of the convergence characteristics has been demonstrated [9]. The use of this formulation can allow for modeling the moving boundary resulting from the tire/road 2 Copyright \u00a9 2007 by ASME url=/data/conferences/idetc/cie2007/71022/ on 05/12/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use D interactions as well as for modeling fully nonlinear dynamic characteristics of tires using a relatively small number of degrees of freedom. As shown in Fig. 2, circumferential and radial springs and dampers are defined between the flexible belt and rigid rim in order to account for the sidewall stiffness of tires, while the tangential tire force is defined by introducing contact nodes attached to the curved belt elements. In the finite element absolute nodal coordinate formulation as shown in Fig. 3, the global position vector ier of an arbitrary point on element e of body i can be defined as ( )ie ie ie ie ie=r S x T p (1) where ( )ie ieS x is the element shape function matrix, [ ]ie ie ie Tx y=x is the vector of coordinates defined in the element coordinate system, Tie is a constant matrix that accounts for the slope discontinuities [10], and iep is the vector of nodal coordinates of element e on body i", + " 4, and ijK and ijC , respectively, define the stiffness and damping matrices as [ ], [ ]ij ij r t r tdiag k k diag c c= =K C (23) where the subscript r refers to stiffness and damping in the radial direction, while t refers to those in the circumferential direction. Using Eq. 20, the sidewall force ij mF given by Eq. 22 can be defined with respect to the global coordinate system by ij j ij m m m=F A F . Accordingly, the generalized forces associated with the sidewall flexibility can be defined for flexible belt body i and rigid rim body j as follows: , T T Ti ie ie ij j j ij m m m m m m= = \u2212Q T S F Q B F (24) where [ ( / ) ]j j j j m Pm \u03b8= \u2202 \u2202B I A u . Tire/Road Interaction As shown in Fig. 2, contact nodes are defined within the flexible belt element in order to account for tread/road interactions. To each contact node, the normal contact force is defined using an elastic contact approach, while the tangential force is defined using a creep function model. That is, the normal contact force at each contact node is defined by spring and damping forces which can be expressed by a function of penetration and its velocity between the contact node and road contact plane. The spring constant for the normal contact force can be determined using measured static load-displacement relationship" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002200_09544054jem699-Figure5-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002200_09544054jem699-Figure5-1.png", + "caption": "Fig. 5 Process of SSM", + "texts": [], + "surrounding_texts": [ + "The constraint condition of a TP, namely equation (8), specifies the lower limit of the information content of a TP, whereas the upper limit of the information content of a TP is constrained by the control condition, controller choice, and price cost, as well as the control efficiency to be discussed hereafter. To measure the control efficiency, a measure of the system function is needed. In this paper, the summation of the information content of the FR is used as the measure of system function according to IR \u00bc X j IRj \u00f013\u00de In the above equation, j indicates the number of FRs and IR is the measure of the system function, which is comparable between different design schemes. Based on IR, the control efficiency between different designs can be evaluated by defining the control efficiency h as h \u00bc IR M \u00f014\u00de h represents the control efficiency of the product control system. When M is decreased by reducing the number of TPs or TP information content, or when IR is increased by integrating or reconstructing structural components, h will increase. The control efficiency h is an important parameter to measure the design quality of a mechatronic system. By this standard, the second principle of the mechatronic design can be proposed: the maximumcontrol-efficiency principle states that an appropriate design for mechatronic system, which satisfies the independence axiom, the information axiom, and the constraint condition of the information content of a TP, is the design with maximum control efficiency. Since the comparison of the information contents of functions is constrained to be that of congeneric functions, the comparison of the control efficiencies of TPs is constrained to be that of congeneric systems, especially between different design schemes of the same system. The above restriction does not hinder the application of the maximum-controlefficiency principle, which is needed in system design to make a choice between many schemes. Whether a design scheme is good or bad can be confirmed only through comparison with other schemes of the congeneric systems. There is not an absolute criterion which can cover all system functions." + ] + }, + { + "image_filename": "designv11_69_0000347_isatp.1997.615388-Figure3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000347_isatp.1997.615388-Figure3-1.png", + "caption": "Figure 3: Stable object without force closure", + "texts": [ + " Force closure is defined as a state that an object can be given forces to any direction by a robot fingers. That is to say that the robot can generate any directional force and acceleration to the object. Therefore, it can resist any directional (small) disturbance applied on it. As an index of the object stability, we use this characteristics of force closure that the object can resist any directional disturbance. However, this stability condition can be achieved even if it is not in force closure. For example, the objects in Figure 3-(a) and (b) don\u2019t move even if any small disturbance is applied. They can resist to the disturbance while friction conditions are satisfied. They are also in stable. In Figure 3, the floor and the gravity force support the object. Floors or any other environment cannot apply any active force to drive the object, but it can support the object against disturbance in place of the fingers. The gravity force also support the object. In some situation, the gravity force may apply some disturbance to destroy stability, but in this example, it works as a support of the object. A difference between force closure and this stability is that the degree of the stability margin can be increased if the object is in force closure. It is controlled by forces from the fingers. In Figure 3, friction conditions provide the degree of the stability margin. Here, we point out the necessity of a new index of stability when we think about graspless manipulation. Graspless manipulation uses environment around the object to reduce operation forces. So, to evaluate graspless manipulation, we need some index how dexterous the environment is used. We define this stable state as \u201cGravity Closure.\u201d It means that an object is in enclosure with gravity force. Gravity force press an object against fingers and environment", + " Definition 1: Gravity closure i s a s tate which a n object can resist t o a n y directional manute applied disturbed force. T h e resis tant force i s generated b y f ingers , env ironment and gravity. Nakamura introduced \u201cwork closure\u201d as a state that fingers and environment can generate any directional force against disturbance[Nakamura90]. Our (\u2018gravity closure\u201d is similar to this closure. But in work closure, fingers and environment construct \u201cform closure\u201d and it doesn\u2019t include cases like as Figure 3-(b). Mason et al. also defined similar closure \u201cstatic closure\u201d [Mason93]. They defined \u201ckinematic closure,\u201d \u201cstatic closure\u201d and \u201cquasi-static closure\u201d to lead \u201cdynamic closure\u201d which was their main target. Aiyama et al. [Aiyama93] described graspless manipulation as \u201ca robot attempts to change the position and the posture of an object without grasping it.\u201d Here we define \u201cgraspless manipulation\u201d with the concept of \u201cgravity closure.\u201d Definition 2: Graspless manipulation is a m a - nipulat ion method that a n object contact wi th env i ronment around and i s in gravity closure while it i s manipulated" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002982_978-3-540-72699-9_46-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002982_978-3-540-72699-9_46-Figure2-1.png", + "caption": "Fig. 2. (E1, E2, E3) is a frame attached to the VTOL at its center of mass G. h is the vertical distance between center of mass and propeller center. d is the horizontal distance between center of mass and propeller axis. The transformation (e1, e2, e3) \u2192 (E1, E2, E3) is defined by (i) rotation of angle \u03b71 around axis 1, (ii) rotation of angle \u03b72 around axis 2, (iii) rotation of angle \u03c6 around axis 3.", + "texts": [], + "surrounding_texts": [ + "i.e. they are zero along the whole trajectory. Hence, the NE-controller reduces to:\n\u2206u(t) = \u2212K(t)\u2206x(t) (15) K = R\u22121FT\nu S (16) S\u0307 = \u2212Q\u2212 SFx \u2212 FT x S + SFuR \u22121FT\nu S (17) S(tk + T ) = P (18)\nwhich can be viewed as a time-varying LQR. Note that, if the local system dynamics are nearly constant, the NE-controller is well approximated by a LQR with a constant gain matrix K. In contrast, if the system is strongly time-varying, it is necessary to compute the time-varying NE-controller (15)-(18).\nThe simulated example is a VTOL structure. The structure is made of four propellers mounted on the four ends of an orthogonal cross. Each propeller is motorized independently. The propeller rotational velocities are opposed as follows (when top viewed, counted counterclockwise): propellers 1 and 3 rotate counterclockwise, while propellers 2 and", + "4 rotate clockwise. The angle of attack (AoA) of the blades and the positions of the propellers are fixed relative to the structure. The VTOL is controlled by means of the four motor torques. The states of the system are:\nX = [ x y z \u03b71 \u03b72 \u03c6 x\u0307 y\u0307 z\u0307 \u03b7\u03071 \u03b7\u03072 \u03c6\u0307 \u03c1\u03071 \u03c1\u03072 \u03c1\u03073 \u03c1\u03074 ] (19)\nVariables [ x y z ] give the position of the center of gravity G in [m] within the\nlaboratory referential (e1, e2, e3). Variables [ \u03b71 \u03b72 \u03c6 ] give the angular attitude of the structure in [rad], with the transformation from the laboratory referential to the VTOL referential, (e1, e2, e3) \u2192 (E1, E2, E3), being described by the matrix \u03a6(\u03b71, \u03b72, \u03c6) = Re3 (\u03c6)Re2 (\u03b72)Re1 (\u03b71), where Re(\u03b1) is a rotation of angle \u03b1 around the basis vector e.\nVariable \u03c1\u0307k is the speed of the propeller k in [rad/s]. The model of the VTOL can be computed by means of analytical mechanics. The aerodynamical forces and torques generated by the propellers are modeled using the standard squared velocity law. The resulting model is rather complicated and will not be explicited here. The reader is referred to [9] for details. The model is nonlinear, and its local dynamics are strongly time varying.\nA simplified model can be computed by removing certain non-linearities, which is well justified in practice. Introducing the notations\nv1 = 4\u2211\nk=1\n\u03c1\u03072 k v2 = 4\u2211 k=1 (\u22121)k\u03c1\u03072 k v3 = \u03c1\u03072 1 \u2212 \u03c1\u03072 3\nv4 = \u03c1\u03072 2 \u2212 \u03c1\u03072 4 v5 = 4\u2211\nk=1\n\u03c1\u0307k", + "the simplified model can be written as:\u23a1\u23a2\u23a3 x\u0308 y\u0308\nz\u0308\n\u23a4\u23a5\u23a6 = Cxyz \u23a1\u23a2\u23a3 sin(\u03b72) \u2212cos(\u03b72)sin(\u03b71) cos(\u03b72)cos(\u03b71) \u23a4\u23a5\u23a6 v1 \u2212 \u23a1\u23a2\u23a3 0 0 g \u23a4\u23a5\u23a6 (20)\n\u03c6\u0308 = C\u03c6v2 (21)\n\u03b7\u03081 = 1\nC1 \u03b71 cos(\u03b72)2 + C2 \u03b71\n(Cdsin(\u03b72)v2 + Cs d cos(\u03b72)(sin(\u03c6)v3 + cos(\u03c6)v4)\n\u2212 IM A cos(\u03b72)\u03b7\u03072v5) (22) \u03b7\u03082 = C\u03b72(\u2212cos(\u03c6)v3 + sin(\u03c6)v4 + IA Mcos(\u03b72)\u03b7\u03071v5) (23) v\u0307k = uk k = 1, ..., 4 (24)\nwith the constants\nCxyz = Cs\nM5 + 4m C\u03c6 =\nCd\n4IA M + 4md2 + I5 M\nC1 \u03b71 = \u2212I5 M + I5 S \u2212 4IA M + 4IA S \u2212 2md2 + 4mh2 C2 \u03b71 = IM 5 + 4IA M + 4md2 C\u03b72 = Csd\n4IA S + 2md2 + 4mh2 + I5 S\nThe inputs uk, k = 1...4, do not represent the physical inputs (the motor torques Mk), but are related to them by invertible algebraic relationships. The numerical values of the parameters used in the simulations are given in Table 1. Parameters I5\nM , I5 S , IA M , IA S are the inertias of the main body and the propellers,\nrespectively. Parameters Cs and Cd are the aerodynamical parameters of the propellers. Parameters M5, m are the masses of the main body and propellers, respectively.\nThe simplified model is flat. The flat outputs are: Y = [ x y z \u03c6 ] .\nTable 1. Model parameters\nCs 3.64 \u00d7 10\u22126 Ns2 d 0.3 m Cd 1.26 \u00d7 10\u22126 Nms2 I5 M 181 \u00d7 10\u22124 Nms2 M5 0.5 kg I5 S 96 \u00d7 10\u22124 Nms2 m 2.5 \u00d7 10\u22122 kg IA M 6.26 \u00d7 10\u22126 Nms2 h 0.03 m IA S 1.25 \u00d7 10\u22126 Nms2\nThe control problem is of the tracking type: the VTOL structure must be driven smoothly from some initial configuration to another predefined configuration. A" + ] + }, + { + "image_filename": "designv11_69_0000064_mma.1670030120-Figure3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000064_mma.1670030120-Figure3-1.png", + "caption": "Fig. 3 Regions of the x - r plane Fig. 4 Triangular initial shape", + "texts": [ + " If a\u201d(x) > 12 the string moves at least during one pseudo-period and 4 Case of the pinched string We assume the string initially at rest on the edges OA and AL of a triangle OAL: Fig. 4. The initial conditions are ( a u / a t ) (x,O) = 0 and u(x, 0) = ax for0 Q x g a u(x, 0) = p(1 - x ) for a g x < 1, (37) (38) (I = - and ap>O. The determination of the solution u(x, t ) of equations and inequality (2) satisfying those conditions, can be made by degrees in the following region (R,) of the x - t plane (Fig. 3). with a + B Region 6n + 1 2n - (I g x + t Q 2 n + a x 2 0 -2n - a g x - t Q - 2 n + a Region6n + 2 2n + a g x + t Q 2 ( n + l ) - a x g l -2n + a Q x - i g -2(n - 1) - a Region6n + 3 2n + a Q x + t g 2 ( n + l ) - a -2n - a Q x - t g -2n + a Region6n + 4 2n + a Q x + t g 2 ( n + l ) - a x 2 0 -2(n + 1) + a < x - t Q -2n - a Region6n + 5 2(n + 1) + a x g 1 -2n - a g x - t g -2n + a Region 6n + 6 2(n + 1) + a -2n - a g x - t Q -2(n + 1) + a The function u(x, t ) being continuous on the characteristics which u1 (x, t ) = ax, 2(n + 1) - a Q x + t Q 2(n + 1) - a < x + t ,< separate the'former regions can be determined by steps" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002628_cdc.2007.4434082-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002628_cdc.2007.4434082-Figure1-1.png", + "caption": "Fig. 1. The iRobot Packbot, used by e.g. US Armed Forces in Afghanistan.", + "texts": [ + " The resulting control mode is very similar to how characters are controlled in FPS games, such as Doom, Quake, Unreal, Half-Life, Counter-Strike and Halo, [8]. When the translational movements are controlled relative to the point of gaze, navigation in confined spaces and camera movements needed to improve depth perception, are much easier to 1-4244-1498-9/07/$25.00 \u00a92007 IEEE. 5794 perform. This is one step towards the better coordination between UGV motion and orientating perceptual functions, that is described as necessary for efficient control in [9]. With differential drive UGVs, we mean e.g. the UGVs depicted in Figure 1-3, with either tracks as in Figure 1, a castor wheel as in Figure 2, or skid to turn wheels as in Figure 3. Throughout this paper, we furthermore assume that the UGV has a pan camera, ideally one that is free to turn 360 degrees. The proposed control mode builds on the combination of two ideas. Firstly, the idea of inverse kinematics for robot arms, see [10], that enables the user to directly control the absolute position and orientation of the end effector, while disregarding the actual joint angles and motor commands of the arm. Secondly, the idea from [11], where Lawton et al", + " Finally, Section V discuss future investigations and experiments using this approach and conclusions are drawn in Section VI. In this section we will first review two ordinary differential equation (ODE) models capturing the kinematic and dynamic behavior of the robots depicted in Figures 1, 2 and 3 above. We then describe what a standard teleoperation interface looks like, and contrast this to the computer game interfaces used in FPS-games, [8]. We present kinematic and dynamic models for both cases. Consider the general UGV model in Figure 4. If we identify the two tracks of the Packbot in Figure 1, with the main wheels of the Scout in Figure 2, and the two pairs of wheels on each side of the ATRV in Figure 3, we find that the following model is applicable to all three robots. z\u03071 = v1 + v2 2 cos \u03b8 (1) z\u03072 = v1 + v2 2 sin \u03b8 \u03b8\u0307 = v1 \u2212 v2 d \u03c6\u0307 = k, where z = (z1, z2) and \u03b8 are the position and orientation of the vehicle, \u03c6 is the orientation if the camera relative to the vehicle, v1, v2 are velocities of the wheels/tracks and d is the width of the vehicle. Finally, k is angular velocity of the camera, relative to the vehicle" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003931_j.physd.2009.01.005-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003931_j.physd.2009.01.005-Figure1-1.png", + "caption": "Fig. 1. Each motor domain has two structural states corresponding to the nucleotide state, weakly bound (light) and strongly bound (dark). The structural state is described by the equilibrium position of the head with respect to the neck (\u03c8). The cargo linker and neck linkers are approximated by elastic elements.", + "texts": [ + " Brownian motion causes the head to move the remaining distance to the next binding site, working against the (time- and space-varying) internal stresses in the molecule. Once the diffusing head reaches the binding site, the head is assumed to have a high affinity to the microtubule. The majority of the time is spent in this mechanochemical state, with both heads having high affinity to the microtubule, but in opposite structural states. After a dwell time governed by the chemical kinetics, the structural states of both heads switch and the cycle repeats. The structure of the kinesin-1 molecule is approximated as shown in Fig. 1. The neck linker domains of the protein are approximated by elastic elements that connect themotor domains (heads) to the neck. The equilibrium position of these elastic elements is determined by the chemical state of themotor domain. Another elastic element connects the neck to the cargo (bead), approximating the cargo linker domain of the protein. Due to the small length scales of the problem, inertial effects are assumed negligible compared to the viscous forces which, hence, dominate the dynamics [2]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000779_j.crvi.2004.05.006-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000779_j.crvi.2004.05.006-Figure1-1.png", + "caption": "Fig. 1. Construction of a crystal (bold line), from the shape of the surface tension in polar coordinates. The contour of the crystal is the podal of the surface tension, i.e., the envelope of the normal to the rays of length \u03b3 (\u03b8).", + "texts": [ + " In simple terms, when trying to rock an element of surface, it is found that it is energetically easier to do it when the element is oriented parallel to specific crystallographic planes; conversely, if a crystal plane is not oriented parallel to some directions, there is a thermodynamic driving force that tends to modify the ori- entation. The actual shape of the crystal is obtained by the Wulff construction [5], which consists in drawing the envelope of the normal to the rays of length OM = \u03b3 (\u03b8) (Fig. 1). This curve, called the pedal, solves exactly Eq. (2) [5]. A similar construction exists in 3D [6,15]. One would dearly desire an equivalent construction for a \u2018crystal\u2019 made with biological tissue, in order to model, for example, the growth of a finger, a plant meristem, a radish, an avocado stone or an almond, or to use it as ingredient for a model of growth of kidney ducts or lung airways. In the equilibrium condition of a crystal surface, the \u22022\u03b3 (\u03b8)/\u22022\u03b8 term comes from the torque exerted on a surface which would not be aligned with the crystallographic axes" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003245_physreve.81.061704-Figure3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003245_physreve.81.061704-Figure3-1.png", + "caption": "FIG. 3. Tilted molecules in a Sm-C layer, with the longitudinal dipoles in equal number in opposite directions while the hindered rotation leaves an average transverse dipole.", + "texts": [ + " The uniaxial orientational order parameter OOP is expressed as Sij =ninj \u2212 1 3 ij where n is the director. When it is written in a frame for which the normal to the smectic layers is taken as the z direction it reads 2 Sij = \u2212 1/3 0 0 0 \u2212 1/3 0 0 0 + 2/3 , ij = aSij . 1 B. Sm-C If the preferred layer thickness decreases with temperature and becomes smaller than the length of the molecules, they have to tilt in one direction giving in the simplest case the phase predicted by Meyer 1 , the smectic C Sm-C where all the molecules are parallel Fig. 3 . The transverse dipoles give birth to the macroscopic polarization PS when summed up over at least ten layers. The longitudinal ones have to average to zero but they still sum up in a macroscopic quadrupole which main axis is tilted with respect to the layer normal. If one approximates the OOP Qij of the Sm-C to be the same Sij as in the Sm-A with its eigenvector 3 tilted at an angle with respect to the layer normal in the azimuthal direction 0, as detailed further in Appendix B for the biaxial case, one gets *marcerou@crpp-bordeaux" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003410_2013.26820-Figure9-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003410_2013.26820-Figure9-1.png", + "caption": "Figure 9. PTO driveline model.", + "texts": [ + " The density and shear mo\u2010 dulus of the component material used for the calculation were assumed to be 7,850 kg/m3 and 83 GPa, respectively. The tor\u2010 sional damping coefficient was calculated using the pre\u2010 viously obtained mass moment of inertia and torsional stiffness to meet the assumed damping ratio of 0.008 for steel (Neville, 1965). Table 2 shows the physical properties of the PTO driveline components (fig. 1). The PTO driveline model was constructed by combining its component models as shown in figure 9. It comprised of 24 parameters, which constitute a non\u2010linear 10\u2010degree of freedom model. The non\u2010linear parameters were backlash, mesh stiffness, and impact force. The backlash was the same for all pairs of gears. The mesh stiffness was assumed to be a harmonic function, since it changes with the relative space between the meshed gears. The impact force between the driving and driven gears was assumed to be described by non\u2010 linear equation 14 (Wang, 1997; Padmanabhan et al., 1995). The PTO driveline model was used to simulate the relative space and angular velocity between the driving and driven gears in the PTO gearbox in response to the torque fluctua\u2010 tions of the engine under no PTO load" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002609_icma.2008.4798870-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002609_icma.2008.4798870-Figure1-1.png", + "caption": "Fig. 1. Measurement system records an image of one block of four points on the considered calibration object.", + "texts": [ + " Note that an extended forward kinematic model incorporating both geometric and elastic effects is used for the parameter identification. We introduced this model in [6]. The problem formulation has a great impact on the results of the calibration and depends on the experimental setup. Existing approaches make use of an external measurement system in order to determine the robot base. A highly accurate determination of the robot base, however, is not feasible. 1) Experimental Setup: We propose an experimental setup consisting of an industrial robot and several calibration objects (Fig. 1). The only measurement system is a camera mounted on the robot flange. The robot executes various motions based on different joint configurations. In each joint configuration, the camera records an image of a block of points. Under the circumstance that at least 3 points not lying on a common line are measured within the same image, the orientation of the image frame w.r.t. the camera frame can be calculated. In some cases, this information is indispensable. Therefore, it is crucial that the camera measurements are reliable and accurate" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000955_cca.2003.1223144-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000955_cca.2003.1223144-Figure2-1.png", + "caption": "Fig. 2. The perspective camera model", + "texts": [ + " has been done in order to ensure the model similarity with robot parameters and affine camera model. The paper is organized in the following manner. In Section [I, the. camera model and robot ' model are developed. The calibration process for camera Calibration, hand-eye calibration and pose measurement aie derived in Section 111. Simulation is presented in Section lV,.followed by discussion in Section V. . . _ . . i 11. dAMER.4 AND ROBOT MODEI . . i. ~ A. CamemMOdel : ' One important camera model is thepinhole or perspective camera, which is used in this paper as shown in Fig.2. The. ;-perspectivecamera model consists o f a point 0, called center of p;_ojection, and the im,e plane..The distance between-0 and image-plane is the focal length f: The line perpendicular io image plane that goes through 0 is the optical axis T . For convenient, a point in . image plane and in world coordinate frame will be represented in small letters and capital letters respectively. A 2D point in image plane is denoted by m = [ , ~ , y r , and a 3D point is denoted by hi = [x,r,z]l. And the augmented vectors are = [" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001652_icar.2005.1507413-Figure4-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001652_icar.2005.1507413-Figure4-1.png", + "caption": "Fig. 4. Configuration of parallel mechanism with adjustable link parameters", + "texts": [ + " This transformation is called \u201dparallel transformation\u201d The other is adjusting the links upon rotating the endplate about one of the principle axes (Fig.3 (b)). This trasnformation is called \u201drotating transformation\u201d Thus, the links of the transfigured mechanisms are not equal and vary depending upon position or angle by which the end plate is transfigured. All these are adjusted passively. III. WORKSPACE ANALYSIS Workspace of parallel mechanism depends upon the mechanical parameters such as link size, size of endplate, size of base plate and stroke of the actuator. For our analysis we have set these parameters as shown in Table I and Fig.4 . Workspace can be divided into rotational and translational workspace. Since there are two types of adjustments possible for the mechanism, combinations of two types of workspace with two types of adjustments give four possible analyses of workspaces. To compare the workspace of normal mechanism with the transfigured mechanisms, we define the ratio of workspace volume U as follows: U = Vt Vn (1) Where Vt is the workspace volume of transfigured mechanism and Vn is the workspace volume of normal mechanism" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003212_educon.2010.5492353-Figure3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003212_educon.2010.5492353-Figure3-1.png", + "caption": "Figure 3. Different examples of detailed designs.", + "texts": [ + " In this way, materials are chosen according to the initial estimations of resistance needed for the different components and parts. Figure 2 shows as example the conceptual design. C) Detailed design Once the most appropriate solution has been chosen from the different pre-designs, the different parts must be exactly defined. Following the concepts explained in the theory classes, the students must use a design approach oriented towards manufacture and assembly, in line with the current trends in Concurrent Engineering. The results of some different tasks are shown in Figure 3. To check that the chosen materials are suitable, the estimates need to be compared, using simplified theoretical models, with the information provided by Computer Aided Engineering programs. The use of thermoplastic material injection simulation programs, which are habitually used prior to the construction of the moulds, is also important, in order to check that the choice of materials and injection conditions are appropriate, as well as an optimum distribution of the cavities in the mould, material inlets, filling channels and cooling system" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003617_ijtc2009-15255-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003617_ijtc2009-15255-Figure1-1.png", + "caption": "Figure 1 Experimental setup for studying effect of current passing through bearing loaded by Axial load (weight), Waveform of current passing through the bearing", + "texts": [ + " The author presents evidence of lubricant degradation giving rise to acidic radicals, which attack the metallic bearing elements. This leads to a change in lubricant chemistry as well as changes on the metal surface topography. Tests have been carried out for over 2000h to produce fluting. Har Parshad et. al have also worked out analytical formulae for fluting pattern. In this work 7204 angular contact bearings are tested with/without different cage materials (no cage, brass & plastic cage) at 50 N axial load, no radial load, rotating at 2700 rpm, lube oil and grease. Figure1 shows the construction of the fixture to hold and load the bearing. Note that the brush contact is insulated from the drive flange so that the bearing sees the full voltage. Current limit is set and voltage varies with bearing impedance. Four experiments were done at different level of voltage current, cage and bearing material (Table1) \u2013Current leads were fixed to the housing leading to ground and brush was loaded against the rotating shaft to apply a current \u2013 the only path for current was through the bearing. (Figure1) The current source was set at 0.1amp DC; the level of current was determined so that the voltage limit is not exceeded. The voltage was allowed to vary depending on the bearing impedance. Voltage values were noted down periodically and plotted as a function of time. The aim of these tests was to identify a voltage above which bearing currents are seen. A high value of this voltage would enable application of high dV/dt. During the screening tests, however, it was observed that even with DC voltage, there were bearing currents, which caused surface pitting characteristic of EDM damage" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002370_msec2007-31089-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002370_msec2007-31089-Figure1-1.png", + "caption": "Fig. 1. Schematic of a typical laser deposition system", + "texts": [ + " The simulation is applied to Ti-6Al-4V and simulation results are compared with experimental results. Laser deposition is an extension of the laser cladding 1 tp://proceedings.asmedigitalcollection.asme.org/ on 01/27/2016 Te process. This additive manufacturing technique allows quick fabrication of fully-dense metallic components directly from Computer Aided Design (CAD) solid models. The applications of laser deposition include rapid prototyping, rapid tooling and part refurbishment. As shown in Fig. 1, laser deposition uses a focused laser beam as a heat source to create a melt pool on an underlying substrate. Powder material is then injected into the melt pool through nozzles. The incoming powder is metallurgically bonded with the substrate upon solidification. The part is fabricated in a layer by layer manner in a shape that is dictated by the CAD solid model. During the laser deposition process, several defects, such as porosity and cracks, should be paid attention to. Cracks initiate corrosion fracture and reduce fatigue strength of the deposited parts" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003928_s10409-010-0364-1-Figure8-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003928_s10409-010-0364-1-Figure8-1.png", + "caption": "Fig. 8 All force components on the helical filament. Bending moment M1, M2; torque T ; shear force Q1, Q2; tension P1", + "texts": [ + " However, we still need to answer the following questions regarding this experiment: (1) In the viscous flow, the filament is subjected to several kinds of forces, such as torque, bending, tension and shear forces, which force dominates the transition process of the filament, such as nucleation and the propagation of the new phases? (2) After the nucleation of the new phase, what kinds of criteria make the interface to propagate in this experiment? In the following analysis, we will answer the above two questions. 2.3 Mechanical analysis of the cyclic transition of filament in the experiment of Hotani Figure 8 shows the force components at the filament cross-section, where M1 and M2, the bending moments, are, respectively, along and perpendicular to the direction of the curvature of the filament, T is the torque, Q1 and Q2 are the shearing forces, respectively, and P is the tension in the filament. The shear forces are neglected in the analysis. As to viscosity, the frictional constants in the normal and the tangent direction are different. We analyzed force conditions for the left handed normal form and the right handed semi-coiled form separately" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001236_s10778-005-0063-y-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001236_s10778-005-0063-y-Figure1-1.png", + "caption": "Fig. 1", + "texts": [ + " Such an approach allows us to simplify the mathematical model by eliminating imperfect constraints, which preserves the structure of the system and, hence, the order of the differential equations. However, this mathematical model is inapplicable to some real dynamic systems. For example, some design improvements in seismic isolation mechanisms of buildings [3, 10\u201312] may increase their rolling friction, which is important for effective damping of small vibrations. Such friction can be provided by filling base 1 (Fig. 1) and partially the spherical recess in body 2 with sand or by coating spheres 3 with rubber or plastic. Otherwise, it is necessary to use special dampers such as hydraulic shock absorbers. Figure 1 shows a design model of the system of interacting bodies described in [6]. It can be used, for example, in seismic isolation mechanisms of buildings [3, 4, 12]. Body 2 can only translate, since it bears with its spherical recesses upon identical spherical bodies 3. These balls roll over the flat surface of base 1 oscillating along the Ox-axis. Since balls 3 are identical in size and in position relative to body 2, it is enough to consider only one of them. The translational acceleration of base 1 may cause several modes of motion [6] of the variable-structure system [1] under consideration", + " Projecting the vector equation of plane translational motion of body 2 onto the x- and z-axes and introducing the relative coordinate u x x= \u22121 2 instead of x2 , we obtain the following two differential equations: / u F m x+ =1 2 1, /z N m g2 1 2\u2212 =\u2212 . (2) One of the coordinates (u or z2 ) is redundant and can be eliminated with the help of the holonomic constraint between bodies 1 and 2. If there are both frictional bonds 1\u20133 and 2\u20133, then it is convenient to represent the corresponding constraint equation in differential form, expressing the vector of vertical velocity of the contact point A2 (Fig. 1) in terms of translational and relative velocities: k i v z u A A2 2 1 = + , (3) where z2 is the projection of the velocity of body 2 onto the Oz-axis; vA A2 1 is the velocity vector of relative motion, i.e., rotation of the point A2 of body 2 about the point A1 of body 1; and k and i are the unit vectors of the z- and x-axes. Projecting (3) onto the axis perpendicular to vA A2 1 and considering that the acute angle between the vectors i u and vA A2 1 is equal to \u2212\u03b2 / 2 (Fig. 1), we obtain the holonomic-constraint equation tan /z u2 =\u2212 \u03b2 2. (4) System (2), (4) should be supplemented with the following three equilibrium equations for the inertialess ball 3 whose plane motion is defined as the projections of forces onto the Ox- and Oz-axes and the sum of moments of forces about, say, the center of a ball of radius r. These equations are valid for the states of rest, rolling, or sliding of the ball [6] N F F2 1 2 0sin cos\u03b2 \u03b2+ + = , N N F1 2 2 0\u2212 + =cos sin\u03b2 \u03b2 , N N F F1 1 3 2 2 3 1 2 0tan tan\u03b1 \u03b1k k+ \u2212 + = , (5) where the angles of rolling friction \u03b11 and \u03b12 are defined by \u03b1 \u00b5i i r=arctan ( / ), i =1 2, . When the ball rolls without sliding, in addition to the unknowns u z N N F, , , ,2 1 2 1, and F2, Eqs. (2), (4), and (5) implicitly include the angular coordinate \u03b2, which can be determined from an additional geometric-constraint equation. Considering the dependence cos /u A A= v 1 2 2\u03b2 = r\u03c9 \u03b23 1( cos )+ (Fig. 1), where v A A1 2 is the velocity of revolution of the point A1 of body 1 around the point A2 of body 2 and using the inverse motion formula (as that for planetary gears [2]) to eliminate \u03c9 3 , we obtain the second geometric-constraint equation in differential form [ ] / ( )( cos )\u03b2 \u03b2= \u2212 \u2212 +u R r 1 . (6) Integrating the system of two equations (4), (6), we obtain both holonomic-constraint equations in ordinary form, i.e., as relations between coordinates: u R r= \u2212 \u2212 +( )( sin )\u03b2 \u03b2 , z R r2 1= \u2212 \u2212( )( cos )\u03b2 , (7) which can also be derived in other ways. Thus, when ball 3 rolls without sliding over bodies 1 and 2, the motion of this plane system (Fig. 1) is described by the two differential equations (2), the three static-equilibrium equations (5), and the two geometric-constraint equations (4) and (6). Equations (4) and (6) may be replaced with Eqs. (7). To eliminate one of the coordinates u or z2 , it is necessary to differentiate both sides of Eq. (4) with respect to time: tan ( )( cos ) u z u R r \u03b2 \u03b22 1 2 2 2 + = \u2212 + . (8) Irrespective of mathematical description, the system has one degree of freedom and its linearized form has one natural frequency", + " Mode (iv) exists for | | | |\u03b2 \u03b2= 2m , \u03b2 u < 0; the motion in this mode is described by the first differential equation in (2), the first two equations in (5), and the relations , , ( )( cos )z z z R r m2 2 2 20 0 1= = = \u2212 \u2212 \u03b2 , | | | |\u03b2 \u03b2= 2m , F N f N m g2 2 2 1 2= \u2212 =sign \u03b2, . (35) Mode (v) is described by the same equations as those in the category B system. We will not consider here systems with f f f f1 0 2 0> <, , f f f f1 0 2 0< \u2265, , f f f1 2 0, < . As an example, let us examine a system (Fig. 1) with the following characteristics: \u03b1 \u03b11 2= = 0.05, m2 55 10= \u22c5 kg, 35 10= = \u22c5 \u2212 m, and g = 9.8 m\u22c5sec\u20132. The initial conditions: \u03b2( )0 0= , u( )0 0= , and ( )u 0 0= . In contrast to [4], the perturbation x1 of base 1 is specified in a convenient form: ( )[ ] ( )[ ]{ } exp sinx A t t t ti i i i i i i 1 2 0 2 0 1 7 = \u2212 \u2212 \u2212 = \u2211 \u03b5 \u03b5 \u03c9 , Ai [m\u22c5sec\u20132]: 0.28, 1, 4, 2, 3, 0.8, \u03b5 i : 1, 0.5, 1, 1.5, 1, 0.4, 1, t i0 [sec]: 1, 3, 5, 6.5, 8, 10, 10.5, \u03c9 i [sec\u20131]: 6, 8, 10, 4, 9, 7, 6, | | | |\u03b2 \u03b21 2m m= \u2032 = 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002979_978-3-540-70534-5_7-Figure14.3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002979_978-3-540-70534-5_7-Figure14.3-1.png", + "caption": "Figure 14.3: Basic manipulator joint types", + "texts": [ + " Each actuated joint is called a degree of free- 1 0 0 1 0 1 0 2 0 0 1 3 0 0 0 1 1 0 0 0 0 0 1 0 0 1 0 0 0 0 0 1 \u2212 Trans vx vy vz, ,( ) 1 0 0 vx 0 1 0 vy 0 0 1 vz 0 0 0 1 = Rot x \u03b8,( ) 1 0 0 0 0 \u03b8cos \u03b8sin 0 0 \u03b8sin \u03b8cos 0 0 0 0 1 = Rot y \u03b8,( ) \u03b8cos 0 \u03b8sin 0 0 1 0 0 \u03b8sin 0 \u03b8cos 0 0 0 0 1 = Rot z \u03b8,( ) \u03b8cos \u03b8sin 0 0 \u03b8sin \u03b8cos 0 0 0 0 1 0 0 0 0 1 = \u2212 \u2212 \u2212 Robot Manipulators 208 14 dom, and so a manipulator with six joints is a six-degree-of-freedom manipulator, or in short: 6-dof. We start by introducing the standard manipulator joints and their unambiguous drawing norm. There are three basic types or manipulator joints (see Fig. 14.3): rotational joints with the rotation axis along the link, rotational joints with the rotation axis perpendicular to the link (hinge joints), and prismatic joints (telescopic joints). All other joints, e.g., a more complex ball joint, can be described as a combination of these basic types. The Puma 560 from Unimation/St\u00e4ubli is a standard 6-dof manipulator that is frequently used as a model in textbooks. Figure 14.4 shows a simulation screenshot in RoboSim [Br\u00e4unl 1999] of this robot and its conceptual drawing, labeling its joints \u03b81, " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000428_6.1979-2033-Figure14-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000428_6.1979-2033-Figure14-1.png", + "caption": "FIGURE 14. OCEAN-GOING SURFACE-EFFECT SHIP CONCEIVED BY THE LlPPlSCH RESEARCH CORP. WHICH", + "texts": [], + "surrounding_texts": [ + "D ow\nnl oa\nde d\nby F\nre ie\nU ni\nve rs\nita et\nB er\nlin o\nn D\nec em\nbe r\n12 , 2\n01 6\n| h ttp\n:// ar\nc. ai\naa .o\nrg |\nD O\nI: 1\n0. 25\n14 /6\n.1 97\n9- 20\n33", + "FIGURE '2 BOEING CONCEPT OF ASW AIRCRAFT, ~ 1 ~ - 1 9 6 0 s ( ~ ~ )\nFIGURE 13. THE LONG-RANGE RAM I GE-STOL VEHICLE, PROPOSED IN ABOUT 1965(17) FIGURE 15. CYGNE 10, LARGE AIR FREIGHT CARRIER\nCONCEIVED BY J. BERTIN CO., FRANCE, 1969(15)\nD ow\nnl oa\nde d\nby F\nre ie\nU ni\nve rs\nita et\nB er\nlin o\nn D\nec em\nbe r\n12 , 2\n01 6\n| h ttp\n:// ar\nc. ai\naa .o\nrg |\nD O\nI: 1\n0. 25\n14 /6\n.1 97\n9- 20\n33", + "FIGURE 16. CYDNE 14. LARGE AIR FREIGHT CARRIER\nJ CONCEIVED BY J. BARTIN CO., FRANCE, 1973(15)\nFIGURE 17. PROPOSED LAYOUT OF WINGED SURFACE EFFECT VEHICLES (WSE'V), WATER\nRESEARCH COMPANY, 1974 115)\nFIGURE 20. XI14 AEROFOIL BOAT DEVELOPED IN 1977 BY RHEIN-FLUGZEUBAU GMBH ( R F B ) ( ~ ~ )\nD ow\nnl oa\nde d\nby F\nre ie\nU ni\nve rs\nita et\nB er\nlin o\nn D\nec em\nbe r\n12 , 2\n01 6\n| h ttp\n:// ar\nc. ai\naa .o\nrg |\nD O\nI: 1\n0. 25\n14 /6\n.1 97\n9- 20\n33" + ] + }, + { + "image_filename": "designv11_69_0000597_cbo9780511547126.019-Figure17.1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000597_cbo9780511547126.019-Figure17.1-1.png", + "caption": "Figure 17.1.2: Profiles of rack-cutter for Novikov gears with two zones of meshing.", + "texts": [ + " Novikov\u2013Wildhaber gears have been the subject of intensive research [Wildhaber, 1926; Novikov, 1956; Niemann, 1961; Winter & Jooman, 1961; Litvin, 1962; Wells & Shotter, 1962; Davidov, 1963; Chironis, 1967; Litvin & Tsay, 1985; Litvin, 1989; Litvin & Lu, 1995; Litvin et al., 2000c]. New designs of helical gear drives 475 Cambridge Books Online \u00a9 Cambridge University Press, 2009available at https://www.cambridge.org/core/t rms. https://doi.org/10.1017/CBO9780511547126.019 Downloaded from https://www.cambridge.org/core. University of Warwick, on 18 Sep 2018 at 01:19:11, subject to the Cambridge Core terms of use, P1: JXR CB672-17 CB672/Litvin CB672/Litvin-v2.cls February 27, 2004 0:58 Figure 17.1.1: Previous design of Novikov gears with one zone of meshing. Cambridge Books Online \u00a9 Cambridge University Press, 2009available at https://www.cambridge.org/core/t rms. https://doi.org/10.1017/CBO9780511547126.019 Downloaded from https://www.cambridge.org/core. University of Warwick, on 18 Sep 2018 at 01:19:11, subject to the Cambridge Core terms of use, P1: JXR CB672-17 CB672/Litvin CB672/Litvin-v2.cls February 27, 2004 0:58 Figure 17.1.3: 3D model of new version of Novikov\u2013Wildhaber gears. are now based on application of a double-crowned pinion tooth surface. Crowning in the profile direction enables localization of the bearing contact. Crowning in the longitudinal direction provides a predesigned parabolic function with a limited value of maximal transmission errors [Litvin et al., 2001c]. Profile crowning but not doublecrowning was applied in the initially proposed Novikov gears (Figs. 17.1.1 and 17.1.2). Therefore, the noise of such gears was inevitable. A new version of Novikov\u2013Wildhaber gears (Fig. 17.1.3) that is free of the disadvantages of the existing design is presented in this chapter. The chapter covers (i) various methods for generation of pinion and gear tooth surfaces of the new design, (ii) avoidance of undercutting, and (iii) stress analysis. The proposed new version of helical gears is based partially on the ideas that have been presented in the patent [Litvin et al., 2001c] and in the literature [Litvin et al., 2000c, 2002d] as follows: (1) Two mismatched parabolic rack-cutters are applied instead of rack-cutters with the circular arc profiles proposed for Novikov\u2013Wildhaber gears" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001701_amc.2006.1631675-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001701_amc.2006.1631675-Figure1-1.png", + "caption": "Figure 1. A Wheeled Mobile Manipulator", + "texts": [ + " Although it generally yields suboptimal results, the RPA is remarkably flexible and computationally competitive. It is applicable in various types of problems, including those with discontinuous dynamic models involving friction efforts and also for various types of cost functions and constraints. Solutions have been obtained in the case of fixed-base manipulators [27][28], wheeled mobile robots [29] and bipedal robots [30]. A deterministic version of this method, based on SQP optimization, has been used in the case of grasping moving objects [31] and closed-chain robots [32]. Consider the WMM shown in Figure 1. It consists of a mobile platform and a series-chain multi-link manipulator. Generally, the mobile platform may belong to any type of the nonholonomic wheeled mobile robots discussed in [33]. Therefore, the type of the mobile platform is not specified here. Let qp Rnp be the coordinates of the mobile platform, qa Rna be the coordinates of the manipulator and q = [qp T; qa T]T Rn be the vector of generalized coordinates of the whole system. The nonholonomic constraints subjected by the wheeled mobile platform can be written as [33]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002789_aim.2007.4412486-Figure4-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002789_aim.2007.4412486-Figure4-1.png", + "caption": "Fig. 4. Forces and torques acting on the axle and on the roll.", + "texts": [ + " Its value can be calculated as the product between the normal component of the contact force Ns and the adhesion coefficient \u03bc, whose value depends on several uncertain and difficultly quantifiable parameters (for example, it is sensitive to the environmental conditions). In the proposed model, the adhesion is related to the relative sliding \u03b4, [12], [11], [10], Fig. 3 shows an example of adhesion function. The maximum \u03bc0 and the asymptotical value \u03bcasy can be set by the user as function of the (simulated) traveled distance, in order to simulate different conditions during a test. When the locomotive is running on the rig, the dynamics of the axle and the roller is described by the following equation (Fig.4): JB\u03c9\u0307 = C \u2212 u r R , (4) where JB = [ J + JR ( r R )2 ] represents the total inertia of the axle/roller system, calculated with respect to the wheel rotation axis, JR is the roller momentum of inertia, \u03c9\u0307 is the actual wheel angular acceleration, u is the torque applied by the roll motor (control torque), C is the actual value of the locomotive axle torque, and R is the roll radius. This equation is obtained taking into account the condition that, 1-4244-1264-1/07/$25.00 \u00a92007 IEEE 3 if a pure rolling motion exists between the wheel and the roller, r\u03c9 = R\u03c9R" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000345_iros.2000.895270-Figure3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000345_iros.2000.895270-Figure3-1.png", + "caption": "Figure 3: Virtual sticks.", + "texts": [ + " The manipulators are controlled to make contact with the object safely considering the structural compliance of the manipulators in phase 3. In phase 4, the angular velocity of the object is slowed gradually to prevent slipping between the manipulators and the object. when the hands grasp the object. At the beginning of phase 5, the object is motionless. The object is then translated and rotated to the desired position and orientation. 3 Kinematics and Statics We consider constant vectors called virtual sticks shown in Figure 3. The kinematics based on the virtual sticks is originally proposed for two arms holding an object in [lo]. In this paper, we use the kinematics approximately while the hands are near the object shown in Figure 3. The radius of the cylindrical object is measured by laser displacement sensor or vision system, and the length of the virtual sticks is equal to the radius. The zero point of C, is equal to the middle point between the tips of one virtual stick and the other stick. The orientation of C, is synchronized with the motion of the two hands. CO represents the base coordinate system. We use the task vectors based on the virtual sticks. The task vectors of generalized forces, generalized velocities and generalized positions are defined as [lo]: 7 7 = [ Of: -2037- where Of,, Os, and 'pa, respectively, represent external forces, absolute velocities and absolute positions of the tip of the virtual sticks relative to CO, \"f,, \"As, and \" Ap, , respectively, represent internal forces, relative velocities and relative positions of the tip of the virtual sticks relative to C a " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002478_gt2007-28219-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002478_gt2007-28219-Figure1-1.png", + "caption": "Figure 1. Schematic drawing of the hydrostatic bearing system.", + "texts": [ + "org/ on 01/31/2016 T U runner (worktable or rotor) velocity u velocity component in x direction u non-dimensional velocity component, Uuu v velocity component in y direction v non-dimensional velocity component, Uuv x , y Cartesian coordinates, hyxyx ,, x , y non-dimensional Cartesian coordinates, hyxyx ,, iw weighted function land-width ratio, La c capillary restriction parameter, 3 c 4 cc h32d3 o orifice restriction parameter, 3 s 2 odo hPdC 23 jet-strength coefficient, 2 s UP absolute viscosity of lubricant density of lubricant The present study is concerned with the flow characteristics within three rectangular recesses as illustrated in Fig. 1 that would be used as hydrostatic pockets in a hydrostatic bearing or slidway system. The lubricant is supplied from a constant pressure source flowing into each recess through a restrictor into the recess and sills. A buffer pocket is installed between the restrictor and recess in order to reduce the effect of jet impingement on the runner (which is rotor or worktable). The runner runs with a constant velocity. Hence, the bearing generates hydrostatic and hydrodynamic effects through lubricant film" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000615_acc.2004.1383676-Figure10-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000615_acc.2004.1383676-Figure10-1.png", + "caption": "Fig. 10. Proof-of-the-concept demonstration. Using EEG control, a subject should utilize a mentally emulated (1-to-2)(2) EEG redundant demultiplexer to control two motors of a robot arm to move the arm from a start region A towards a goal region B avoiding an obstacle C along the way.", + "texts": [ + " 9 shows an EEG demultiplexer with an EEG sentence encoded in two frames. In this realization, both the address and the data d are decoded from the same EEG frame. Address will be binary valued, and data d will be integer valued. In each EEG frame the subject tends to generate an EEG feature (e.g., increased amplitude of the alpha rhythm) that will be interpreted by both the EEG-to-address and the EEG-to-data decoders. To give a proof of the EEG demultiplexer concept, a demonstration scenario is designed as shown in Fig. 10. The subject observes a scene in which a robotic arm should move from a start region A to a goal region B, while avoiding an obstacle C along the way. The scenario uses a 5-motor robotic arm in which only two motors are allowed to move (i.e., two degrees of freedom are allowed), while the other three motors are fixed in some predefined positions. The controlled motors are motor (torso) for azimuth (horizontal) movement, and motor (wrist) for elevation (vertical) movement. The robot arm moves horizontally, and at the point in which the robotic arm reaches the obstacle, its horizontal projection should be less than , where is the horizontal distance between the robot arm base and the obstacle" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002690_robio.2007.4522277-Figure7-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002690_robio.2007.4522277-Figure7-1.png", + "caption": "Fig. 7. Manipulating a thin object by two robot fingers in 2D plane", + "texts": [ + " Physical parameters of the finger and object and control gains of this simulation are shown in Tables III and IV respectively. The noteworthy difference between the non-redundant case of Fig.4 and redundant case of Fig.6 is that the convergence speed of the non-redundant case is faster than the redundant one. However, the redundant finger is capable to realize stable grasping and angle control simultaneously with the aid of redundancy in finger joints. Next consider precision prehension of a 2D rigid object by a pair of single DOF robot fingers with soft ends (see Fig.7). Almost all symbols used in this section (see Fig.7) are defined with extended meanings of the symbols already appeared and explained in Section II. In a similar way to that in Section II , \u0394xi and xi can be expressed as \u0394xi = ri + li + (\u22121)i (x\u2212 x0i) T rX , i = 1, 2(29) xi = x0i \u2212 (\u22121)i (ri \u2212\u0394xi) rX , i = 1, 2 (30) and rolling constraints between finger tips made of soft material and the object surfaces can be expressed as \u2212 (ri \u2212\u0394xi) d dt \u03c6i = d dt Yi, i = 1, 2 (31) where \u03c6i = 3\u03c0/2 \u2212 (\u22121)i\u03b8 \u2212 qi1(i = 1, 2). By setting the Lagrangian as L = K\u2212P in the same way as in Section II and applying the variational principle that is similar to equation (15), we can derive Lagrange\u2019s equation of motion of the overall finger-object system: Iiq\u0308i1 \u2212 (\u22121)ifiJ T i (qi1) rX +\u03bbi { (ri \u2212\u0394xi)\u2212 JT i (qi1) rY } = ui, i = 1, 2 (32) M x\u0308\u2212 (rX , rY ) (f1 \u2212 f2,\u2212\u03bb1 \u2212 \u03bb2) T = 0 (33) I\u03b8\u0308 \u2212 f1Y1 + f2Y2 + l1\u03bb1 \u2212 l2\u03bb2 = 0 (34) By considering and extending the proposed control signal of eq" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002034_2007-01-2860-Figure6-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002034_2007-01-2860-Figure6-1.png", + "caption": "Figure 6: Camber angle", + "texts": [ + " Detailed investigation of all these basic characteristics during vehicle conception phase is essential. Camber and toe parameters directly interfere with vehicle behavior in maneuvers. A brief description is made about this influence in vehicle handling behavior. Camber (\u03b5w) is the angle formed between the vertical line of each wheel and its respective center axis, in the vehicle front view. By definition, camber is positive if the top part of the wheel slopes outwards, and the opposite slope is defined as negative, as shown in Figure 6. When the vehicle has a load equivalent to three people, for example, slightly positive camber values should be used because, in wavy roads, the wheels trend to be in a perfect vertical position, thus decreasing roll and tire wear (Reimpell 1986). Negative camber values are preferred in unloaded vehicle projects, mostly aiming to improve behavior in curves, improving outer wheel behavior. In independent and twist-beam rear suspensions, negative camber angles tend to be caused by the vertical bump movement, as in Figure 7, explaining the option the initial camber values, as discussed" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002249_ijtc2007-44023-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002249_ijtc2007-44023-Figure2-1.png", + "caption": "Figure 2. Test Bearing Configuration", + "texts": [ + " iF : Excitation force (N) Schematic diagram of the test rig which is used on the experimental study is shown in Figure 1. Test bearing is settled in the middle of the test rig enclosing journal with 300mm diameter. Rotating shaft is with 2m long and supported with two ball bearings at the end of the shaft. Air bellows is located under the bearing casing to supply the load to the bearing. A 225kW electric DC motor is connected with shaft. Lubrication system supplies oil of ISO VG32 grade with test bearing. Test bearings are shown in Table1. Figure2 shows configuration of test bearings. 1 Copyright \u00a9 2007 by ASME ghts Reserved. l=/data/conferences/ijtc2007/71808/ on 07/19/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use Copy Down To evaluate the static characteristics of partial tilting pad journal bearing, oil film pressure, oil film thickness and pad temperatures are measured during the test. Thermocouples installed at bearing pads are presented in Figure 3. Table2 shows test conditions Table2 Test Conditions Case Rotation Speed, rpm Load, kN Flow Rate, liter/min" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001950_1-84628-179-2_5-Figure5.2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001950_1-84628-179-2_5-Figure5.2-1.png", + "caption": "Fig. 5.2. Thrust vector of the main rotor.", + "texts": [ + " The vector thrust of the main and tail rotors are described by TM = T 1 ME1 + T 2 ME2 + T 3 ME3 (5.2) TT = T 1 T E1 + T 2 T E2 + T 3 T E3 (5.3) However, it is well known that the tail rotor has no swashplate. Then, the thrust vector of this rotor always has the same direction, i.e. in the direction of the E2 axis, so equation (5.3) can be rewritten as TT = T 2 T E2 (5.4) The components of the thrust vector for the main rotor can be defined as a function of an angle \u03b2 called the flapping angle which denotes the tilt of the main rotor disk with respect to its initial rotation plane (Figure 5.2). This angle is formed by the tilts of angle a (longitudinal flapping) and angle b (lateral flapping) that we have assumed to be measurable and controllable variables (Assumption 5.3). By simple geometric calculus, we have tan2 \u03b2 = (T 1 M )2 + (T 2 M )2 (T 3 M )2 (5.5) tan2 \u03b2 = tan2 a + tan2 b (5.6) 1 \u2212 cos2 \u03b2 cos2 \u03b2 = sin2 a cos2 a + sin2 b cos2 b (5.7) Simplifying the last equation, we obtain the expression cos \u03b2 = cos a \u00b7 cos b\u221a 1 \u2212 sin2 a \u00b7 sin2 b (5.8) The term T 3 M is directly obtained by the projection of the thrust vector on the axis E3", + " The first part is composed of the total translational forces applied to the coaxial helicopter and the second part is related to the sum of the rotational torques. More details on coaxial helicopter dynamics can be found in [78]. Denote by TA and TD the thrusts generated by the above (A) and down (D) rotors respectively. Then, the thrust vectors are defined by TA = T 1 AE1 + T 2 AE2 \u2212 T 3 AE3 (5.46) TD = T 1 DE1 + T 2 DE2 \u2212 T 3 DE3 (5.47) It is desirable to express the thrust components in terms of the cyclic tilt angles a and b which form the system inputs. Define \u03b2 as the measure of the tilt of the rotor disk (see Figure 5.2). Thus, the thrust components can be directly computed by the projection of TA or TD onto the axis of the body fixed frame by TA = G(a, b) \u00b7 |TA| (5.48) TD = G(a, b) \u00b7 |TD| (5.49) where G(a, b) are defined in equation (5.13). Another force applied to the coaxial helicopter is the gravitational force given by (5.15). Denote by f the total translational force applied to the helicopter and expressed in the inertialfixed frame: f = RG(a, b)(|TA| + |TD|) + mgE3 (5.50) The torques generated by the thrust vectors TA and TD, represented by \u03c4A and \u03c4D respectively, are due to separation between the centre of mass and the rotor hubs" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002461_bf02916353-Figure6-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002461_bf02916353-Figure6-1.png", + "caption": "Fig. 6. Schematic diagram of gimbal system.", + "texts": [ + " It is designed to adjust the flow provided to the hydraulic cylinder by servo valves (NG6,BOSCH). The stability part is designed to make it possible to individually operate and control the movement in the direction of roll and pitch by universal joint. A tiltmeter is set to measure a range of \u00b130\u00b0(\u00b15 VDC) changes of the horizontal angle of each plate. For the control part, two Data Acquisition Board(DR1010) is set up in a computer (Pentium ) and the control program is written in using LabWindows based C language. Figure 6 is the picture of the gimbal system and Fig. 7 is the schematic diagram of electro-hydraulic system which is used in this research. The nominal values of the hydraulic system parameters are listed in Table 1. This paper simulates and experiments using the PI 0.9 0.001, , 9 1 350, 0.002 s a d n sW W s s (29) where H controller can be obtained as in Eq.(30)." + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000878_iscas.2004.1329099-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000878_iscas.2004.1329099-Figure2-1.png", + "caption": "Fig. 2. Compass gait biped robot.", + "texts": [ + "2 The associated Jacobian entry DF5\u2212(xk 0) is also evaluated at the return point, see (27). The matrices gxx, gyx, gxy and gyy are usually extremely sparse. It has been found that often the error introduced into DF by ignoring them has negligible effect on convergence. However situations can arise where they do affect convergence. Efficient computation of these matrices is discussed in [5]. A model of the compass gait biped robot is discussed in detail in [9], with a summary given in [10]. The biped robot can be treated as a double pendulum. Figure 2 provides a schematic representation and identifies important parameters, including the incline angle \u03b3. The robot configuration is described by the support angle \u03b8s and the non-support angle \u03b8ns. Dynamic equations describe the evolution of the state vector x = [\u03b8ns \u03b8s \u03b8\u0307ns \u03b8\u0307s] T \u2208 4 during the swing phase. An event occurs when the non-support (swinging) leg collides with the ground. This establishes the triggering condition \u03b8ns + \u03b8s + 2\u03b3 = 0. The biped robot is therefore an example of a hybrid system, with walking motion corresponding to a periodic orbit" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001493_iros.2006.281826-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001493_iros.2006.281826-Figure2-1.png", + "caption": "Fig. 2 Application of a Magnetic Hair Pad.", + "texts": [ + " We can build such a device by making a structure similar to a gecko foot and by attaching a magnet to each spatula (hair). The Van der Waals interaction becomes negligible by scaling the spatulas by a factor of 100 or bigger. We call this hypothesis force substitution hypothesis. The rest of the paper is based on the validity of this hypothesis. 1-4244-0259-X/06/$20.00 C)2006 IEEE magnet, (when d -* 0). This maximum force is also known as the break away force. The distance expressed in the X-axis is normalized by the radius of the magnet. E. Magnetic Hair example Based on D, Fig. 2 represents a model of a 5 finger magnetic hair pad, modeled after a gecko foot, where the function of Wan der Waals force has been substituted by magnetic force. Each spatula ends up in a cylindrical magnet. These spatulae are around 3 orders of magnitude bigger than the original ones. However, the question remains, how good is the solution proposed in Fig. 2 compared to other systems? Table 1 reviews competing systems for wall mobility. TABLE I COMPARISON OF FOUR WALL CLIMBING MECHANISMS Device Gecko foot Gecko Hair Magnetic Hair IB Magnet[5] Force type Van der Van der Waals Magnetic MagneticWaals Hamr 0.2 ,um 80nm-0.2,um 0.2 mm lcm Not apply Production Live Forming, method creature lithography Cost High Low Low Impractical Scalable Difficult to Usability [6] miniaturize Ferromagnetic Ferromagnetic Places structures structures. where it can eno italenfor only", + " In the area left to \"Tack mode\" the hair skids. The bending of the hair that occurs in the area right of \"Release mode\" is unrealistic in a close packed (carpet-like) arrangement of hairs. (Fig. 3) From Fig. 12 we conclude that hairs in Tack mode are softer and therefore can \"stick\" better. Hairs in release mode are stiffer and thus are easier to release. E. The Moment Distribution Another effect of the characteristic curvature of the hairs is the ability to distribute economically [9] a big load (as see in Fig. 2) into smaller loads to each hair (MAdh of Fig. 5). The importance of this function, (the complementary of the peeling effect of section C), comes to the fore in systems that don't ensure uniform load distribution. In Fig. 13b a load of 63Kg on a magnetic pad causes a momentum that is not distributed uniformly into the magnets. This causes an unwanted detachment (peeling) of the inferior part that in turn causes a spiraling (and dangerous) loss of adhesion of the system as a whole. The characteristic curved shape gecko hair seems naturallyfitted [9] to avoid this issue by conveying loads and tensions acting in one end of the hairs to the substrate efficiently" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000258_peds.1997.618715-Figure3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000258_peds.1997.618715-Figure3-1.png", + "caption": "Fig. 3 GeometIy of the VCM", + "texts": [ + " Section I1 of this paper describes the method of obtaining the control model; section I11 describes the control system structure; sections IV and V are the actual implementation and the results of the proposed system. 2. Obtaining the Dynamic Model of the Actuator In this section, a full model of the dual VCM actuator is constructed. The model includes inter-axis coupling effect, spring effect, friction, and inertia. 2. I Calculation of spring effect The movement of the VCM has a nonlinear relationship with the spring position, as shown in figure 3. In this project, the force-position relationship of this spring effect is calculated from the geometry of the VCM and the spring constant. This relationship has been used in the spring compensation block of the controller. The amount of spring force induced on the voice coil is calculated by considering the free body diagram of the wedge and the coil bearing as shown in figure 4(a) and 4@). angular surface \\ \\ \\ Fspring '\\\\\\ // 2.2. Estimation offriction on the small VCM Due to the small size and the geometry of the small VCM, friction becomes a dominant factor" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000096_s1044-5803(03)00014-7-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000096_s1044-5803(03)00014-7-Figure1-1.png", + "caption": "Fig. 1. Schematic drawing showing the sectioned portion of the samples used for the impact test.", + "texts": [ + " Different substrate temperatures were generated by cooling or preheating substrate to 0, 200, or 400 C before the laser cladding treatment. The substrate temperature was determined using Pt\u2013 Rh thermocouples during the laser cladding treatment [6]. The samples of the laser-clad bronze coating were prepared using the following parameters: laser power P= 2 kW; beam diameter d = 3 mm; relative powder flow rate G = 20 mg mm 1; and scanning velocity ranging from 6 to 16 mm s 1. Laser-clad single tracks with dimensions of 2 mm width, 8 mm length, and 10 mm height were taken out by wire cutting in the directions shown in Fig. 1. The arch part of the clad was ground to a flat surface such that the coating thickness was about 1 mm. The final total sample thickness was measured using a vernier caliper. Cross-sections were taken of the samples, then polished and etched\u2014the aim being to observe the cracking morphology during the impact test. The etching reagents were: for the iron\u2013base coating, 70 vol.% HNO3 + 30 vol.% HF, and for the bronze coating, 50 vol.% HNO3 + 50 vol.% CH3COOH. The impact loading experiments were carried out using an apparatus shown in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000006_jsvi.2002.5135-Figure4-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000006_jsvi.2002.5135-Figure4-1.png", + "caption": "Figure 4. The sketch for locking joint with clearance.", + "texts": [ + " With the cubic Hermite interpolation functions for the beam element [18] in the interval \u00f0 1=2; 1=2\u00de j1 \u00bc 1 2 3 2 x\u00fe 2x3; j2 \u00bc l\u00f01 8 1 4 x 1 2 x2 \u00fe x3\u00de; j3 \u00bc 1 2 \u00fe 3 2 x 2x3; j4 \u00bc l\u00f0 1 8 1 4 x\u00fe 1 2 x2 \u00fe x3\u00de; \u00f011\u00de where x \u00bc x=l; the stiffness matrix of the element can be obtained as Ke \u00bc EA l 0 0 EA l 0 0 12EI l3 6EI l2 0 12EI l3 6EI l2 4EI l 0 6EI l2 2EI l sym EA l 0 0 12EI l3 6EI l2 4EI l 2 6666666666666666664 3 7777777777777777775 : \u00f012\u00de The generalized force can be transformed into the global system Qe si \u00bc RePe si; Qe sj \u00bc RePe sj : \u00f013\u00de The locking joint with clearance is shown in Figure 4. ya and yb denote the clearance angles. When y4y1; the beam enters the region of locking, and a serial of impact might take place. While the condition ya > y > yb is satisfied, the beam is in the clearance and is free from impacts. When condition y5ya or y4 yb is satisfied, the impact takes place. The continuous contact force model with hysteresis damping in reference [16] is used for the locking joint. The original contact model is for linear contacts. In this paper, a contact model for angular contacts is presented", + " The damping coefficient C \u00bc 2Mon ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi \u00f0ln e\u00de2=\u00bd\u00f0ln e\u00de2 \u00fe p2 q ; \u00f019\u00de where e is the coefficient of restitution appropriate for the initial impact velocity, and on \u00bc ffiffiffiffiffiffiffiffiffiffi k=M p is the natural frequency of the beam on the linear spring. In the case of revolute locking joint, which has the angular form, the contact force model has to be transformed into the angular form. As shown in Figure 4, let yc be the angular form of the indentation d; the following equation is obtained: yc \u00bc d xp : \u00f020\u00de The occurrence of contact between the beam and locking joint is determined by evaluating variable yc at any time during the numerical integration of the system equations of motion as yc \u00bc y ya if y5ya; y\u00fe ybj j if y4 yb; 0 if yb5y5ya: 8>< >: \u00f021\u00de Equation (14) has the following form: f \u00bc K\u00f0xpyc\u00de3=2 \u00fe CT\u00f0yc\u00dexp \u2019yc; \u00f022\u00de where the damping function T\u00f0yc\u00de is T\u00f0yc\u00de \u00bc \u00bd\u00f0yc \u00fe ycj j\u00de=2yc exp\u00bdf\u00f0yc ec\u00de yc ecj jg\u00f0Q=ec\u00de ; \u00f023\u00de in which ec \u00bc e=xp: The generalized force due to the impact between bodies i and j are Qi \u00bc 0 0 N 2 64 3 75; Qj \u00bc 0 0 N 2 64 3 75; \u00f024\u00de where N is given by N \u00bc Kt\u00f0y\u00fe ypre\u00de if y > y1; xpf \u00fe Kt\u00f0y\u00fe ypre\u00de if y15y5ya; xpf \u00fe Kt\u00f0y\u00fe ypre\u00de if y24y4 yb; Kt\u00f0y\u00fe ypre\u00de if yb5y5ya: 8>>< >>: \u00f025\u00de With flexible beam modelling in section 3" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001809_2005-01-0384-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001809_2005-01-0384-Figure1-1.png", + "caption": "Fig. 1 FE Model of a Tire", + "texts": [ + " Finite element methods have been used to analyze stiffness, vibration, strength, crashworthiness and other characteristics of vehicle bodies, and are recognized as effective measures in the vehicle development process. Finite element analysis could also be an effective method of finding the nonlinear characteristics of the tires, for example, cornering force and braking torque, due to the large tire deformation that occurs under extreme vehicle operating conditions(2). However, it has not been applied to analyze vehicle dynamic performance since it is difficult to stably compute the large deformation of rotating tires. In this study, the FE model of a tire shown in Fig. 1 was used to simulate the dynamic behavior of a vehicle. In this tire model, the rubber parts were modeled with solid elements, and the belts and fibers were modeled with shell elements. The detailed structure of tire, for example, the tread block patterns, was ignored to obtain an incrementation time of 1.0-6 s with an explicit finite element method for solving vehicle dynamic behavior in seconds. Air pressure in the tire was modeled by applying pressure to segments on the tire's inner surface. The tire pressure was constant over time in this simulation", + " The three test conditions A, B and C were simulated using the measured vehicle speed as the initial velocity and the measured steering profile as the boundary condition. Table 1 gives the measured initial velocity in each test, and Fig. 8 shows the measured rack displacement. The data in Table 1 and Fig. 8 were the inputs used in the vehicle cornering simulations conducted in this study. Table 1 Initial Velocity in Cornering Simulations Fig. 8 Measured Rack Displacement Simulation of vehicle braking behavior Vehicle braking behavior was also simulated on a flat, dry asphalt road surface using the FE tire model shown in Fig. 1 and the vehicle model shown in Fig. 3. The vehicle speed and braking torque applied to the wheels were first measured in a driving test. The measured vehicle speed just before braking was input as the initial velocity and the braking torque as the load condition of the four wheels. Figure 9 shows the measured braking torque used in the braking simulation. In this measurement, the driver applied a steady force to the brake pedal. However, braking torque fluctuations at the wheels are observed in Fig", + " The maximum acceleration in Test C was close to the upper limit of acceleration of the vehicle on the dry asphalt road surface used in the tests. The maximum acceleration calculated under both conditions, including the extreme condition, showed good agreement with the driving test data. Figure 19 shows the peak angular velocity at the vehicle\u2019s center of gravity in Tests A, B and C. It is seen in the figure that the calculated maximum angular velocity agreed well with the measured values. Simulation of vehicle braking behavior Vehicle braking behavior was simulated with the FE tire model shown in Fig. 1, the vehicle model shown in Fig. 3 and the measured braking torque applied to the wheels in Fig. 9. Figure 20 shows simulated vehicle attitudes before braking and during braking, and it is seen that the nose of the body dives due to braking. Figure 21 shows the calculated and measured longitudinal acceleration at the vehicle\u2019s center of gravity. The calculated time history of longitudinal acceleration shows good agreement with the driving test results. A sedan type vehicle was modeled using finite elements to represent the tires and kinematics of the front and rear suspension systems in order to simulate the dynamic behavior of the vehicle" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002766_sice.2008.4654924-Figure6-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002766_sice.2008.4654924-Figure6-1.png", + "caption": "Fig. 6 Coordinates in the interface", + "texts": [ + " In this paper, we fix only position of the robot on the center of the monitor. When the robot rotates, the robot model on the monitor is rotated based on the 3D motion sensor. In addition, the north-up local environment map is always displayed on the monitor. The north direction can be measured by the 3D motion sensor. The map data and the posture of the robot are represented by the robot coordinates. On the other hand, the touched point data is represented by the display coordi- - 1635 - nates. We define these coordinates as shown in Fig. 6. The positions of the origins of these coordinates are the same place. However, the orientations of these coordinate frames are different since the local environment map is always the north-up one. We transform the touched point which is inputted by the operator to robot coordinates from display coordinates. The transformation is expressed as \ufffd xRd yRd \ufffd = \ufffd cos(\u2212\u03b8) \u2212 sin(\u2212\u03b8) sin(\u2212\u03b8) cos(\u2212\u03b8) \ufffd\ufffd xDd yDd \ufffd , where xDd and yDd are the target point with respect to the display coordinates, xRd and yRd are the target point with respect to the robot coordinates and \u03b8 is the relative angle between two coordinate frames" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000995_rissp.2003.1285644-FigureI-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000995_rissp.2003.1285644-FigureI-1.png", + "caption": "Figure I . Environment with a mobile robot and a moving obstacle", + "texts": [], + "surrounding_texts": [ + "searching, we used for reference the idea of punish-function, the sign-bit is set on the each item, then originally target function change into:\nk\n/(x) = /(X)+P 0w(10,10) * Z p , ( 5 ) id\nit is named for expanded target function, where p i is\nthe sign ofthe constraint function. From the formulation (5) , we know that expanded target function of the satisfaction the condition is the same with the originally target function and that of the dissatisfaction the condition is evaluated a great number, and the solution of the dissatisfaction the constraint condition is wash out. Aftertaxis, the minimum-value of the independent variable is the initial of the next time cycle and that is record the optimal solution, Optimalx,\nin approximately searching nowadays. Of course, the term of the chaotic consequence shall he enough more. Step 4 . Better searching. After step 3, fmt, by Logistic mapping, the variables are translated into the interval of [-1,1] in the form\nYln+l = -1+2xi,,+, ( 6 )\nthen second mapping\nxj,\u201d =optimal++a,y;., , *random (7)\nwhere, ai is the adjustable coefficients, commonly it is 0.01-0.05, random is random, Optimalxi is the\noptimal solution currently. It is helpful to speedup the speed of the searching that the introduced pseudorandom can further short the interval of the searching because of the adjustable coefficients. In the searching process, the function-constraint and the variable-constraint shall be satisfied at the same time during the second searching since the variable-constraint may dissatisfaction more than the approximately searching. If subject to the constraint and f(i) < f(0p)ptimalx) then updating the better solution Optimalxi and the better values currently, otherwise to evaluate f ( 2 ) a\nconsiderable value. Step 5 . Optimal searching. After step 4, first, searching in\nthe form:\nx,? = oplimalx, +a,y:,, *random3\nwhere y,+l =i-p~y, is Ulam-von Neumann mapping\nand it is further shorten the searching interval that Pseudorandom random is introduced, second, estimating subject to the constraint, and to record optimal solution and optimal values currently. Step 6. To shorten the computation time, if optimal values don\u2019t vary a lot of times, then to shorten the interval of searching on purpose. Step 7. Repeat step 4, 5, 6. If its terminaterule is contented, then searching is end and printing optimal solution and optimal values currently; otherwise to return step 4.\n( 8 )\n2 .\n3 Mathematical Modeling and Problem Statement\nThe mathematical model about the motion planning for mobile robot[8]is stated as follows: Performance Criterion:\n1 2 J ( i ) = -[f(s,) - TI\u2019\nDynamic Model:\nWhere\nX ( S ) = [~(s) ,v(s)]\u2019 , V(S) = -, t ( ~ ) , V ( S ) is\nthe time and velocity respectively when the robot is at s.\nx\u2019(s) = A[+)] + B[x(s) ]u(s )\nds dt\n1 - 1 1\nWhere m is the mass of the robot, I is the moment of the rotational inertia of the robot, u(s) is the steering force of the tangent to r ( s ) at every instant, j ( s ) is the curvature ofthe trajectory r ( s ) .", + "Assume the Initial-Final Conditions:\nx ( so ) =[o,v,l', =[free,v,l\nInput Constraints:\n-U , 5 U m i n ( S ) 5 u ( s ) 5 u,(s) S U I\nState Constraints:\nV,i,(S) 5 4 s ) v,,,(s)\nCollision Avoidance Constraints: d(s,t) 2 do\nWhere u,,,'(s), umi,(s),v,,,(s),vmi.(S) is the limit\nof the steering force of the tangent to r(s) and the\nvelocity respectively; U, , U, is the max i \" thrust and break forces respectively; do is the safety constant for\nthe robot, it is decided for the shape of the robot and the\nobstacles.\n4 Numeral Simulation\nIt is difficult to solve the problem of motion planning for mobile robot directly. To simply the problem, we use the\nmethod of the descretization the domain[O,s,] [9]with\nA ~ , =L, where Lis the length of r ( s ) , N IS the steps of N\nthe descretization, x, is the controllable variable, the\ntimeduring'rohotmoving AY, =fi, i=1 ,2 ; . . ,N\nThen the above nonlinear problem can be change into the form as follows:\nxr\nWhere, ..,=F 8 7 is the friction coefficient\nbetween the wheels and the floor,\nlh)\nVmln,vma is the minimum velocity and maximum\nvelocity respectively. f ( s , ) is the curvatnre of the robot\nat the i th step, T is the time of the nominal plan for\nthe r(s) . d(s,,x,), do is the distant between the\nobstacle and the robot at s i , a safety constant for the\nrobot respectively. Consider the environment of figure 1\nWhere the curvatnre each step is\nf ( s , 1 = ' ' ' f(s,,) = 0\nf ( f 3 8 ) =...f(saz)=O f(s,,) = ...f( s,,) = 112 f ( s 7 4 ) = ... f(s113 1 = 0 f ( s , , , ) =...f(s121)=1 f ( s l , , ) = . . . f ( S l 4 6 ) = o\nf ( s Z 6 ) =. . . f ( s3?)= 213\nOthers parameters:", + "N = 1 4 6 , T=12.1753, m=60kg\nI = 32kg. m 2 , U\\ = 140N, vo = vI = 0, q = 0.3, L = 14.6m\nvmin = O.lm/sec, v,, = Smlsec\ndo = 0.7ni\nU2 = 60N\nThe computed result by the algorithm is in the figure 2, figure 3\n5 Conclusion\nThe ergodicity and mndomicity of the improved chaotic optimization based on the chaotic motion of the inherent regularity, its process of search is decided that regularity of chaos itself and Ulam-von NeumaM mapping can be ergodic in the range of the constraints. Further, there is simple structure and less computation on the improved chaotic optimization algorithm and the convergent speed of tbe method is faster and its accuracy is better. It can\nresolve large scale linear or nonlinear problems (there are 146 variables here) . But the parameter in the algorithm is decided by real problem, it is further research that a universal method of selecting parameter is found.\nReferences\nB. Haq From parabla-An introduction to chaotic dynamical system, Shanghai Scientific and Technological Education Publishing House, 1993. B.Li, W. Jiang, \u2018Chaos Optimization Method and its\nApplication\u201d, Control Theory and Applications, vol. 14, no.4, pp 613-615, Aug. 1997. W. Tang, D. Li, \u2018Cbaatic Optimization for Economic Dispatch of Power Systems\u201d, m Proceedings of /he CSEE, 2Mx). 20(10),pp. 3640. T. Wang, H. Wang, Z. Wang, \u2018Mutative Scale Chaos Optimization Algorithm and its Application\u201d, Control and DecisioR vo1.14,no.3, pp. 285-288, May 1999.\nK Aihara, T. Takabe M. Toycda, \u2018Chaotic Neural Nework\u201d, Physics Letter A, volI44, m.6, pp. 333 -340, 1990. Z. Wang, T. Zhang, H. Wang, \u201cSimulated Annealing Algorithm of Optimization Based on Chaotic Variable\u201d, Contml and Decision, vol.14, no.4,pp. 381-384. 1999.\nY Wag, J . Liu, Y. Sun, \u201cHybrid Genetic Algoritlim Based on Mutative Scale Chaos Optimization Strategy\u201d, Control and Decisioh vo1.17, no.6,pp. 958960,2002.\nI. Konstantinos, \u2018\u2018a Supervisory Control Strategy for Navigation for Mobile Robots in Dynamic Environments\u2019\u2019, Rensselaer Polytechnic Institute Ph. dissertation, 1991. M. Sun, P. Cheng, \u201cRule-based Reaktime Motion Planning for Mobile Robots\u201d, Control and Decision, vol. 12, no.4, pp. 322-326, July 1997.\n[IO] Z. Chen, H. Shi, \u201cApplication of Chaotic Optimization Algorithm to Problems with Constrained Optimization\u201d, Control and Decision, vol. 17, no.Suppl.2, pp 111-114,\n2w2.\nJ . Latombe, \u2018Robot Motion Planning\u201d , Kluwer Academic [ I l l Publishers, 1991" + ] + }, + { + "image_filename": "designv11_69_0001205_pime_proc_1972_186_090_02-Figure12-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001205_pime_proc_1972_186_090_02-Figure12-1.png", + "caption": "Fig. 12. Seal configuration", + "texts": [ + " Part-sectioned view of the experimental engine Proc lnstn Mech Engrs 1972 Vol \u2018I 86 62/72 at UNIV OF CINCINNATI on June 4, 2016pme.sagepub.comDownloaded from ROTARY PISTON MACHINE SUITABLE FOR COMPRESSORS, PUMPS AND I.C. ENGINES 749 a Rotor close to casing. b Rotor away from casing. Fig. 14. Section through third design of apex seal that this particular rig did not exploit the potential for high speed operation and, at the same time, it made heavy demands on the standard of sealing required. 7.1 Seals Fig. 12 shows the seal grid in schematic form to illustrate the names used for the various seals and the position of the enlarged section shown in Figs 13 and 14. The sealing of each chamber is made up from four pieces. (1) A leading apex seal. (2) Half the circumference of the hub ring. (3) A trailing apex seal. (4) A tip seal. The two apex seals are the least conventional because, like the apex seals of the Wankel engine, they tilt as they slide and are therefore limited to line contact as opposed to Prac I nstn Mech Engrs 1972 surface contact" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001580_ramech.2004.1438922-FigureI-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001580_ramech.2004.1438922-FigureI-1.png", + "caption": "Fig. I., Schematic description of SA-PM", + "texts": [], + "surrounding_texts": [ + "I. INTRODUCTION\nParallel manipulators have the advantages of high stiffness, high accuracy. low moment of inertia and high payload ca pacity. However, the six motion freedoms of the end-effector are usually coupled together due to the multi-loop kinematic structure in the manipulator. Hence, motion planning and can\u00b7 tral of the end-effector for such lype of manipulators become very complicate. Here we propose a new kinematic architec ture for a 6-DOF parallel manipulator that can decouple the three rotational and three translational motions. This parallel manipulator relies on a newly developed 2-DOF actuator that can produce both rotation and translation along the same axis. With different actuator motions, the manipulator may generate different types of movements. We term this manipulator a Selectively Actuated Parallel Mallipulators, or simply an SA PM. With decoupled rotation and translation, the kinematics of the SA-PM is simplified so as to facilitate its motion planning and control. This will facilitates the development of high precision optical alignment tools that requires both orientational and positional accuracy.\nThe development of SA-PMs is originated from various types of 3-DOF parallel manipulators (PM). In the literature, the 3-DOF PMs can be classified into three groups according to the motion of the moving platform, i.e., PMs with three translational DOFs ([1], [2], [3], [4], [5], [6]), with three rotational OOFs ([7], [8], [9], [10]) or with hybrid three OOFs ([11], [12], [13]). In the aspect of translational PMs, Tsai and Joshi [4] enumerated a class of PM with only translational DOFs and studied the kinematics of the 3-UPU PM. Subsequently, Ii and Wu [61 studied the kinematics of an offset 3\u00b7UPU translation PM, and demonstrate that the\nforward kinematics has 16 solutions instead of two in the zero offset manipulator. Korig and Gosselin [3] proposed a CRR translational parallel mechanism which is isotropic in its entire workspace. For PMs with three rotational DOFs, Gosselin et al. [14] studied the kinematics a 3\u00b7RRR spherical manipulator which shows there are 8 solutions for the forward kinematics of this manipulator; Karouia and Herve [IOJ proposed a 3- UPU PM for spherical motion; Di Gregorio.studied the 3-URC wrists [8J and the 3-RUU [9] PM for 3-DOF rotations.\nIn [16]. we have studied type synthesis and singularity analysis of the new SA-PM. This paper will focus on the kinematics of the SA-PM. It will be demonstrated that the position and orientation can be computed separately due to the decoupled motion architecture. The closed-form solutions for both inverse and forward kinematics have been fonnulated, which provides a foundation for further investigation of the SA-PM.\nII. STRUCTURE DESCRIPTION OF SA-PM\nThe SA-PM (Fig. 1) is built with a moving platform and a base connected by three identical serial chains in parallel. Each of the three 2-DOF actuators (Ail, Ai2), (i = 1,2,3, representing the number of the limb) is mounted on the base, followed by passive joints P (Ai3), P (Ai4), R (As), R (Ai6) (P, R represent prismatic and revolute joints respectively). Note\n0-7803-8645-0/041$20.00 \u00a9 2004 IEEE 231", + "Fig. 3. The 2-DOF actuator module [15]\nFig. 4. The first limb of the SA-PM\nthat the three actuator axes are mounted mutually orthogonal and intersect at point O. In each limb, the two passive p. joint axes are both perpendicular to the actuator axis, and the last two passive R-joints in every limb must intersect at one common point 0'. To verify its feasibility, a CAD model of the PM is constructed as shown in Fig. 2. Figure 3 shows the 2-DOF actuator module that combines a pure rotation and a pure translation On a single shaft [15]. Depending on the application scenario, the actuator can produce I\u00b7 DOF rotary motion, I-DOF linear motion, l-DOF screw motion, or the combined 2-DOF motion using one or both motors residing in the actuator.\nIII. KINI3.MATICS ANALYSIS OF SA-PM\nA. Frames of Reference and Vector Representation\nAs shown in Fig. 1, the base frame {a - xy z} is located at point a with its x-, y- and z- axis along with the first, second and third actuator joint axis respectively. The moving platfonn frame {O' - xIY'Z'} is located at point 0'. It's z' -axis is perpendicular to the plane detennined by three points Aw, A26,A36 and points to the centroid point NI of the equilateral triangle A16A26A36. It's y' -axis is located in the plane determined by three points 0', A 16 and M and has the opposite direction with vector 0' Ai6. The x' axis is obtained by the right-hand rules.\nTaking limb I as an example, the local frames associated wilh each link assembly are defined as shown in Fig. 4, which shows the initial configuration of limb 1 (Olj = 0 (i = 1; j = 1,2,\",,6)). In this configuration, frame {Aid and frame {Ai2} are in coincidence, and the plane determined by vectors 0' Ai5 and 0' Aio is perpendicular to the plane detennined by the three point\ufffd A16, A26 and A36. Note that the origin of frame {Aid and frame {Ai6} is located at the common point 0', and thex,6 axis is in the plane determined by three points, namely, 0', Ai6 and !vI. Because of the symmetrical structure, the local frames associated with each joint of limb 2 and 3 are the same as limb I. The geometry of the moving platform can be thought of as a pyramid with triangular base (A16A26A36) as shown in Fig. 5.\nFor describing the geometry, some notations are defined as follows:\n\u2022 $ij: a unit screw associated with the jth joint axis of the ith limb with respect to local frame (i = 1,2,3; j = 1,2, . . \u00b7,6); \u2022 Bij: a vector along the jlh joint axis of the ith limb with respect to base frame (i = 1,2,3; j = 1,2, . . . ,6); \u2022 L 1: position displacement of frame {Ail} and base frame in initial configuration; \u2022 L2: position displacement of frame {A;3} and {An} in initial configuration; \u2022 L3: position displacement of frame {Ai4} and frame {A;3} in initial configuration; \u2022 a: the angle between Si3 and 8i4; -y: the angle between Si5 and Si6; \u2022 [P4x,P4y,P4z]: the position of frame {Ai5} relative to frame {Ai4} in initial configuration;" + ] + }, + { + "image_filename": "designv11_69_0003180_s12289-009-0660-0-Figure3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003180_s12289-009-0660-0-Figure3-1.png", + "caption": "Fig. 3 a Stress State in the Cup Wall. (a1 and b1 represents the dimensions namely width and height). b Wrinkling tendencies shown in stress and strain space", + "texts": [ + " Figures 4a\u2013d show the sample distribution of the radial and hoop strains for different Aluminum alloy blanks of grade Aluminum alloy 5005 which is subjected to four different types of heat treatments as said above. From these Figures it is noted that the slope of the er\u2212e\u03b8 strain curve changes suddenly when a wrinkle develops on the blank. The reason for this is that there is an increase in circumferential strain (e\u03b8) when the wrinkle is about to develop on the blank (Fig. 4). Theory of plasticity The wall section of a truncated Conical cup (a partially drawn cup) is approximated to a flat strip of width \u2018a1\u2019 and height \u2018b1\u2019 as shown in the Fig. 3a, corresponding respectively to the mean circumference and the generator of the cone. The strip is subjected to plane stress, specified by \u03c3r and \u03c3\u03b8. as shown in Fig. 3b Assuming the material to be rigid-plastic and obeying the Hill new yield criterion [8] together with the Levy-Mises flow rule, the increments of strain can be expressed as follows: de1 \u00bc dl s1 \u00fe s2\u00f0 \u00dea 1\u00fe 1\u00fe 2R\u00f0 \u00de s1 s2\u00f0 \u00dea 1 h i de2 \u00bc dl s1 \u00fe s2\u00f0 \u00dea 1\u00fe 1\u00fe 2R\u00f0 \u00de s1 s2\u00f0 \u00dea 1 h i de3 \u00bc dl 2 s1 \u00fe s2\u00f0 \u00dea 1 h i \u00f02\u00de Where \u2018a\u2019 is yielding behaviour constant. The ratio of the in-plane strain increments can be expressed as de2 de1 \u00bc s1 \u00fe s2\u00f0 \u00dea 1 1\u00fe 2R\u00f0 \u00de s1 s2\u00f0 \u00dea 1 s1 \u00fe s2\u00f0 \u00dea 1\u00fe 1\u00fe 2R\u00f0 \u00de s1 s2\u00f0 \u00dea 1 \u00f03\u00de Expression (3) can be rewritten as b \u00bc 1\u00fe a\u00f0 \u00dea 1 1\u00fe 2R\u00f0 \u00de 1 a\u00f0 \u00dea 1 1\u00fe a\u00f0 \u00dea 1\u00fe 1\u00fe 2R\u00f0 \u00de 1 a\u00f0 \u00dea 1 \u00f04\u00de where a \u00bc s2 s1 \u00bc sq sr and b \u00bc de2 de1 \u00bc deq der \u00f05\u00de The constant \u2018a\u2019 provided in the Eqs", + " As per Hill\u2019s old yield criterion [8], putting a=2, expression (4) becomes b \u00bc 1\u00fe a\u00f0 \u00de 1\u00fe 2R\u00f0 \u00de 1 a\u00f0 \u00de 1\u00fe a\u00f0 \u00de \u00fe 1\u00fe 2R\u00f0 \u00de 1 a\u00f0 \u00de\u00f0 \u00f06\u00de In expression (6) a takes the value of zero to infinity. Substituting the limiting values b \u00bc R 1\u00fe R whena \u00bc 0\u00f0 \u00de and b \u00bc 1\u00fe R\u00f0 \u00de R when a \u00bc 1\u00f0 \u00de \u00f07\u00de When R takes the value of unity: 2 < b < 1=2 \u00f08\u00de This relationship is true irrespective of whether the material work-hardens or not. The above relationship [refer Eq. (7)] is independent of the yielding behaviour constant \u2018a\u2019. This can be checked without substituting the value of 2.0 for the above constant \u2018a\u2019, in the Eq. (4), provided in the above. As shown in Fig. 3b, this situation can be represented in principal stress space and principal strain space. The Eq. (6) can be rearranged as follows: a \u00bc R 1\u00fe b\u00f0 \u00de \u00fe b R 1\u00fe b\u00f0 \u00de \u00fe 1 \u00f09\u00de Using the Eq. (9), for the known values of \u03b2 and R, the value of a can be determined. As described elsewhere [10], the effective strain increment can be written as follows: d e \u00bc ffiffiffiffiffiffiffi 2=3 p 2\u00fe R\u00f0 \u00de 1\u00fe R\u00f0 \u00de 1\u00fe 2R\u00f0 \u00de de2r \u00fe de2q \u00fe 2R 1\u00fe R derdeq 0:5 \u00f010\u00de The Eq. (10) can also be written as follows: d e der \u00bc ffiffiffiffiffiffiffi 2=3 p 2\u00fe R\u00f0 \u00de 1\u00fe R\u00f0 \u00de 1\u00fe 2R\u00f0 \u00de 1\u00fe deq der 2 \u00fe 2R 1\u00fe R deq der ( )\" #0:5 \u00f011a\u00de Similarly, d e deq \u00bc ffiffiffiffiffiffiffi 2=3 p 2\u00fe R\u00f0 \u00de 1\u00fe R\u00f0 \u00de 1\u00fe 2R\u00f0 \u00de 1\u00fe der deq 2 \u00fe 2R 1\u00fe R der deq ( )\" #0:5 \u00f011b\u00de For the known values of der and deq, the effective strain increment d e, the ratio d e der and the ratio d e deq can be determined" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003021_jae-2010-1280-Figure3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003021_jae-2010-1280-Figure3-1.png", + "caption": "Fig. 3. (a) Z-folded tube model; (b) Sketch of folded location; (c) Folded location with artificial wrinkle; (d) Z-folded model with wrinkled elements.", + "texts": [ + " Considering the stowed state as the initial condition, the inflating gas enters the first control volume, and in turn, flows to other contiguous volume through dummy partitioned membranes. This method has basically eliminated unrealistic simulations under low flow rate. Three packaged configurations of cylindrical tube, namely Z-folded, coiled and telescopically-folded configurations, are modelled in this section. All the models are polyethylene tubes with the diameter of 0.07 m and the length of 0.8 m. Table 1 presents the material properties of cylindrical tubes. The model for Z-folded tube is displayed in Fig. 3a. The fold lines have the role of natural constriction for the gas flow, so the four dummy partitioned membranes are assigned at the fold lines. The five sub-control volumes are discretized of the model from base cap to top, which are named CV1, CV2, CV3, CV4, and CV5. The boundary of base cap is fixed, and base cap forms a little initial volume about 7.8E-5 m3. In order to avoid initial penetration of the contact segments, certain initial gap should be assigned between the tube skins or the adjacent layers of folded tube. Sketch of folded location is shown in Fig. 3b, and T1 is the gap distance, L is the distance of adjacent folded layers, \u03b1 is the folding angle. The skin length errors between inner skin and outer skin is \u2206L = 2 ( L\u2032 \u2212 L1 ) = 2T1/cos \u03b1 (8) Introduce an artificial wrinkles in folded locations can modify the geometric modelling error. Figure 3c and 3d show the detail of artificial wrinkles in the folded location. The coiled tube in Fig. 4 is rolled in spiral with a rolling diameter of 0.1 m in the form of an Archimedean linear spiral shape governed by the relationship, R = R0 + a\u03b8, in which, a = (T1 + T2)/2\u03c0, and T1 is the spacing of tube skins, T2 is the spacing of adjacent layers of folded tube), R0 is the initial radius, and \u03b8 is the sweep angle. The coiled model is built with five connected control volumes links to each other through dummy partitioned membranes", + " On the base of the merit for the two inflation rates, we get a suggestion, in which a low inflation rate is employed during the early stages of inflation, and the weak inflation spreads slowly the folded tube into a preliminary shape of inflation beam, then a high inflation rate can be used to deploy the shapes into a steady support inflation beam rapidly. The effect of membrane thickness is also investigated using the Z-folded model as described above. The skin thickness of the model shown in Fig. 3d is modified as 200 \u00b5m, which is increased about 53%. Other material properties are the same as in the preceding example. Membrane thickness of 130 \u00b5m and 200 \u00b5m represent a thin and a thick wall, respectively. Deployment of Z-folded tube with the thick wall is shown in Fig. 18. Comparing Figs 9 and 18, the deployment process with the thick wall is nearly the same as the thin wall before 0.065 second. The later deployment shapes are quite different (in Figs 9 and 18). The \u201ckink formation\u201d occurs between 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002672_ical.2007.4338895-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002672_ical.2007.4338895-Figure1-1.png", + "caption": "Fig. 1. three 7R robots", + "texts": [ + " 10 equations are derived from the matrix that has the same complex exponential forms. Then extract the coefficient of every term in polynomials and construct a 16\u00d716 real number matrix. The eigenvalues and eigenvectors of the 16\u00d716 matrix are computed for inverse kinematics of the 7R Robot in stead of expanding its determinant. An example shows that program written in C++ language runs fast enough and can be used in the real time control of an industrial robot. II. INVERSE KINEMATICS A. Geometrical Model As shown in Fig.1, there are 3 robots named Telbot. Telbot 1 and 2 hold the tool which moves around the sphere with tool\u2019s z-axis pointing to the sphere centre. Telbot 3 holds the radioactive object, without moving, which is collinear to z-axis of world frame. World frame is defined in Fig.1 so that Fig.2. D-H coordinate and structure of 7R robot the original point is located at sphere center. Longitude and latitude of sphere is defined as geography. D-H coordinate 1-4244-1531-4/07/$25.00 \u00a9 2007 IEEE. 1963 system and the general 7R robot is shown in Fig.2.The original point of the base frame is located at the cross point of 1st joint z-axis and 2nd joint z-axis .The z-axis of the base frame is coincide with z-axis of the 1st joint and the x-axis is parallel to the cross vector of 1st z-axis and 2nd z-axis at the start point" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000347_isatp.1997.615388-Figure8-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000347_isatp.1997.615388-Figure8-1.png", + "caption": "Figure 8: A case of a slope", + "texts": [ + " When it decides to get over, it changes to tumbling operation to get over the step. 4.3 A Case of Slope To estimate of inclination of the object, the robot has a half circular fingertip as shown in Figure 5 . When an incline of the floor changes and then the posture of the object followed it, the contact point between the object and the fingertip also changes on the tip surface. This contact point can be calculated from the forces and the moment from the force sensor. So the robot can estimate the posture of the object and push the object to the direction as shown in Figure 8. 4.4 Experiments With an experimental system shown in Figure 9, we made experiments on operation switching according to each situations. A 5 DoF manipulator has a 6 axis force/torque sensor which measurement range is 50[N] and 5[N.m]. The fingertip has a half circular tip with the radius of 10\". First experiment is a case of change of the friction shown in Figure 6. The object is a 50[mm] copper cube, the floor 1 is aluminum with the coefficient 0.26 and the floor 2 is rubber with coefficient 0.67. In Figure 10, when the estimated coefficient increase over 0.5, the robot switches the operation to tumbling. Second one is a case of step shown in Figure 7. The object is a brass rectangular with 50[mm] x 50[mm] x 100[mm], 2.1kg. Against three steps with 10,20,40[mm] height, the object can get over them with about 10% error of estimation. Third case of slope shown in Figure 8 uses same object with first one. Against a slope with 15[deg], we obtain a measurement result shown in Figure 11 and success to move the object to follow the slope. We merge these three cases. With an algorithm shown in Figure 12, the robot can follow any cases. 5 Step Passing with Vision Sensor If a robot uses only a force sensor, it cannot sense down steps. An object doesn't face a step edge and fall down the step. So, some sensors are needed to sense ahead and find down steps. We adopt a vision sensor to this purpose" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001783_wcica.2006.1713764-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001783_wcica.2006.1713764-Figure2-1.png", + "caption": "Fig. 2 Flight Experimental Platform", + "texts": [ + " The box, which is fixed near the center of mass to offset the helicopter center of inertia as little as possible, includes a computer system(PC104 1-4244-0332-4/06/$20.00 \u00a92006 IEEE Plus based P 500MHz), some image process and data transmitting units. To measure helicopter attitude accurately, many sensors are installed on helicopter, such as digital compass, gyroscopes, DGPS. The precise attitude is obtained by multiple sensors data fusion. Extended Kalman filtering algorithm is also used for improving computation accuracy. The flight experimental platform is shown in Fig. 2. An unmanned scale helicopter is mounted on a universal rotating shaft which in turn is connected to a pole. The center of gravity of the helicopter can be aligned with the universal rotating shaft. The universal rotating shaft allows the attitude (such as pitch, yaw and roll) of the helicopter to freely vary. The pole can glide consistently with helicopter flight on vertical orientation. The pensile weight compensates the weight of the vertical pole. It is assumed that the helicopter is in an out ground effect (OGE) condition, and the effects of the compressed air in take-off and landing are neglected" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001979_50037-5-Figure7.125-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001979_50037-5-Figure7.125-1.png", + "caption": "Figure 7.125 Symmetry constraints applied to the model.", + "texts": [ + " Also, Contact Manager summary information frame should be closed. Before the solution can be attempted, solution criteria have to be specified. As a first step in that process, symmetry constraints are applied on the quarter-symmetry model. From ANSYS Main Menu select Solution \u2192 Define Loads \u2192 Apply \u2192 Structural \u2192 Displacement \u2192 Symmetry BC \u2192 On Areas. The frame shown in Figure 7.124 appears. Four surfaces at the back of the quarter-symmetry model should be selected and button [A] OK clicked. Symmetry constraints applied to the model are shown in Figure 7.125. The next step is to apply constraints on the bottom surface of the block. From ANSYS Main Menu select Solution \u2192 Define Loads \u2192 Apply \u2192 Structural \u2192 Displacement \u2192 On Areas. The frame shown in Figure 7.126 appears. The bottom surface of the rail should be selected. After selecting required surface and pressing [A] OK button, another frame appears in which the following should be selected: DOFs to be constrained = All DOF and Displacement value = 0. Selections are implemented by pressing OK button in the frame" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002018_004051756903900209-Figure6-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002018_004051756903900209-Figure6-1.png", + "caption": "Fig. 6. Statically deflected tire showing exposed cords in the outer ply.", + "texts": [ + " At increased inflation pressures, the cord force variations obtained were qualitatively the same as those shown in Figures 7 and 8, but were shifted upward relative to the horizontal axis. At the crown in the inner ply, the measured cord forces are symmetrical about the center of the foot print (0\u00b0 position). At the sidewall in the outer ply, the pattern is different-cord tension increases (above that imposed by the inflation load) forward of the footprint and decreases aft of the footprint. A somewhat similar phenomenon can be expected to Static Defiection Problem The inflated, deflected, nonrotating tire shown in Figure 6 is not a surface of revolution, the applied loads are not distributed in a rotationally symmetric For well-exercised tires, all four crown trans ducers in tire I gave results at any inflation pressure that deviated from one another by less than 0.1 lb. Similar results were obtained from the four sidewall transducers in tire II. This accuracy is quite good, considering that there are small, unavoidable varia tions in end count and cord angle along the crown or any other circumferential line of an inflated, but otherwise unloaded, tire with embedded transducers", + " Finally, although the transducers were not cali brated in compression, any comprehensive cord forces that were monitored were plotted by using the tensile calibration characteristics of that particular transducer. The tensile and compressive cord force variations, when plotted in this fashion, seemed to be smooth. Compressive forces were easily developed, especially at the outer-ply sidewall location aft of the footprint, when the tires were overloaded and underinflated. Rolling Tire As with the static deflection problem, no meaning ful cord load formula exists for the inflated, deflected, occur at the same point in the inner ply. This can be explained by referring to Figure 6; forward of the footprint in the inner ply (and aft of the footprint in the outer ply), tensile relief forces are propagated along the cord from the load region to the transducer. Some measure of the decrease in the initially imposed inflation tension in the inner ply at the crown of the tire in the center of the footprint (0 0 position) is shown in Figure 9 where cord tension is plotted against tire load. Again, the curves all have the same shape, but are shifted upward from the hori zontal axis as the inflation pressure is increased" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002660_1081286506068823-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002660_1081286506068823-Figure2-1.png", + "caption": "Figure 2. Local coordinate system in reference and deformed configurations.", + "texts": [], + "surrounding_texts": [ + "Let represent the surface of a rigid, infinite half-space and define the right-handed orthonormal triad e1 e2 e3 such that e3 is normal to the half-space. Next, let X X1e1 X2e2 X3e3 denote the position of any point in 3 with respect to the origin O . In addition, let s X3 represent a cylindrical coordinate system with origin O and polar axis e1 and define the right-handed orthonormal triad es e e3 such that es cos e1 sin e2 and e e3 es (1) The two coordinate systems are related by the equations X1 s cos and X2 s sin (2) Consider an homogeneous, elastic membrane that on one side adheres to the half-space and on the other side is attached to the base 0 of a rigid cylindrical shaft. Delamination is possible by translating and/or rotating the shaft base to a new configuration . Let denote the midplane of the delaminated portion of the membrane that is not in contact with the shaft and identify its boundary with the function [0 2 ] through the representation X 3 : s [0 2 ] (3) The function , which defines the amount of delaminated material, may also be used to define the surface in its natural configuration: at UNIV OF MONTANA on April 4, 2015mms.sagepub.comDownloaded from 0 X 3 : s [ ] [0 2 ] X3 0 (4)" + ] + }, + { + "image_filename": "designv11_69_0002660_1081286506068823-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002660_1081286506068823-Figure1-1.png", + "caption": "Figure 1. Illustration of shaft loaded membrane (a) reference configuration, (b) delamination under normal translation, (c) delamination under rotation.", + "texts": [ + "ey Words: Non-linear membrane theory, stationary principles, complementary energy, Hencky strain measure A shaft-loaded membrane, such as the one illustrated in Figure 1, may be detached from a substrate either by pulling the shaft perpendicularly from the substrate (equivalent to the blister test of [1]) or by rotating the shaft about the substrate plane. Systems related to the normal pull-off mode have been studied [2, 3] and are relevant to applications ranging from MEMS devices [4] to intracellular binding [5]. There has also been recent interest in the role of shaft-supported thin plates found in the adhesive system of wall climbing geckos [6, 7, 8, 9] (more general information on gecko adhesion can be found in the popular literature [10, 11])" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001509_detc2005-84462-Figure4-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001509_detc2005-84462-Figure4-1.png", + "caption": "Figure 4: A limb of virtual PM 3-PvRCRR As a lower-mobility PM, the Jacobian and Hessian for the 3-RCRR are not square or cubic which consequently adds obstacles to achieve the reverse kinematic modeling. To obtain the accurate solutions, the virtual mechanism principle (VMP) [18, 19] is adopted in this paper for forward/reverse velocity and acceleration analyses. According to the VMP, every limb of a 6-DOF virtual PM should be built by adding prismatic pair or revolute pair to the initial lower-mobility PM until every limb has 6-DOF single pair. For the 3RCRR, a virtual prismatic pair is added to every limb for that its every limb has five single-DOF pair. Since all three limbs in the mechanism have one common constrain that constrains the translation along the z-axis, the axis directions of all virtual prismatic pair are the same and along the z-axis. Thus, a virtual 6-DOF 3-PvRCRR PM (Pv denotes the virtual prismatic pair) is built up, which can be represented as P0R1(P2R3)R4R5 shown in Fig. 4.", + "texts": [], + "surrounding_texts": [ + "Given position and orientation of the moving platform, the rotation center Or and the points D(i) in three limbs are derivable. Given coordinate of the rotation center Or , points A(i) and B(i) are derivable. Then, according to the Eqs. (8) and (9), there are two unknown variables ( 3\u03b8 , 4\u03b8 ) in the variable expression of the points C(i), which can be solved with following equations )( ))(( )( 43423342312 4342334233434312 )( 43423342334343 )( cssccs csccscssscD csccsssscD i y i x \u2212 \u2212+\u2212= ++= (14) Then all points are solved and the reverse position analysis is available. The symbol expressions for the 3\u03b8 and 4\u03b8 are very long. Considering the limit of the paper length, they are not given here." + ] + }, + { + "image_filename": "designv11_69_0000361_iros.1998.724587-Figure10-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000361_iros.1998.724587-Figure10-1.png", + "caption": "Figure 10: Eigenspace in several tasks.", + "texts": [ + " We make the following matrix by the upper n eigenvectors: (4) where n is determined so that the contribution factor W, becomes larger than the given threshold. n m W n =CXi / Cxi (5) i = l i = l The state vector xi is projected to the n dimension eigenspace by E and gi is obtained. gi = E T Z i E Rnxl ( 6 ) Because the eigenspace can reduce the dimension of the state vector, m, to the eigenspace dimension, n, while keeping the important features, we call this eigenspace a \u201cmotion feature space\u201d. In the eigenspace, g1 - gN make some tracks for each task (Fig.10). For the sake of memory efficiency, we pick up points g1 - g, so that the distance between each point becomes larger than the given threshold on these tracks. These Doints are the motion features 3.1 Motion Feature Space obtained in training. As the angles of the legs express the motion pattern of the hexapod, it is natural to use the angles of the . In this study, we prepare the eigenspaces for every combination of observed legs by simulation as training. For example, the eigenspace constructed by the left three legs in the representative tasks such as hexapodwalking, quadruped-walking, and manipulating a box with the forelegs is shown in Fig.10, where m is 9 and n is 3. 3.2 Motion Recognition In the recognition mode, we generate the state vector z p from the joint angles of the legs observed by vision. By using (6 ) , we calculate g p and obtain g, as the point closest to g p among the motion features g l , . . . , gr . This g, is the motion feature observed at this moment. The motion feature for the joint angles obtained in Fig.8 is shown in Fig.10 as a point e on the track of the hexapod-walking. By repeating such motion feature observation for a while, the present motion of the hexapod is recognized as hexapod-walking. 4 Conclusion In order to cooperate with a robot controlled by a human operator, autonomous assistant robots must recognize its motion in real time. Joint motion of each leg and occlusion by another leg make it difficult to recognize and track the motion of the hexapod robot. In this study, we proposed a method by which we could recognize and track the 3D position and posture of the robot by vision" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003424_s12206-009-1173-y-Figure3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003424_s12206-009-1173-y-Figure3-1.png", + "caption": "Fig. 3. Strain gauge and signal wire at the bottom of the piston ring.", + "texts": [ + " 2 shows the positions of the intake and exhaust ports around the top and second rings. In order to measure the strain on the bottom sides of the rings as they slid over the ports, we used a strain gauge with a grid width of 0.84 mm and a grid length of 2.0 mm. We attached the strain gauge on the bottom sides of our rings over the center of port width. To protect the lead wires of our strain gauge from breaking, we set terminals next to the strain gauge, then connected our lead wires and heavy-duty signal wires via terminals, as shown in Fig. 3. In order to prevent the strain gauge from damping, the strain gauge was covered with silicone sealing, as shown in Fig. 4. We ground and drilled the ring grooves and the piston crown, as shown in Fig. 5, then installed the rings and strain gauge, and finally drew signal wires from the strain gauge through the inside of the piston, and out of the piston crown. To prevent the signal wires from breaking under engine operation, we attached the signal wires to a steel sheet spring [1] with epoxy adhesive, as shown in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002877_epe.2007.4417421-Figure5-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002877_epe.2007.4417421-Figure5-1.png", + "caption": "Fig. 5: Eccentric sleeve to adjust fixed dynamic eccentricity", + "texts": [ + " 4 a stator slot and the rotor are depicted balanced illuminated for a proper measurement. A local air gap length of 550\u00b5m was measured. Up to now only the adjustment of static eccentricity has been considered. The realization of dynamic eccentricity is doe using a single eccentric sleeve between bearing and rotor shaft. This option was chosen as the arrangement with the double eccentric sleeves had made a reduction of the shaft diameter necessary preventing the usage of standard rotors for further investigations. This single sleeve arrangement is depicted in Fig. 5. In total four sets of these rings have been manufactured with eccentricity of zero, 50\u00b5m, 150\u00b5m and 250\u00b5m respectively. Each ring has a marker to indicate the direction of the eccentricity as shown in the picture. In order to investigate the properties of an induction machine reference to airgap asymmetry it is thus necessary to consider the above mentioned non-ideal properties of the whole mechanical setup. Tolerances in the manufacturing of the machine as well as the eccentric sleeves have to be identified" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001125_cdc.2005.1583435-Figure4-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001125_cdc.2005.1583435-Figure4-1.png", + "caption": "Fig. 4. Exact and asymptotic plot of W\u22121 p (\u03c9).", + "texts": [ + " P From the resultant control scheme in Figure 3 we can compute the transfer matrix function, say N , that relates the input signals [di do]T with the weighed output signals, [u\u2032 y\u2032]T , N = [N11 N12 N21 N22 ] (8) where N11 = \u2212W1MrRYrNrM \u22121 r N12 = \u2212W1MrRYr N21 = Wp(Nr(I \u2212 R)M\u22121 r + NrRXr) N22 = Wp(I \u2212 NrRYr) (9) and R is defined in (6). The weighing matrix, Wp, for the Sensitivity transfer function may be represented by Wp = s/g + \u03c9b s + \u03c9b\u03b5 (10) where it is observed that |W\u22121 p (j\u03c9)| is equal to \u03b5 1 at low frequencies and it is equal to g \u2265 1 at high frequencies. The asymptote crosses 1 at \u03c9b, which is approximately the bandwidth requirement. An asymptotic plot of W\u22121 p over the frequency is illustrated in Figure 4. On the other hand, considering unstructured uncertainty (i.e., \u2206 is a full complex matrix of appropriate dimensions) and a multiplicative input uncertainty description as the one considered in (7), the Small Gain Theorem [8] imposes a condition on the \u221e-norm of the transfer matrix function from the input disturbance, di, to the weighted control signal, u\u2032, in order for robust stability to be guaranteed, i.e., \u2016N11\u2016\u221e < 1 (11) This fact restricts the design of the feedback controller K1 since, form (8) and condition (11), we have robust stability iff \u2225\u2225\u2212W1MrRYrNrM \u22121 r \u2225\u2225 \u221e < 1 (12) In this point, we could face step 1 by finding an optimal feedback controller K1 minimizing N21 in (8) subject to the constraint (12)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003212_educon.2010.5492353-Figure6-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003212_educon.2010.5492353-Figure6-1.png", + "caption": "Figure 6. Development process followed.", + "texts": [ + " Using the rapid prototyping technologies available in the UPM\u2019s Product Development Laboratory (http://www.dim.etsii.upm.es/ldpdim/) brings students closer to these new technologies now becoming more widespread in industry, thereby giving added value to their training and allowing them to physically check the validity of their CAD designs. Figure 5 shows some of the prototypes made for visual and assembly checks, paying special attention to tolerances, any possible interferences and empty or useless spaces. Additionally Figure 6 includes a schematic diagram of the development process followed, from initial drafts to final pre-production prototype. They also enable certain working trials to be performed, but the fragility of the epoxy resin materials of which they are made has to be taken into account. For tougher trials, a second prototyping stage can be carried out, involving the manufacture of silicone moulds for polyurethane vacuum casting. During course 2008-2009 the costs for manufacturing the 10 prototypes of the selected best toy-designs laid around 2500 to 3000 euros" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002716_robio.2007.4522162-Figure4-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002716_robio.2007.4522162-Figure4-1.png", + "caption": "Fig. 4. Geometrics relationship of navigation", + "texts": [ + " An analytical localization method has been proposed to solve the key problem. The method is to project the three-dimension space relationships of localizing lines into the two-dimension geometric plane and resolve their projecting relationships. And then the twodimension geometric relationships are reduced into the threedimension space to solve localizing lines. It is supposed in the situation without wind and other disturbances that localizing lines can be stable. Based on the perception precision of obstacle sensors as illustrated in Fig.4, the distance between the angle of rotation center and the center of front wheel is 150mm. Thus the angle of rotation on the shift tower can be confirmed, which is defined as by [19] . 2 2 1 1 2 2 1 1 cos cosarctan sin sin R R R R (6) \u2014angle of rotation on shift tower; R1, R2\u2014radii of the navigating handle center line to front and back sensors; L0\u2014distance between angle of rotation center and the center line of navigating handle; 1 \u2014a rotation angle when the navigating handle \u201cidentifies\u201d the lines; ' 1 \u2014error angle of 1 ; 2 \u2014a rotation angle when the navigating handle \u201clocalizes\u201d the lines; ' 2 \u2014error angle of 2 ; \u2014rotation angle of localization compensation" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002347_15397730701404684-Figure4-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002347_15397730701404684-Figure4-1.png", + "caption": "Figure 4. The model of substructure.", + "texts": [ + " It is especially suitable for the cyclic symmetric structure to adopt the substructure technique to make an analysis and meshes only need to be divided for one substructure mode to create the global stiffness matrix, computed results can be taken to other super elements being of the same construction. This gains higher computational efficiency of the substructure mode and leads to a significant reduction of computer cost. The more the repeated structures are, the higher computational efficiency will be achieved. Table 1 gives the overall dimensions of the impeller taken for analysis. The computational model of the impeller is shown in Fig. 4. The mesh of substructure is shown in Fig. 5. Figure 6a shows the side elevation of the finite element mesh of the impeller. Figure 6b shows the front elevation of the finite element mesh of the impeller. In order to truly represent the impeller and improve the accuracy in analysis, the computational model used a four-node tetrahedron solid element with three degrees of freedom at each node. The effects of temperature and pressure loading were not considered, as they are small in comparison with the centrifugal force due to the high speed of rotation (Ramamurti et al", + " 7 and 8, solid curves represent the contact stress distribution of the external surface of shaft sleeve (surface between impeller and shaft sleeve), dashed curves represent the contact stress distribution of the internal surface of shaft sleeve (surface between shaft sleeve and shaft). Contact stresses distribution of both the external and internal surface of shaft sleeve along the axial direction, are illustrated in Fig. 7 ( 1 = 0 03mm, 2 = 0 04mm) with four rotational speeds. The contact stress of right edge of the external surface of shaft sleeve is higher than of other regions because the contact region between impeller and shaft sleeve does not extend to the end of impeller. There is suspended region in the impeller at the place of x = 47 0mm (see Fig. 4), not maintaining contact with shaft sleeve. From Fig. 7, it is clearly seen that a stress concentration is generated due to geometry discontinuity of the root of the suspended region of the impeller. In Fig. 7, it can be seen that for the same amount of interference, contact stress of mating surfaces decreases with the increase of rotational speed due to the increase of centrifugal forces. Variation of the contact stress of mating surfaces edge is small and the change trend of the contact stress of internal and external surface is uniform" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002406_bf00902341-Figure3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002406_bf00902341-Figure3-1.png", + "caption": "Fig. 3", + "texts": [ + " ax ~ - - ax\"- g G3t ~ OX 2 3X 2 H e r e r 0 and q a r e the r a d i u s of g y r a t i o n and the weight p e r unit length of the rope ; g i s the f r e e - f a l l a c e e l e r a - l ion; the plus s ign in f ron t of ~ c o r r e s p o n d s to the a s - r e nd ing b r a n c h of the rope , the minus s ign c o r r e s p o n d s to the descend ing b r a nc h . In w r i t i n g out the equat ions of r ope mot ion in the s e g m e n t s A/i w h e r e t h e r e i s s l i ppage and s l id ing of the rope ove r the pu l l ey [2] we m u s t a l so t ake account of the f r i c t i on f o r c e s :(z) (x, t) dx = tsll p . tslid~X (4) ac t ing on the r o p e e l e m e n t s (Fig. 3). The c o m p o n e n t s f (il)ipdX of the f r i c t i on f o r c e s f (i) (x, t)dx a r e due to the s l i ppage of the rope ove r the pul ley ; : ( i ) , dx a r e due to the s l id ing of the the componen t s ] s l i a r ope on the pu l ley . As we know, the f r i c t i o n a l f o r c e s a r e d i r e c t e d op - pos i t e to the r e l a t i v e ve loc i t y of the moving b o d i e s . Hence, / ~ (x, t) dx = f ~ (x, t) dx cos y~ - sJip f(:) ~(x, t) dx = fr (x, t) dx sin ~, --- slid (d~ (x, t) f(~) (x, t) dx V ~ + r ~'~ rOi (x" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001535_12.666370-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001535_12.666370-Figure1-1.png", + "caption": "Figure 1: Schematic structure of the interdigitated micro-sensor based on hydrazine-sensitive poly (3-hexylthiophene) conducting polymer.", + "texts": [ + " In addition, the effects of hydrazine concentration and P3HT thin film thickness on the sensitivity of the sensor were investigated. The micro-sensor was composed of 3 layers. The upper sensing layer was P3HT conducting polymer thin film, of which electrical conductivity decreased when exposed to hydrazine vapor. The middle layer was interdigitated gold electrode pairs which worked as a transducer to collect and transfer electrical signals. The bottom layer was an oxidized silicon wafer substrate that was used for support and electrical insulation. The schematic of the sensor structure is shown in Figure 1. The design of the interdigitated electrodes was based upon the electrical conductivity of the P3HT thin film. A sensor platform had totally 25 electrode pairs, each electrode finger was 22 microns wide and 2985 microns long, and the spacing between the fingers was 15 microns. Microelectronic fabrication techniques were used to pattern and deposit interdigitated electrodes onto the 4 inch silicon wafer. Then the whole wafer was diced into individual dies. The dimensions of the test sensor platform were 1" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002560_robot.2007.363935-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002560_robot.2007.363935-Figure2-1.png", + "caption": "Fig. 2. Setting for detecting a RDMap", + "texts": [ + " Finally, experiments with a humanoid are conducted to demonstrate obstacle recognition while walking with the proposed method . A. What is a relative disparity map? [11] An RDMap is defined as follows. A range image sensor observes a plane as a reference in advance, and, when it observes a target scene, it detects disparities that are relative to the reference plane. When the range image sensor with multi-spots is used, the RDMap is easily measured as the shifts of the projected spots. This is illustrated in Fig.1. The pose of a sensor for a plane (typically a floor) can be given as shown in Fig.2. The height from the plane is h; the angle between the z-axis (optical axis) of the sensor and the plane is \u03b8; and the rotation angle of the sensor around its z-axis is \u03c6. Supposing the measuring direction (of each spot) is [ s t 1 ]T as shown in Fig.3, the pose parameters of the reference plane are h0, \u03b80, \u03c60, and the disparity of the reference plane for infinite distance is k0(s, t). Then, the 1-4244-0602-1/07/$20.00 \u00a92007 IEEE. 3048 relative disparity \u2206k = k \u2212 k0 is expressed as follows. \u2206k (s, t) = \u03b1 ( cos \u03b8 sin \u03c6 h \u2212 cos \u03b80 sin \u03c60 h0 ) s+ \u03b1 ( cos \u03b8 cos \u03c6 h \u2212 cos \u03b80 cos \u03c60 h0 ) t + \u03b1 ( sin \u03b8 h \u2212 sin \u03b80 h0 ) \u03b1 = b\u00b7f p (1) b: baseline length (distance between the projection center and the lens center) f : focal length of the lens of the CCD camera p: width of each pixel of the CCD This equation shows that the relative disparity is linear for s and t and, therefore, forms a plane in the s\u2212t space, which means that a plane in real 3D space also becomes a plane in the RDMap", + " A Hitachi IP-5005 is used for image processing, such as for the detection of spots and measurement of their center of gravity. Fig.5 shows an example of an experimental scene. Obstacles are put on a flat floor, and the sensor projects spots on the floor. And Fig.6 is an image for Fig.5 obtained by the CCD camera. An RDMap is obtained from such an image. We first show experimental results to detect obstacles by plane fitting. Fig.7 shows the pose parameters of the sensor attached on the humanoid (same as Fig.2). The parameters are changed as shown in Table I by rotating the humanoid\u2019s head. \u03c6 was fixed to 0. Fig.8 shows the RDMaps of a flat floor for the three conditions in Table I. It is shown that the RDMaps remain planar when the sensor pose is changed. On the other hand, Fig.9 shows the RDMap with obstacles of 25(height)\u00d750\u00d725mm3 for condition 2. Fig.9(b) corresponds to Figs.5, 6. One of the obstacles in Fig.9(b) was set so that its height would be 50mm. It is shown that the disparities at the obstacle are out of the plane" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000250_eeic.1991.162566-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000250_eeic.1991.162566-Figure2-1.png", + "caption": "Figure 2. Scanner Relationship to Stator Bore.", + "texts": [ + " Interpretative 39 1 intelligence, normally the domain of skilled personnel, can be provided by a suitably-customized expert monitoring system. The hydro scanner is an in situ system that is sensitive to the various electrical, mechanical, thermal, acoustic, and magnetic signatures that result from the progressive degradation of stator coils and core illustrated by Figure 1. The various sensors are distributed along a structural member, called a bridge, that is situated between two adjacent rotor poles as shown in Figure 2. The length of the sensor-array bridge is sufficient to provide complete coverage of the stator bore from end turn to end turn. In practice, the nose piece of the bridge (containing most of the sensors) is positioned two inches (5 cm) from the stator surface. RFI sensors, however, are mounted away from the bridge at the top and bottom of an adjacent pole on pole centerline. The purpose is to place these sensors as close as possible to the coils having highest induced voltage and therefore maximum probability of exhibiting corona or partial discharging" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001632_pac.2005.1591644-Figure8-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001632_pac.2005.1591644-Figure8-1.png", + "caption": "Figure 8: Pitch-corrected view of EB and pole.", + "texts": [ + " To gain pipe-to-diverging-beam tangency, pipe rotation about the world-x is by the kick angle (\u03b1EB CL) less the EB y-radii divergence angle (\u2206\u03b1). To relate ELS pole coordinate system (POP CS) to the customary perspective for a conventional dipole in the WCS, one must perform a series of two coordinate transformations. Mathematically, the matrix algebra of Fig. 7 is solved for both the divergence-corrected pitch angle (\u03b1LOT) and the roll angle (\u03b3) given constraints on lattice geometry (\u03b2 & \u03b8 of Fig. 2 & 3). Visually, we first rotate the viewing direction in Fig. 5 by the corrected pitch angle (Fig. 8), then by the roll angle (Fig. 9). We may resolve the velocity vector v of an arbitrary particle entering ELS field into components orthogonal to POP; i.e., aligned with the flux lines (POP-y axis, perpendicular to the pole), or lying in the POP (POP x-z). Fig. 10 depicts a particle on the LOT path. The vector v is resolved into components orthogonal to the POP CS. The resultant of vectors parallel to the POP is v // POP. OPTICS AND GEOMETRY 0-7803-8859-3/05/$20.00 c\u00a92005 IEEE 3848 Bend radius due to ELS field is proportional to v // POP ; but the component perpendicular to POP remains constant along the particle path" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000258_peds.1997.618715-Figure4-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000258_peds.1997.618715-Figure4-1.png", + "caption": "Fig. 4 Free body diagram of (a) the spring in relation to the angular surface, and @) the coil bearing", + "texts": [ + " I Calculation of spring effect The movement of the VCM has a nonlinear relationship with the spring position, as shown in figure 3. In this project, the force-position relationship of this spring effect is calculated from the geometry of the VCM and the spring constant. This relationship has been used in the spring compensation block of the controller. The amount of spring force induced on the voice coil is calculated by considering the free body diagram of the wedge and the coil bearing as shown in figure 4(a) and 4@). angular surface \\ \\ \\ Fspring '\\\\\\ // 2.2. Estimation offriction on the small VCM Due to the small size and the geometry of the small VCM, friction becomes a dominant factor. It has been found that friction is a discontinuous nonlinear function which depends on the velocity of the small VCM. To estimate the amount of frictional force on the small VCM, a friction learning program has been developed to run the small VCM in open loop mode. By applying different current profiles to the small VCM, the acceleration, position, and current profiles at various velocities are recorded" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002034_2007-01-2860-Figure12-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002034_2007-01-2860-Figure12-1.png", + "caption": "Figure 12 \u2013 Profile twist movement, generating camber", + "texts": [ + " Figure 11: Positive toe value in the rear wheel Great kinematic variations can be observed in function of crosspiece profile. There is great sensitivity of camber and toe curves due to small alterations in crosspiece profile. To mention some: a) Symmetrical vertical movement - camber Two main variables affect camber in symmetrical movements of the suspension: 1 - Inertia of the crosspiece (Ix): Profiles tend to have smaller camber angles in function of an increase in inertia relative to the profile\u2019s centroid. As in figure 12, camber is generated mainly by twist of the main crosspiece around the X-axis. Therefore, camber must be proportional to the profile inertia relative to the main axis (X). 2 - Stiffness of spring support The spring fixation structure can deform individually, due to high values of the reaction force, thus reducing camber. This fact can indeed modify the trend discussed in item 1, as in Figure 13. b) Symmetrical vertical movement - toe The main cause of toe in symmetrical movements is the relation between the moments of inertia Iz and Ix relative to the profile\u2019s centroid" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002906_s11172-007-0101-5-Figure4-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002906_s11172-007-0101-5-Figure4-1.png", + "caption": "Fig. 4. Comparison of the lower limits of determined concentra tions of phenolic compounds with peroxidases of different origin.", + "texts": [ + "31 The above described established different (either in hibitor or a second substrate) functions of phenols in the horseradish peroxidase catalyzed oxidation of aromatic diamines was verified in the same reactions in the pres ence of other plant peroxidases isolated from peanut cells, the xylotropic fungus Phellinus igniarius, the alfalfa cell culture Medicago sativa, and soya hulls. It was shown that different peroxidases exhibit different sensitivities with respect to the action of these organic compounds (which is of the same type in all cases) (Fig. 4).21\u201424 Different sensitivities of enzymes isolated from differ ent sources with respect to the same inhibitors were also observed for alcohol dehydrogenases. A study of the in fluence of a number of metal ions on the catalytic activity of alcohol dehydrogenases from baker\u00b4s yeast (ADH I) and horse liver (ADH II) in the oxidation of ethanol with nicotinamide adenine dinucleotide (NAD+) has shown that HgII, AgI, CdII, CuII, and ZnII inhibit the catalytic activity of both ADHs in different concentration ranges (Table 5)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000282_tmag.2003.810347-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000282_tmag.2003.810347-Figure1-1.png", + "caption": "Fig. 1. Structural analysis model.", + "texts": [ + " Then, in consideration of the magnetic characteristic of material deteriorating by stress, the claw-poled permanent magnet stepping motor is analyzed using the three\u2013dimensional (3-D) finite-element method with edge elements, which code is made by us. In addition, the influence that it has on the torque characteristic is investigated. First of all, the portion of the stator that stress concentrates is clarified by the structural analysis. Next, the magnetic field of the claw-poled permanent magnet stepping motors is analyzed. The analysis model used in the structural analysis is shown in Fig. 1. The analysis model is for one pole of the stator, and a force as shown in Fig. 1 is applied. The stress distribution of a central section of a tooth is shown in Fig. 2. From Fig. 2, the pulled stress at magnet side and compressed stress at coil side is added. The part of the largest stress is a part of \u201cA\u201d at the coil side as shown in Fig. 2. Therefore, it becomes possible by taking into consideration that the magnetic characteristic at the Manuscript received June 18, 2002. Y. Okada is with Motor Technology Center, Motor Company, Matsushita Electric Industrial Co., Ltd, Osaka 574-0044, Japan (e-mail: yukihiro@mot" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000031_s0307-904x(81)80025-x-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000031_s0307-904x(81)80025-x-Figure2-1.png", + "caption": "Figure 2 shows that the radial deflection u(\u00a2) must be even. Consequently, v(\u00a2) must be odd. Thus:", + "texts": [ + " If we assume the circumferential displacements take the form: f;(r, \u00a2) = v(r - rB)/(r, -- rB) then: l: ere = v/2r(rl/rB -- I)) and Us(V) may now be computed numerically. The coefficient of the total circumferential stiffness is: k~ = k~; + Vs(v)/(.av ~) Lateral coefficient, kw The volume of the axially deformed model can be obtained by solving the couple of problems (B) corresponding to Zn +- w. Let the corresponding volumes be V(w) or 1I(-w), respectively. Equation (2) immediately yields: kw =-p[O.5(V(w) + V(- w))- V(O)]/(Traw 2) Radial loading The planar contact of the model under vertical load is shown in Figure 2. As is well known ~2 the circumferential component of the strain tensor is: e [ ~ V + u ) / r The changes of the belt length are negligible, i.e. e~, ~ 0 for rl ~ r ~< a. For the present we shall relate the symbols u, v to the equator only (r = a). Therefore, u and v become functions of one variable and: u(\u00a2) = --/~(\u00a2) = - dv(~b)/d\u00a2 (4) The energy of the deformed model can be approximated as follows: i - - ~ - - ' f \u00a2 if + - k , , [ ( a : - a - u ) 2 - ( a : - a ) 2 1 a d ~ 2 where the first term on the right hand side denotes the work of the belt tension T, the second term expresses the work of circumferential displacements and the last one is the work of radial displacements. Further, f = - u&) (Figure 2) and a / i s the equator radius of the free carcass, i.e. with removed belt. By means of (4) we obtain U as a functional depending only on the function v and its derivatives: lr U(f) = a f Q(v, i~, i3) de ~t - - 1 T with: Q(v, i), iJ) = t2[T(ii/a)2 + kui) 2 + kvv2l (5) With respect to obvious symmetry it is sufficient to consider the interval 0 ~< \u00a2 < 7r only. In the contact zone 0 ~< \u00a2 ~< 7r the radial displacement is determined by the supporting surface. In the planar case: u ( \u00a2 , ) = - ( a - a ) / c o s \u00a2 - a By means of (4) we can easily obtain v(\u00a2),/~(\u00a2) and//(\u00a2)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000105_s0263574702004630-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000105_s0263574702004630-Figure2-1.png", + "caption": "Fig. 2. Schematic representation of the single link arm dealt with the Ding-Holzer method.", + "texts": [ + " For conservation of elasticity, the stiffness matrix of the field is chosen to be the same as the stiffness matrix of the original link. The mass distribution is approximated by two stations at either end of the link. For mass conservation the masses of the two stations amount to the total mass of the link plus the tip mass, m+M. We also require that the inertia of the system, about the revolute joint, is the same before and after simplification. This implies that the station at the origin must have a mass 2m/3, and the station at the tip will have a mass M+m/3 (see Figure 2). In general, we are only interested in the response of the tip of the arm; the behaviour of the middle of the beam is not too important. This approximate method greatly simplifies the system, yet provides reasonably accurate results. As the arm moves, forces and torques act on the tip and at the joint. The torque provided by the motor will be denoted . The forces and torques can be written as six dimensional vectors, called wrenches. Wrenches can be thought of as element of the space dual to the Lie algebra of the group of rigid body motion, sometimes called the space of co-screws" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001814_1.3547897-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001814_1.3547897-Figure2-1.png", + "caption": "FIG. 2. \u2014 Above: 2-D mesh (210 elements, 257 nodes) [rebars placed at the top of colored elements-cord angles w/r to crown: 64\u00b0/ 16\u00b0/-16\u00b0/-14\u00b0]. Below: 3-D mesh (6301 elements, 7715 nodes) [ABAQUS].", + "texts": [ + " The image resulting from scanning was edited in Adobe Photoshop for a size reduction and then converted to the Data Exchange Format (DXF) through Adobe Streamline (Figure 1). A DXF file of the tire image was imported into GID where a mesh for the tire cross section was created. The node-element information was then imported into ABAQUS where the remainder of the pre-processing was carried out. The definition of the 2-D geometry was completed by identifying elements with reinforcement. Rebars representing the cords and the carcass were placed at the element interfaces. Figure 2 shows the two-dimensional mesh that consisted of 210 4-node hybrid elements with torsion and 257 nodes. The three-dimensional mesh, which was generated by revolving the 2-D mesh around the circumference, consisted of 6301 8-node brick elements and 7715 nodes. Many analysts who perform 3-D simulation of tires assume symmetry at the mid-plane and model only half of the tire. This simplification might play a minor role when the simulation is concerned with only the overall parameters such as structural stiffness and pressure distribution", + " The motivations for this are to prevent convergence difficulties that might occur at the beginning of the steady-state rolling analysis following the static loading, and to allow a gradual change of frictional forces for small load increments in the solution process. The rim was assumed to be fixed since the effect of rim boundary condition on the crack behavior around the belt-edge is not significant. In this section, we present the free rolling analysis of the global tire model as well as that of several local areas with and without cracks. The global-local scheme allows economical yet detailed analysis of critical regions, therefore is very useful for studying crack initiation and crack growth in tires. The analysis of the global model shown in Figure 2 was carried out for loading steps described in the previous section. The steady-state transport analysis requires both the translational ground velocity, v0, and the spinning angular velocity, \u03c9. Since the combination of v0 and \u03c9 that results in free rolling is not known in advance, \u03c9 has to be determined in an indirect manner. Free rolling is defined as the state at which there is no torque applied around the axle. There is no provision in ABAQUS for specifying the value of torque, only the rate of rotation can be input" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001897_bf01974774-Figure5-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001897_bf01974774-Figure5-1.png", + "caption": "Fig. 5. (Tll)R,L VS. shear rate for aqueous ET597 solutions with a capillary diameter of 2.44 mm", + "texts": [ + " 6 % / 0 / 10-3 i i 1111111 i i i i I i I i i i I i i 1 i i i I 0 10 2 10 3 10 4 Apparont Shear Rate (sac- I ) Fig. 3. Flow curves for aqueous ET 597 solutions of various concentrations holds over the range of shear ra tes investigated. I t is to be no ted t h a t the flow curves in fig. 3 are cons t ruc ted wi thou t mak ing end-correct ions, which is justified because of the large capi l lary l eng th- to -d iamete r (L/D) ra t ios ( > 200) employed. P lo ts of the axial no rma l stress, (Vll) R,L, VS. t rue shear r a t e are given in fig. 4 for D : 1.52 m m , and in fig. 5 for D ---- 2 .44mm. Note t h a t these figures are p r epa red b y the use of eq. [ l l ] and figs. 2 and 3. I t is interest ing to see, f rom figs. 4 and 5, tha t , a t a fixed shear rate , (Tll) R,L decreases as solution concent ra t ion is increased. This is a t t r i - bu tab le to the fac t t h a t the magn i tude of the second t e r m on the r igh t -hand side of eq. [11] increases as the concent ra t ion is increased. I n order to clearly observe the dependence of (Tll)R,L on solut ion concentra t ion, cross plots of fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002267_robot.2007.363971-Figure5-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002267_robot.2007.363971-Figure5-1.png", + "caption": "Fig. 5. PBOT model description", + "texts": [], + "surrounding_texts": [ + "PBOT (Fig. 1) is a four-link (foot, shank, thigh, torso) legged robot that has 3 actuators and 8 degrees of freedom forplanarmotion(x, y, Oo, 01, 02, 03, Z1, z2),whereziis a spring force coordinate for ankle and knee actuator. Linear actuation mechanism is used for the ankle and knee joints, which use series elastic actuator[ 12], where a ball screw and spring (6OkN/m) are serially connected with a motor to allow impact shock tolerance. This linear actuator is connected to the robot by forming a closed kinematic chain around the" + ] + }, + { + "image_filename": "designv11_69_0000151_proc-657-ee8.3-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000151_proc-657-ee8.3-Figure1-1.png", + "caption": "Figure 1. TiNi microactuator Figure 2. Setup of fatigue test", + "texts": [ + " The TiNi films are co-sputtered at a deposition rate of ~1.0 \u00b5m/hr, using 75 mm diameter TiNi alloy and elemental Ti targets. Both targets can be controlled independently, which allows for tuning of deposition rate and film composition. The as-deposited TiNi films are amorphous. Shortly after sputtering, in situ annealing is performed at 420\u00b0C for 15 minutes to crystallize the amorphous as-deposited films; Ar is then evacuated from the deposition chamber and the wafer cools in vacuum. The patterned TiNi microactuator shown in Fig. 1 was fabricated using standard micromachining techniques. A TiNi path connects an outer Si frame with a central Si island. The reason to use patterned TiNi films, as opposed to continuous films, is to reduce the thermal mass, while retaining high actuation force. This actuator can be used as the moving part in microvalves or micropumps. In the microvalve described in [5], a microfabricated silicon spring was used to bias the TiNi microactuator such that in the martensite phase, the TiNi was deformed, and the Si island covered the valve orifice" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002381_s11665-007-9185-1-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002381_s11665-007-9185-1-Figure2-1.png", + "caption": "Fig. 2 Schematic representation of a PL/b vs. d plot for a bimaterial system. The slope of the curve corresponds to the bimaterial elastic response m (Eq 12)", + "texts": [ + " Thus, the axial displacement of the examined specimen may be written as, u\u00f0x\u00de \u00bc P b D \u00f0AD B2\u00de x \u00f0Eq 10\u00de where A, B, and D, are defined in Eq 7. Naming u\u00f0x \u00bc L\u00de \u00bc d, the maximum axial displacement is given by, u\u00f0x \u00bc L\u00de \u00bc d \u00bc PL b D \u00f0AD B2\u00de \u00f0Eq 11\u00de It is worthwhile to notice that when E1 \u00bc E2 the coupling stiffness element vanishes (B = 0) and the conventional strength of materials formula d \u00bc PL=Ebh\u00f0E \u00bc E1 \u00bc E2; h \u00bc h1 \u00fe h2\u00de is recovered. For convenience, Eq 11 may be written as, PL b \u00bc \u00f0AD B2\u00de D d \u00bc md \u00f0Eq 12a\u00de with m \u00bc AD B2 D \u00f0Eq 12b\u00de Equation 12a defines a parameter m as the initial slope of the PL/b vs. d curve, see Fig. 2. Thus, m encompasses the stiffness of the bimaterial \u00f0m \u00bc EBim\u00f0h1 \u00fe h2\u00de, where EBim is the modulus of the bimaterial) and may be obtained plotting the load-displacement data of an actual tensile test in the format presented schematically in Fig. 2. The objective here is to use this bimaterial elastic response (m) to extract the unknown Journal of Materials Engineering and Performance Volume 17(4) August 2008\u2014483 elastic modulus of one of the materials, say E2, assuming that E1 and all the geometrical parameters \u00f0L; b; h1; h2\u00de are known. To simplify the calculations, let us define a new variable n \u00bc E2h2 (with E2 as unknown) and the following constants, u1 \u00bc E1h1 \u00f0Eq 13a\u00de u2 \u00bc h1=2 \u00f0Eq 13b\u00de u3 \u00bc E1h2 \u00f0Eq 13c\u00de u4 \u00bc E1 h31 \u00fe 3h22h1 =12 \u00f0Eq 13d\u00de u5 \u00bc h22 \u00fe 3h21 =12 \u00f0Eq 13e\u00de Notice that the parameters u1 through u5 depend only on E1, h1, and h2, which are assumed to be known", + "75% the PR in the bimaterial breaks (perpendicularly to the applied load) and the bimaterial curve drops. A tendency of the curve to recover the PP behavior is observed subsequently. Since the PR in the bimaterial broke at a strain of about 1.75%, tensile testing of the bimaterial was not extended to large deformations and the test was stopped at e 4%. For determination of the bimaterial elastic response (m), the load-displacement curve generated from testing of the bimaterial was plotted in the format sketched in Fig. 2. In order to use the bimaterial model to extract the elastic model of one of the materials, the PP was assumed to act as a substrate with known elastic modulus (454.5 MPa) and the PR was assumed to be the material with \u2018\u2018unknown\u2019\u2019 elastic modulus. With the dimensions of the PP and PR, elastic modulus of PP, and bimaterial elastic response (m) obtained from tensile testing of the bimaterial, the parameters ui ( i \u00bc 1 . . . 8) can be calculated from Eq 13a-e and 16a-c and the elastic modulus of the PR can be computed solving Eq 15" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000118_s0043-1648(03)00330-2-Figure3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000118_s0043-1648(03)00330-2-Figure3-1.png", + "caption": "Fig. 3. Updating the deformations of the bristles (ux), three cases.", + "texts": [ + " The normal pressure p(x) is calculated according to Hertz theory for line contacts. qx(x) = Ktux(x) (14) qx,max(x) = \u00b5p(x) sign(ux(x)) (15) The displacement between the surfaces depends on the motion of the bodies and is calculated step by step for short time intervals. During each step the surfaces move a distance \u03b4x relative to each other. The deformation ux is updated depending on the condition of stick or slip and the direction of \u03b4x (relative to ux) at each point. There are three possible cases (Fig. 3): 1. Stick: ux,n = ux,n\u22121 + \u03b4x if |qx,n| > |\u00b5px,n| then ux,n = ( \u00b5p(x) Kt ) sign(ux,n) 2. Slip and sign(\u03b4x) = \u2212sign(ux,n\u22121): ux,n = ux,n\u22121 + \u03b4x if |qx,n| > |\u00b5px,n| then ux,n = ( \u00b5p(x) Kt ) sign(ux,n) 3. Slip and sign(\u03b4x) = sign(ux,n\u22121): ux,n = ux,n\u22121 Since the contacts are subjected to rolling and sliding motions, new points will enter or exit the contact at every step. Points will enter or exit the contact on either side depending on the direction of the relative velocity between the surfaces. The friction force calculated with the brush model is given by Eq" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001047_ias.2003.1257638-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001047_ias.2003.1257638-Figure1-1.png", + "caption": "Figure 1. Rotor cage equivalent circuit [2].", + "texts": [ + " One particular advantage of this approach is that it is applicable to rotors with both integer and non-integer number of bars per pole. For simplicity, each loop is defined by two adjacent rotor bars and the connecting portions of the endrings between them [1]. For the purpose of analysis, each rotor bar and segment of end ring is replaced by a R-L series equivalent circuit representing the resistive and inductive nature of the cage. It is also convenient to carry out the analysis using mesh currents as the independent variables. Such an equivalent circuit is shown in Fig.1. 0-7803-7883-0/03/$17.00 \u00a9 2003 IEEE Assuming the rotor of a squirrel cage induction machine to be symmetric, an equivalent model of a wound-rotor machine may be obtained in a synchronously rotating dq reference frame [4]: e qse e ds e dssds pirv \u03bb\u03c9\u03bb ... \u2212+= (01) e dse e qs e qssqs pirv \u03bb\u03c9\u03bb ... ++= (02) ( ) e qrre e dr e drrdr pirv \u03bb\u03c9\u03c9\u03bb ...0 \u2212\u2212+== (03) ( ) e drre e qr e qrrqr pirv \u03bb\u03c9\u03c9\u03bb ...0 \u2212++== (04) ( )e qr e ds e dr e qse iiPT \u03bb\u03bb ... 2 . 2 3 \u2212= (05) ( )e dr e dsm e dsls e ds iiLiL ++= ..\u03bb (06) ( )e qr e qsm e qsls e qs iiLiL ++= " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000253_j.1934-6093.2003.tb00099.x-Figure5-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000253_j.1934-6093.2003.tb00099.x-Figure5-1.png", + "caption": "Fig. 5. Control input u(i, j).", + "texts": [ + "9 A B \u2212 = = \u2212 \u2212 , 0.7408 0.3888 0.1641 , 0.0952 0.9048 0.1237 G H \u2212 = = , 6 1 21.5306 10 , 0.01R R\u2212= \u00d7 = , 4 1 24 0.00220.0896 10 , 0.00330.2061 10 M M \u2212 \u2212 \u2212 \u00d7 = = \u2212 \u00d7 , 0.2002 1.2045 0.1641 0 and 0.5026 2.1056 0.1237 0 G H \u2212 \u2212 = = . It is easy to check that the open-loop system is unstable. Finally, referring to ( , )r i j , the state response of the redesigned sampled-data system x(i, j), output y(i, j) and the optimal control u(i, j) are plotted in Fig. 2 to Fig. 5, respectively. Notice that the output y(i, j) follows reference r(i, j) rapidly. Finally, the minimal cost function can be obtained as * 3,3 64.2611J = . The topic of finding the best practical sampling periods for 2-D systems needs further investigation. In this paper, we have utilized the bisection search method to find the sampling periods used in the example. As a result, the selected sampling periods are in the stable range. Whenever, the relative difference between two consecutive cost functions is smaller than some acceptable tolerance error for different sampling intervals, we can regard the digital controller as the acceptable continuous-time controller" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003347_10402001003658318-Figure3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003347_10402001003658318-Figure3-1.png", + "caption": "Fig. 3\u2014Contact wetness within one revolution for four cases: (a) case 1, (b) case 2, (c) case 3, and (d) case 4. (Figure available online in color) (Continued)", + "texts": [ + " Furthermore, oil molecules have a better chance to adsorb on the surface, reforming the protective film if more oil is around. Therefore, how much oil is within the contact regions will somehow impose a limitation on how high a surface temperature the seal can tolerate. In other words, the greater the oil content available around the contact asperities, the higher temperature the asperities can tolerate. This concept is represented by the contact wetness, which is defined as the average partial oil amount per unit area within the contact region. Figure 3 illustrates the contact wetness for the four cases. For each case, three specific instances within one revolution are presented. At a certain instance, distribution of the oil amount is only shown in the contact regions. The oil partial density is then integrated along the circumference at each radial location. And the average value is plotted versus the radius, reflecting the oil amount in the contact region. The radial range of contact is also indicated in the figure. Among the four cases, case 4 (Fig. 3(d)), in which the seal pair failed at 1800 rpm, has the least cavitation effects and largest average partial density, especially in the contact areas. The larger amount of oil content around the asperities helps the seal pair to operate through the higher speed and survive the higher surface temperature. The second best behaved seal pair in terms of contact wetness is case 1 (as shown in Fig. 3(a)). The average partial density is about 0.7\u20130.8 within the contact areas. As indicated in Table 1, this seal pair failed at the lowest speed and the lowest surface temperature. Therefore, there must be another factor(s) to trigger its failure at a much premature stage. Both case 3 and case 2 have relatively small partial densities within the contact regions. And case 3 has a smaller contact area than case 2 over one revolution. The primary reason for the large discrepancies of the contact wetness for the above four cases is the cavitation effects" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001166_1-84628-559-3_7-Figure2.1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001166_1-84628-559-3_7-Figure2.1-1.png", + "caption": "Fig. 2.1. Gear involute [3]", + "texts": [ + " But, due to restrictions in terms of calculation times, computer memory and computer hardware costs, these approaches were based on simple mathematical functions (polynoms, splines, etc.) and a low point density. Today the continuing dramatically change in available computational resources offers new options in gear metrology and quality control for gear manufacturing processes [3- 10]. They include the complete 3D-model of the whole gear, i. e. the gear body (shaft, wheel) and all gear flanks. The basic mathematical equation is a 3-dimensional extension of an involute function, shown in Figure 2.1. A point P is given in cartesian coordinates by cossin sincos ry rx (2-1) with (2-2) taninv r: base radius, : pressure angle, : roll angle, : helix angle The quantity determines the origin of the involute at the base diameter. For helical gears, these intersection points follow a conventional screw. (2-3) r zz zz tan)( )()( 12 12 A pair of one left and right flanks differ by = 2 b, where b is the tooth space half angle at base diameter. Thus, the 3D description of an involute flank without modifications is generated by a conventional involute function \u201cscrewed\u201d along one helix line (Figure 2" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002001_3-540-45118-8_48-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002001_3-540-45118-8_48-Figure1-1.png", + "caption": "Figure 1. Acrobot model", + "texts": [ + " Section 6 gives a description of the experimental setup and describes the implementation of the acceleration control. D. Rus and S. Singh (Eds.): Experimental Robotics VII, LNCIS 271, pp. 481\u2212490, 2001. Springer-Verlag Berlin Heidelberg 2001 The Acrobot is a two DOF planar robot. The first arm is connected to the fixed environment with a rotational joint, joint 1. The second arm is connected to the first arm through a second rotational joint, joint 2. Only joint 2 is actuated which means that the torque of joint 1 is zero. Figure 1 shows the acrobot with the definitions of the joint angles and some basic physical parameters. As the torque of joint 1 is zero the equations of motion considering the torque of joint 2 as input are (1) Where (2) If the angular acceleration is considered as input the equations reduce to (3) d11v\u00b7\u00b71\u2013 d12v2 \u00b7\u00b7 h1 \u03c61+ + + 0= d21v\u00b7\u00b71\u2013 d22v2 \u00b7\u00b7 h2 \u03c62+ + + \u03c4= a1 m1lc1 2 m2 l1 2 lc2 2+( ) I1 I2+ + += a2 m2l1lc2= a3 m2lc2 2 I2+= a4 g m1lc2 m2l1+( )= a5 gm2lc2= d11 a1 2a2 v2( )cos+= d12 d21 a3 a2 v2( )cos+= = h1 a\u2013 2 v2( ) v\u00b72 2 2v\u00b71v\u00b72\u2013( )sin= h2 a2 v2( )v\u00b71 2 sin= \u03c61 a4 v1( ) a5 v1 v2\u2013( )sin+sin= \u03c62 a5 v1 v2\u2013( )sin= d22 a3= v\u00b7\u00b72 d11v\u00b7\u00b71\u2013 d12u h1 \u03c61+ + + 0= v\u00b7\u00b72 u= which is the form used in this study" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001950_1-84628-179-2_5-Figure5.4-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001950_1-84628-179-2_5-Figure5.4-1.png", + "caption": "Fig. 5.4. A tandem helicopter in sideways flight.", + "texts": [ + " Typically, the two rotors are overlapped by around 20% to 50% of the radius (r) of the rotor disk, so the shaft separation is thus around 1.8r to 1.5r. To minimize the aerodynamic interference created by the operation of the rear rotor in the wake of the front, the rear rotor is elevated on a pylon (0.3r to 0.5r above the front rotor). In a tandem rotor helicopter, pitch moment is achieved by differential change of the main rotors thrust magnitude (by collective pitch), roll moment is controlled by lateral thrust tilt using cyclic pitch (Figure 5.4), yaw moment is obtained by differential lateral tilt of the thrust on the two main rotors with cyclic pitch (Figure 5.5), and the vertical force is achieved by the change of the main rotor collective pitch. For simplicity we will present here the dynamic model of a tandem main rotor helicopter in hovering. We propose a dynamic tandem helicopter model based on Newton\u2019s equations of motion [59] with the assumptions of the standard helicopter with the following changes: 1T The nose rotor blades are assumed to rotate in an anti-clockwise direction when viewed from above and the tail rotor blades rotate in a clockwise direction, see Figure 5" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002065_ecc.2007.7068339-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002065_ecc.2007.7068339-Figure1-1.png", + "caption": "Fig. 1. Body-fixed frame Fb and Earth frame Fe.", + "texts": [ + " The rotation of the ship with respect to the satellite will be defined by earth coordinates. To describe the orientation of the antenna element direction, two frames are defined as Joint frame and Plate frame. \u2022 Body-fixed frame Fb is used to show the effect of the wind and wave disturbances on the ship. The origin of this frame is placed in the point where the base of the antenna is located. The vector xb is the heading vector of the ship, yb is the right hand vector of the ship, and zb is pointing downward the ship. Fig. 1 shows the frame Fb on the ship. 368ISBN 978-3-9524173-8-6 \u2022 Earth frame Fe describes the orientation of an Earth fixed frame to the Body-fixed frame Fb. The origin of this frame is placed on the earth where the origin of Fb is located. As long as the ship does not move and the satellite is a geo-stationary one, it is assumed that Fe does not move with respect to Fb. Furthermore, xe is pointing toward the satellite and ye and ze are such vectors perpendicular to xe. They can also be aligned to yb and zb with two rotations of Fe around yb and zb axes. More details are shown in Fig. 1. \u2022 Joint frame F j describes the movement of the azimuth motor with respect to the frame Fb. The origin of F j is placed on the antenna body. The vector z j is always aligned with zb. The vectors x j and y j are initially aligned with the vectors xb and yb respectively, but they change due to the rotation of the azimuth motor which rotate on zb axis. Fig. 2 illustrates this frame on the antenna body. \u2022 Plate frame F p describes the movement of the elevation motor with respect to F j. The origin of F p is placed on the plate of the antenna", + " The rotation matrix Reb describes this rotation and the entries of the matrix are functions of time. This rotation matrix has the property of R\u0307eb = RebSkew(\u03c9b eb), (1) where the entries of \u03c9b eb are the coordinates of the angular velocity vector of Fb relative to Fe, resolved in Fb [10]. \u03c9b eb = [ p q r ]T , (2) where p, q, and r are pitch, roll, and yaw angular velocities respectively. Also, the function Skew is defined as Skew(\u03c9b eb) = \u23a1 \u23a3 0 \u2212r q r 0 \u2212p \u2212q p 0 \u23a4 \u23a6 . (3) These disturbances are illustrated in Fig. 1. \u2022 The only change between Fb and F j is caused by the azimuth motor which is installed on the base. This rotation can be computed directly from the rotation angle of the motor. Consider \u03b8 j b j as the rotation angle of the azimuth motor as shown in Fig. 2. We can represent this rotation matrix by Rb j = \u23a1 \u23a2\u23a3 cos(\u03b8 j b j) \u2212sin(\u03b8 j b j) 0 sin(\u03b8 j b j) cos(\u03b8 j b j) 0 0 0 1 \u23a4 \u23a5\u23a6 . (4) \u2022 The rotation between F j and F p is caused by the elevation motor which is installed on the plate around y j axis" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002400_s12239-008-0037-2-Figure15-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002400_s12239-008-0037-2-Figure15-1.png", + "caption": "Figure 15. Schematic diagram of the driving simulator.", + "texts": [ + " Comparison for a Ordinary Car (1500 kg in mass) In order to compare the optimum value for the front and rear weight distribution ratio of a car with a ordinary weight, maximum-speed cornering simulation was performed using parameters for a vehicle having a mass of 1500 kg. Table 6 shows the results for the lap time in the simulation. It was found that the optimum front and rear weight distribution ratio was between 55:45 and 50:50 for a vehicle having a weight of 1500 kg. 4.1. Description of Driving Simulator Figure 14 shows a photograph of the driving simulator used for this research, while Figure 15 shows a schematic diagram of the simulator. The CarSim simulation model for vehicle was used as the vehicle model for the driving simulator as well as the simulation model. Regarding the simulator, a visual display system and a sound generation system, which simulates the engine sound of a running vehicle, were used to simulate actual driving conditions. A motion device having two degrees of freedom (roll and pitch) was employed. The rolling movement of the motion device is controlled by inputting the roll angle signal" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000749_pime_proc_1973_187_029_02-Figure9-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000749_pime_proc_1973_187_029_02-Figure9-1.png", + "caption": "Fig. 9. Three-inertia system", + "texts": [ + " When checked against the practical result, such as in Fig. 7, the correlation on frequency could not be obtained with the calculated value of stiffness, and was only achieved by using an artificial value of 6 8 4 Nm/rad instead of the 120 Nm/rad calculated. In an indirect gear the difference between the output torque and the input torque is supplied by the reaction of the gearbox casing. With an engine and gearbox assembly on rubber mountings this whole inertia must participate in the vibration. This system is represented in Fig. 9 as a branched system with three inertias. This simulation gave good agreement on the natural frequency and proved to be adequate for examining many of the possible or postulated causes of judder, but the representation of damping in the system proved to be unsatisfactory. Linear behaviour had been obtained by assuming damping proportional to velocity-and a sufficient amount had to be associated with the vehicle inertia I,. This damping appeared then as a drag to be overcome by the engine, which would have given the car a maximum speed of only 32 km/h" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001118_0471758159.ch5-Figure5.31-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001118_0471758159.ch5-Figure5.31-1.png", + "caption": "FIGURE 5.31 A ferrite bead.", + "texts": [ + " Ferrite materials are basically nonconductive ceramic materials that differ from other ferromagnetic materials such as iron in that they have low eddy-current losses at frequencies up to hundreds of megahertz. Thus they can be used to provide selective attenuation of high-frequency signals that we may wish to suppress from the standpoint of EMC and not effect the more important lower-frequency components of the functional signal. These materials are available in various forms. The most common form is a bead as shown in Fig. 5.31. The ferrite material is formed around a wire, so that the device resembles an ordinary resistor (a black one without bands). It can be inserted in series with a wire or land and provide a high-frequency impedance in that conductor. The current passing along the wire produces magnetic flux in the circumferential direction, as we observed previously. This flux passes through the bead material, producing an internal inductance in much the same way as for a wire considered in Section 5.1.1. Thus the inductance is proportional to the permeability of the bead material: Lbead \u00bc m0mrK, where K is some constant depending on the bead dimensions" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003212_educon.2010.5492353-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003212_educon.2010.5492353-Figure2-1.png", + "caption": "Figure 2. Different examples of conceptual designs.", + "texts": [ + " B) Conceptual design The teams continue to work on the list of requirements to identify any crucial problems and choose the best solution for each one, paying also attention to manufacturability, time optimization and costs reduction. Using CAD programs and drawing sketches by hand (as shown bellow) sees the beginning of the work to obtain a pre-design of the different parts while comparing any possible alternatives. In this way, materials are chosen according to the initial estimations of resistance needed for the different components and parts. Figure 2 shows as example the conceptual design. C) Detailed design Once the most appropriate solution has been chosen from the different pre-designs, the different parts must be exactly defined. Following the concepts explained in the theory classes, the students must use a design approach oriented towards manufacture and assembly, in line with the current trends in Concurrent Engineering. The results of some different tasks are shown in Figure 3. To check that the chosen materials are suitable, the estimates need to be compared, using simplified theoretical models, with the information provided by Computer Aided Engineering programs" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000429_6.2003-5652-FigureA.1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000429_6.2003-5652-FigureA.1-1.png", + "caption": "Fig. A.1: Dual control missile", + "texts": [ + " APPENDIX The generic interceptor model used in this study is based on Freidland's tail control missile example introduced in [10]. The model assumes planar and linear dynamics, with no thrust or drag force and constant speed. Furthermore, it relies on a small angles assumption. These assumptions are reasonable during the end-game period, the time of interest in this study, which is typically small for high performance interceptors. The governing equations of the missile configuration are modified in order to incorporate additional canard controls (see Fig. A.1): t t c c V Z V Z q V Z \u03b4\u03b4\u03b1\u03b1 \u03b4\u03b4\u03b1 +++=> (A.1) ttccq MMqMMq \u03b4\u03b4\u03b1 \u03b4\u03b4\u03b1 +++=> (A.2) The states \u03b1 and q are the angle of attack and the pitch rate, respectively. tc ZZZ \u03b4\u03b4\u03b1 ,, are the linear force coefficients of the angle of attack and the canard and tail angles c\u03b4 and t\u03b4 . Similarly tc MMM \u03b4\u03b4\u03b1 ,, are the linear moment coefficients about the center of gravity CG. qM is the aerodynamic damping coefficient and V is the missile speed. In this representation the forces and moments are normalized by the missile mass and inertia respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001293_rob.20099-Figure4-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001293_rob.20099-Figure4-1.png", + "caption": "Figure 4. Step 1 of guidance example: a step 1.1 and b step 1.2, respectively.", + "texts": [ + " After the long-range motion of the platform, the guidance algorithm first relocates the platform so that at least one laser beam hits the center of a corresponding detector as close as random errors would allow . The subsequent corrective actions move the platform in such a way as keeping the first LOS hitting the center of its respective detector, while repositioning the platform such that the remaining LOS also hit their targets. The overall proofof-concept algorithm, thus, consists of three steps: Step 1 Figure 4 : 1.1 Determine the largest PSD offset, d1. 1.2 Translate the platform along the x or y direction by d1. Step 2 Figure 5 : 2.1 Measure the new offset, d2, along the same detector considered in step 1. Calculate the actual orientation angle of the platform from the x axis, a, and determine the difference = d\u2212 a, where d is the desired platform orientation. 2.2 Translate, by b, and, then, rotate, by , the platform about the center of the fixed side of the platform to compensate for d2 and . This movement should result in the LOS considered in step 1 to keep hitting the center of its detector and, furthermore, the platform achieving its final desired orientation" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002265_12.776490-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002265_12.776490-Figure1-1.png", + "caption": "Figure 1. Configuration of spherical motor.", + "texts": [ + " The resulting motor has a simple configuration, including the holding mechanism via a permanent magnet, and is appropriate for miniaturization; it is driven by a low voltage of less than 3 V. * Corresponding author, E-mail: ueno@aml.t.u-tokyo.ac.jp Behavior and Mechanics of Multifunctional and Composite Materials 2008, edited by Marcelo J. Dapino, Zoubeida Ounaies, Proc. of SPIE Vol. 6929, 69291V, (2008) 0277-786X/08/$18 \u00b7 doi: 10.1117/12.776490 Proc. of SPIE Vol. 6929 69291V-1 Downloaded From: http://proceedings.spiedigitallibrary.org/ on 07/03/2016 Terms of Use: http://spiedigitallibrary.org/ss/TermsOfUse.aspx 2. CONFIGURATION AND PRINCIPLE Figure 1 shows the configuration, and Figure 2 shows a photograph of our motor. It consists of four square rods of Galfenol (Fe81.6Ga18.4), with wound coil, permanent magnet, magnetic fixture, and spherical rotor (iron ball) located on the edge of the rods. The rods, having magnetically preferred axis in the long direction, are bonded on the fixture with an interval of 90\u25e6 and are subject to strain by a magnetic field induced by the coils. The permanent magnet placed at the center of the fixture provides bias magnetic fluxes passing through the rotor, with four rods attached as shown in the leftward diagram in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002200_09544054jem699-Figure6-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002200_09544054jem699-Figure6-1.png", + "caption": "Fig. 6 The first scheme for the SSM system", + "texts": [ + " FR1: TP1 X guide and screw: distance, 420mm; accuracy, 0.01mm FR2: TP2 Y guide and screw: distance, 530mm; accuracy, 0.01mm FR3: TP3 Z guide and screw: distance, 550mm; accuracy, 0.01mm FR4: TP4 heater and temperature controller: maximum temperature, 300 C\u030a, accuracy, 1 C\u030a FR5: TP5 pressing guide and belt: distance, 700mm; accuracy, 0.06mm FR6: TP6 rewind drawing: distance, 550mm; accuracy, 0.01mm FR7: TP7 CO2 laser and its power controller: maximum power, 40 W; accuracy, 1 W The first design scheme is shown in Fig. 6. In this SSM system design, all the TPs are controlled. From equation (9), the desired value of control information in the SSM machine is M \u00bc X10 i\u00bc1 log2 Si \"i \u00bc 84:83 Therefore the corresponding control efficiency for the first design scheme is obtained as h \u00bc IR M \u00bc 112:29 84:83 \u00bc 1:324 In this design scheme, TP2 and TP5 are two Y-axis motion elements. TP2 is used to accomplish X\u2013Y JEM699 IMechE 2007 Proc. IMechE Vol. 221 Part B: J. Engineering Manufacture at UNIV OF CONNECTICUT on June 27, 2015pib" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000933_acc.2004.1384489-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000933_acc.2004.1384489-Figure2-1.png", + "caption": "FIGURE 2. Photographs of the fabricated water dielectric patch antenna. (a) Whole structure, (b) Side view, (c) Antenna under test.", + "texts": [ + " Therefore, it can be seen from the above analysis that for any requirement of practical applications, a proper height of substrate may be selected to achieve the desirable bandwidth. III. RESULTS AND DISCUSSION In order to demonstrate the correctness of the proposed design, a prototype of design I in Section II was fabricated and measured. The input characteristic of the antenna was measured by applying anAgilent Network Analyzer E5701C, while the radiation performance was measured by a Satimo Starlab near-field measurement system. A. WATER DIELECTRIC PATCH ANTENNA WITH METALLIC GROUND PLANE The fabricated prototype is depicted in Fig. 2. A water valve made of plastic is installed at a corner of the small plexiglass box for water injection. From the side view of the prototype shown in Fig. 2(b), it is seen that the whole antenna structure maintains a low profile. The height of the whole antenna including the thickness of the plexiglass box is 19 mm, corresponding to 0.059 \u03bb0 at 0.93 GHz. The antenna under test is illustrated in Fig. 2(c). In our design, the xoz and yoz planes are the E- and H- planes respectively. The simulated electric field distributions across the antenna are shown in Fig. 3. It can be seen that the electric field in the air substrate between the water patch and the ground plane is much stronger than the field in the water dielectric patch. The field is radiated from the two open ends of the water dielectric patch. Therefore, it is confirmed that the proposed water dielectric patch design works as a patch antenna instead of a DR antenna" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003130_978-90-481-9884-9_25-Figure3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003130_978-90-481-9884-9_25-Figure3-1.png", + "caption": "Fig. 3 Position of the centre of mass of the piston i", + "texts": [ + " (50) When (50) is pre-multiplied by JT Di the kinetic component of the generalized force applied to {P} due to each cylinder rotation is obtained in frame {B}: P fCi(kin)(rot)|B = JT Di .Ci fCi(kin)(rot)|B = JT Di . d dt ( ICi(rot)|B .JDi .T ) .B x\u0307P |B|E + JT Di .ICi(rot)|B .JDi .T.B x\u0308P |B|E . (51) The inertia matrix and the Coriolis and centripetal terms matrix of the rotating cylinder may be written as: JT Di .ICi(rot)|B .JDi .T, (52) JT Di . d dt ( ICi(rot)|B .JDi .T ) . (53) If the centre of mass of each piston is located at a constant distance, bS , from the piston to moving platform connecting point, Pi (Fig. 3), then its position relative to frame {B} is: BpSi |B = \u2212bS.l\u0302i +B pi|B +B xP(pos)|B . (54) The linear velocity of the piston centre of mass, B p\u0307Si |B , relative to {B} and expressed in the same frame, may be computed as: B p\u0307Si |B = l\u0307i +B \u03c9li |B \u00d7 (\u2212bS.l\u0302i ) , (55) B p\u0307Si |B = JGi . [ BvP |B B\u03c9P |B ] , (56) where the Jacobian JGi is given by: JGi = [ I \u2212 bS. \u02dc\u0302lTi .\u02dc\u0304li ( I \u2212 bS. \u02dc\u0302lTi .\u02dc\u0304li ) .P p\u0303T i|B ] . (57) The linear momentum of each piston, QSi |B , can be represented in frame {B} as: QSi |B = mS" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003398_ijmr.2009.026577-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003398_ijmr.2009.026577-Figure2-1.png", + "caption": "Figure 2 Problem domain", + "texts": [ + " The second part of the numerical model uses the values obtained from the first part to determine the temperature distribution in the tool insert by means of the finite difference techniques. The friction-heating rate per unit area of the chip-tool contact is given by equation (1) (Childs et al., 2000): fric chipfq V\u03c4= (1) where fric:\u03c4 Friction stress (MPa) chip:V Chip velocity (m/min). It is assumed that some heat fraction *\u03b1 will flow into the chip and that the remaining fraction (1 *)\u03b1\u2212 will flow into the tool. The fraction *\u03b1 and friction-heating rate qf is assumed to remain constant over the contact length l or OB (Figure 2). The question, which arises is what quantity of heat flows into the chip and into the tool. It must be noted that the contact area is common to the tool and the chip. and therefore the temperature should be the same whether calculated for the tool or the chip (Childs et al., 2000). There is heat flow into the tool over a rectangle fixed on the surface of a semi-infinite solid. The rectangular contact area is said to have a depth d (depth of cut) and length l. After considering the mean temperature rise over a rectangular heat source fixed on the surface of a semi-infinite solid, Childs et al", + " (2000) derived the following equation for the temperature rise in the tool: avg.tool avg chip 0 contact (1 *) ( ) f t V l T T s K \u03b1 \u03c4\u2212 \u2212 = (2) where To: Ambient temperature (\u00b0C) Kt: Thermal conductivety of the tool (W/m\u00b0C) Sf: Shape factor l: Chip-tool content length (m). The shape factor is said to be dependant on the contact area aspect ratio (d/l). The aspect ratio is said to increase from 0.94 to 1.82 as (d/l) increases from 1 to 5 (Childs et al., 2000). It is clear that the chip moves past the heat source at the speed chipV as shown in Figure 2. The temperature rise in the chip is therefore based on the theory of a moving heat source. The thermal diffusivity \u03ba of a given material is said to be given by the following equation (Childs et al., 2000); 20.00017m / min w w w K C \u03ba \u03c1 \u03ba = = (3) where Kw: Workpiece thermal conductivity (W/m\u00b0C) \u03c1: Density (kg/m3) C: Specific heat (J/g\u00b0C). The Peclet number Pe is known to be a dimensionless number concerned with the rate of advection of a flow, to the rate of diffusion (Holman, 2002). Childs et al" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003622_s0026-0657(09)70201-2-Figure3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003622_s0026-0657(09)70201-2-Figure3-1.png", + "caption": "Figure 3. Golf ball blow mould. Left: Conformal cooling. Middle: Venting channels (Green). Right: DirectTool cavity. Courtesy: Es-Tec, DemoCenter.", + "texts": [], + "surrounding_texts": [ + "8 MPR September 2009 0026-0657/09 \u00a92009 Elsevier Ltd. All rights reserved.\nin business\nChannelling quality for moulded parts using fast manufacturing Direct metal laser-sintering (DMLS) technology may sound like something out of science fiction. But as its manufacturers and users report, it\u2019s gaining ground among customers as a way to produce complex parts quickly and accurately. Some developments have clear economic advantages, says Liz Nickels...\nDirect metal laser sintering (DMLS) produces solid metal parts by locally melting and solidifying metal powder with a focused laser beam, layer by layer. A 3D computer-aided design model is \u201csliced\u201d into layers, and the lasersintering technology then builds the geometry. The technology\u2019s Munichbased manufacturer, EOS, says that it can create extremely complex metal parts in a relatively short cycle time and is particularly suitable for industries that no longer need to produce a large volume of identical parts.\nPotential applications cited by EOS include tooling, aerospace and automotive industries, designer objects, con-\nsumer goods such as toys, and robotics. New applications are being found in the medical field \u2013 dental prostheses, implants and devices.\nAccording to the EOS website: \u201cLaser-sintering enables a rethinking in product development and production. It is a departure from tool-based, inflexible technologies in favour of generative, flexible methods.\u201d Claudia Jordan, an EOS spokesperson, said: \u201cThe most groundbreaking aspect of the technology, in our opinion, is freedom of design. This makes it possible to produce really attractive products, such as customised implants and lightweight structures for aerospace and professional cycling.\u201d\nBut is the technology \u2013 also called e-manufacturing by the company \u2013 an innovation too far for customers in the current economic climate, where innovation can take second place to the stability of \u201ctraditional\u201d technology?\nOne user of the technology is LBC Laserbearbeitungscenter GmbH, a manufacturer of inserts, parts and prototypes for diecasting and injection moulding. Currently, 20% of its work involves producing prototypes for aircraft, the medical industry or for the automotive sector. It was established in October 2002 and only two years later started using EOS\u2019s DMLS technology.\nIn the injection moulding tooling business, there is hardnosed busi-", + "September 2009 MPR 9metal-powder.net\nness reasoning behind the deployment of innovative technology. The DMLS system allows LBC to build moulds that include internal conformal cooling channels. These are carefully calculated and designed to reduce the time taken to cool the mould in each moulding cycle, making for faster throughput and improving productivity and, therefore, profitability.\n\u201cWe\u2019re very pleased with the technology. It is excellent for mass production moulds,\u201d said Ralph Mayer, managing director of LBC. \u201cWith conformal cooling we can achieve cycle time reduction up to 60%, although the average time reduction is 30% to 40%.\u201d\nConformal cooling is, in fact, a fairly recent commercial reality. \u201cIn theory, EOS has been dealing with the concept of conformal cooling for over 10 years,\u201d said Jordan. \u201cBut it was only when we started to implement MS1 (a maraging steel in fine powder form) at the end of 2006 that EOS could actively offer conformal cooling.\u201d\nDMLS allows for almost any shape in heating/cooling channels and thus can already improve the effectiveness of\ncooling. However, by using conformal cooling with DMLS, routing options for cooling channels are almost infinite.\nEOS goes so far as to claim that certain geometries of products can only be manufactured at required quality standards with conformal cooling. \u201cIt is suitable for making very complicated parts to a very high quality in a much faster time than traditional moulding,\u201d agrees Mayer. \u201cFor example, a gear wheel requires special milling tools to be created, and can take up to four months using existing technology. Using DMLS, it is possible to make a prototype in just one week.\u201d\nBenefits of conformal cooling\nAccording to an EOS white paper, conformal cooling works by creating a suitable cooling channel at a well defined distance to the cavity, which is impossible using a conventional drilled cooling mechanism (Figure 1). Cooling channel cross sections can take almost any shape. Turbulence of the coolant (the desired high Reynolds number)\nwithin the system can thus be controlled by actively choosing different cross sections and by switching between different cross sections. As a consequence, turbulence inside the coolant stream is generated close to the mould cavity along the whole path of the channels.\nChanging the cross sections or forking the cooling channel can be done without splitting up the form. This allows for additional heat/cooling advantages in areas that cannot be reached by conventional methods.\nConformal cooling can also improve quality due to better control of the injection moulding process. Warping and sink marks are minimised by evenly cooling out the plastic melt, thus minimising internal stress. Scrap rates are reduced or eliminated. Avoiding internal stresses helps to produce better parts with the same amount of required material. In fact, certain geometries are only possible to manufacture at required quality standards with conformal cooling, EOS says.\nCombined systems with separated cooling and heating channels are also possible, and the technology can also", + "10 MPR September 2009 metal-powder.net\nperform a split between main systems (for the control of the global temperature) and specific systems (for the handling of close to cavity critical temperatures), opening up the potential for future applications. Heating/cooling at critical parts inside the tool, which often cannot be reached by conventional methods, becomes feasible (eg long and lean cores, areas around hot-runners or small sliders). Using special copper heat conductors or other complex measures becomes obsolete. If necessary, it is also possible to undercool mould cavities, thus reaching optimal cycle times by minimising cool down times in tooling cavities.\nAn evened out temperature level can help to improve tool life time. This becomes relevant especially in die casting tools that are exposed to extreme temperature variations.\nIn the white paper, EOS also sets out the drawbacks of conventional cooling. The distance from cavity to cooling channel differs, as only straight line drilling channels are possible and as a consequence heat dissipation cannot take place uniformly in the material.\nThis can result in uneven temperature levels on the cavity surface, uneven cooling-down processes resulting in internal stresses and thus negative impact on part quality (warpage). The drilling procedure is not without risk: in the case of deep drilling there is always a danger of hitting ejector holes (wandering drill), or the drill can even break. As a consequence, the whole mould insert could become unusable.\nIn conventional cooling, optimising the cooling channels helps condition the tooling temperature, enabling uniformity. This temperature level can be influenced to achieve on the one hand a lower temperature for quicker cooling, or higher temperature for better product surface quality on the other.\nConventional cooling channels are drilled into a tool. This limits design to straight lines, easily accessible by a drill. Tooling cavities can pose limits to position and routing of conventional cooling channels. With DMLS, however, the cooling channels can be positioned freely. The cross sections can be optimised to mould temperature control requirements.\nFuture developments Some of the results already obtained by using DMLS to create injection moulds and mould inserts with conformal cooling include a 20% increase in mould productivity and a 50-hour toolmaking time for a blow-mould (Es-Tec), reduced cycle times from 15 to 9 seconds, enabling a 75% increase in productivity on a four-bottle blow mould with DMLS inserts (SIG Blowtec) and reduced cycle time by two-thirds using a DMLS designed core effectively cooling down a critical hot spot (LaserBearbeitungsCenter).\nEOS says that DMLS can make electric discharge machining (EDM) and milling obsolete in many cases, especially with part geometries where slides, inserts or other tool components with complex characteristics are required. But one issue with the technology is that building rate is still a problem, according to Mayer. \u201cCurrently it is much faster to build components in large numbers using traditional moulding technology. At the moment, using DMLS, we can only produce 8cm3 per hour. However it seems likely that this will improve in the future.\u201d" + ] + }, + { + "image_filename": "designv11_69_0003807_amr.139-141.1079-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003807_amr.139-141.1079-Figure1-1.png", + "caption": "Fig. 1 Tapered roller bearing in section Fig. 2 The geometry of Direct busbar Tapered Roller parameters", + "texts": [ + " With the Finite Element Method(FEM), by analyzing the stress distribution of roller busbar and raceway contact area in the design of different rollers with varied convexity of tapered roller bearing, the paper gets the best solution for the design of convexity of tapered rollers and the cause of roller bearings\u2019 early destroy. The optimal result shows that the service life of this bearing has been improved by 93%. Hence, a more efficient method of improving the service life of rollers is got. Taking advantage of its structural characteristic, tapered roller bearing can take heavy load, and it is well used in the equipments for heavy load, hoist, and metallurgy. As in Fig.1, when Tapered roller bearing is working under certain conditions, the damage to the bearing is mainly given by the contact fatigue on the rolling surface, which is caused by the circulating contact stress. The crucial problem in the bearing design is how to optimize the contact stress distribution between roller and raceway and then improve the fatigue life of the kinematics pair between roller and raceway to avoid early damage to the bearing. In this paper, we are going to make finite element method analysis on the problem of contact of tapered roller bearing by using finite element software, ANSYS" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003381_00022660910983716-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003381_00022660910983716-Figure1-1.png", + "caption": "Figure 1 Schematic diagram of the ALV", + "texts": [ + ", 2007), an algorithm has been presented that its efficiency for identifying the vehicle parameters has been evaluated by flight simulator data. The algorithm is performed with the aid of autoregressive with exogenous input (ARX) model. The ARX model is especially realistic and matches the structure of many real-world processes (Nelles, 2001). Time-variant parameters are tracked by Kalman filter method. In this paper, this algorithm is used for implementation of system identification based on flight test data. In this paper, pitch (longitudinal) channel of the vehicle, as shown in Figure 1, is selected for study due to the fact that the major guidance commands for maneuvers are given in longitudinal plane (Mehrabian et al., 2006). Identified parameters are finally compared with the linear model parameters extracted by two other methods that are: 1 identification based on the recorded data from six-degree- freedom (six DoF) simulation of motion; and 2 linearization of the equations of motion via small- disturbance theory as an analytical method. The current issue and full text archive of this journal is available at www" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000781_bf02644033-Figure20-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000781_bf02644033-Figure20-1.png", + "caption": "Fig. 20-Theoretical deformation behavior in t ransverse compression of (001) monocrystals as a function of orientation of the s t ress axis (a) under no constraint, (b) under the constraint of no component of shear strain on the specimen and (c) under the constraint of plane-strain deformation.", + "texts": [], + "surrounding_texts": [ + "F o r p lane s t r a i n de fo rma t ion , c a s e (c), the l a t e r a l s t r a i n r a t i o dczz/deyy ~ oo for a l l ,I,. Since c a s e (c) has the h ighes t d e g r e e of c o n s t r a i n t among the t h r e e c a s e s , i t i s to be expec ted tha t the T a y l o r f ac to r M or the y ie ld s t r e n g t h should be equal to o r g r e a t e r than those of the o ther two c a s e s . This i s seen to be so f r o m Fig . 20~ The M v s 4, cu rve for c a se (c) l i e s above those of (a) and (b) excep t at 9 = 0 deg and 45 deg. At t hese o r i en ta t ions the p l a n e - s t r a i n condi t ion can be s a t i s f i e d without add i t iona l shear~ To c o m p a r e the e f fec t s of s u p p r e s s i n g the s h e a r component cl~xy METALLURGICAL TRANSACTIONS A vs tha t of s u p p r e s s i n g the l a t e r a l component deyy on M, let us examine cases (a) and (c) in greater detail. As 4, a p p r o a c h e s z e r o , d~zz/d~yy for c a s e (a) a p - p r o a c h e s that of c a s e (c); the d i f f e r ence in M be tween t h e s e two c a s e s i s the s m a l l e s t , a l though dexy to be s u p p r e s s e d in c a s e (c) is the l a r g e s t . As 4, i n c r e a s e s the d i f f e r ence in dezz/deyy between c a s e s (a) and (c) i n c r e a s e s ; whi le the amount of d~xy to be s u p p r e s s e d in ca se (c) d e c r e a s e s , the d i f f e rence in M i n c r e a s e s . At 4, = 22.5 deg the d i f f e r ence in d~zz/deyy between c a s e s (a) and (c) i s the l a r g e s t , al though no s u p p r e s s i o n of d~xy in c a s e (c) i s r e q u i r e d , the d i f f e r ence in M is a l s o the l a r g e s t . T h e r e f o r e , l a r g e r M in p l a n e - s t r a i n d e f o r m a t i o n i s a s s o c i a t e d with the add i t iona l s h e a r r e q u i r e d to s u p p r e s s the l a t e r a l t e n s i l e s t r a i n c o m p o n - ent dcyy r a t h e r than the s h e a r component d~xy~ CONC LUSION The low degree of plastic anisotropy, observed in the transverse compression of DS nickel-base superalloys, expressed as the lateral strain d~zz/dEyy, can be explained qualitatively according to the HosfordBackofen analysis in terms of the scattering in growth directions of the columnar grains and nonrandom rotations of the grains about their growth directions. However, quantitative agreement with the Hosford-Backofen analysis has not been achieved, possibly due to insuf- VOLUME 7A, JANUARY 1976-t5 ficient texture characterization on the columnar grained nickel-base superalloy used in this study. Furthermore deviation of the growth directions from (001) increases the yield strength (or average M) of the DS structure, The yield strength of nickel-base superalloy monocrystals under plane-strain deformation at high homologous temperature was found to be in good agreement with the Bishop-Hilt analysis. Therefore, it indirectly confirms the applicability of the Hosford-Backofen analysis used for the study of the DS nickel-base superalloy since it is essentially an extension of the BishopHill analysis for textured materials. ACKNOWLEDGMENTS The authors wish to thank Mr. R. E. Doiron for specimen preparation and testing and Mr. A. V, Karg for X-ray analysis~ Enlightening discussions with Dr, G. Y. Chin of the Bell Telephone Laboratories and Pro- lessor W. F. Hosford, J r . of the University of Michigan are greatly appreciated. REFERENCES t. F. L VerSnyder, R, B, Barrow, L. W. Sink, and B. L Pieareeu M o t Ca~L, t967, voL 52, p, 68, 2, B. J. Pieareey, B. H. Kear, and R, W, Smashey: Trans. ASM, 1967, voL 60, p, 635, 3. C. H. Wells: Trans, ASM, I967, vol. 60, p. 270, 4. W, Fo Hofford, Jr. and W. A. Baekofen: Fundamentals o f Deformation Processing+ W. A. Backofen, J. J. Burke, L. F. Coffin, N. L. Reed, and V, Weiss, eds. pp, 25902, Syracuse University Press, 1964. 5. J. F. W, Bishop and R, Hill:Phil. Mag.. 1951,vot, 42, p. 414, 6. J, F~ W. Bishop and R. Hill: Phil. Mag., t951, rot, 42, p. 1298. 7, F. L, VerSnyder and M. E. S|lank: Mater. Sei. Eng,, t970, rot, 6, p, 213, 8, D, Lee and W. A. Backofen: Trans. TMS-AIME, 1966, vol, 236, p, 107% 9, J. F, W. Bishop: Phil. Mag., 1953, voL 41, p. 51. tO, W, F. Hosford, Jr.: Trans~ ~MS~d[M~. 1965~ voL 233. p. 329. i I. G, Mayer ~nd W. A. Backofen: Trans. TM.S~4LrCfE. I968, voL 242, p. I587, 12, W, F. Hosford, Jr.: AetaMet. 1966, vo|. t4, p. I085, 13. G, Y, Chin, E, A, Nesbitt, and A, J. Williams: Aeta Met. t966, rot, 14, p, 467, t4. B. C. Wonsiewicz andG. Y, Chin: Met. Trans., 1970, vot. 1, p, 2715, 15~ P. R. Morris and A. 3. Heckler: Advances in X-rey Analysis. voL 11, pp. 454-72, Plenum Press, N.Y., t968. 16-VOLUME 7A, JANUARY 1976 METALLURGICAL TRANSACTIONS A" + ] + }, + { + "image_filename": "designv11_69_0001574_s00422-005-0017-9-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001574_s00422-005-0017-9-Figure1-1.png", + "caption": "Fig. 1 A biomechanical view of an example of neutral state in the bipedal body. When the leg begins the stance phase with zero speed of the body\u2019s center of gravity, the forward or backward falling motion occurs depending on whether the knee joint angle \u03c6 exceeds \u03c60 (the neut ral state), or not", + "texts": [ + " In human bipedal walking, an extension of the leg depends roughly on the angle of the knee joint only. In this study, we will focus on the relationship between the modulation of the knee joint angle and the adaptive change of a walking pattern. In the following section, we consider the extreme case in which ankle torque cannot be fully generated in the ankle, but normally in all other joints. In this case, only a passive falling motion of the leg in stance phase, which is the essential motion as an inverted pendulum, can be used in the generation of the forward propulsive force.As shown in Fig. 1, when the leg begins the stance phase with zero speed of the body\u2019s center of gravity, the direction of the body falling motion (forward or backward) is determined by the knee joint angle. It is obvious that in the BSP, there is a neutral balanced angle (\u03c60 = neutral state) at the knee joint such that it determines whether there will be a fall or sustained walking. Here, we call the knee joint angle in the BSP the initial state (\u03c6). If \u03c6 is larger than the neutral angle \u03c60, the body begins to fall forwards" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003210_j.mechmachtheory.2009.09.007-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003210_j.mechmachtheory.2009.09.007-Figure2-1.png", + "caption": "Fig. 2. An arc of a circle described by Ap .", + "texts": [ + "6b) of Appendix B allows to determine function Up. 3.1. Some geometrical lemmas on the Jacobian of Up Some geometrical lemmas on the Jacobian of Up will now be given, aimed to define a pure geometrical interpretation of results of Appendix A. Lemma 3.1. Let k 2 f1; . . . ; pg; \u00f0hj\u00dej2f1;...;pgnfkg, be p 1 fixed angles and eU a function from R to R2 defined by eU\u00f0h\u00de \u00bc Up \u00f0h1; . . . ; hk 1; h; hk\u00fe1; . . . ; hp\u00de: \u00f05\u00de Then, the range of deU\u00f0hk\u00de is orthogonal to \u00f0Ak 1\u00f0h1; . . . ; hk 1\u00de;Ap\u00f0h1; . . . ; hp\u00de\u00de. Proof. If h is varying (Fig. 2), point Ap \u00bc Up\u00f0h1; . . . ; hk 1; h; hk\u00fe1; . . . ; hp\u00de describes a circle of center Ak 1, crossing by Ap. This curve admits a tangent orthogonal to \u00f0Ak 1\u00f0h1; . . . ; hk 1\u00de;Ap\u00f0h1; . . . ; hp\u00de\u00de. Moreover, the tangent to the curve is also parallel to the range of deU\u00f0hk\u00de, which permits to conclude. h Referring to Definition 2.1, we will present a consequence of this lemma fundamental for the geometrical interpretation of results of Appendix A. For all x 2 F, we denote I \u00bc fi1; . . . ; iqg the set of integers 1 6 i1 < i2 < < iq 6 p corresponding to nonsaturated components of x and J \u00bc fj1; " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001132_0470870508.ch20-Figure20.1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001132_0470870508.ch20-Figure20.1-1.png", + "caption": "Figure 20.1 The Gyricon display consists of spherical beads, with optically contrasting hemispheres, dispersed in a transparent rubber sheet. A spherical cavity around each bead is filled with oil", + "texts": [ + " Because media thickness uniformity is not critical, these large sheets can be economically made into large displays. Subject to the flexibility of the addressing system, a Gyricon display with plastic windows can be rolled up for easy transportation and storage. Flexible Flat Panel Displays Edited by G. P. Crawford # 2005 John Wiley & Sons, Ltd., ISBN 0-470-87048-6 The Gyricon display consists of spherical beads, with optically contrasting hemispheres, dispersed in a transparent rubber sheet. This is illustrated in Figure 20.1. A spherical cavity around each bead is filled with oil, allowing it to rotate in response to an electric field. One field polarity will cause hemispheres of a first color to face a viewing window, reversing the field polarity will allow the hemispheres of the second color to be seen. These bead orientations will persist after the field is removed, allowing long-term image storage. The beads used in Gyricon displays are typically 100 mm in diameter and their coloration comes from the use of pigments", + " The first commercial applications of Gyricon technology are in retail signage and message boards. Longer-term applications are expected to include large-area signs, electronic newspaper and book readers and displays for handheld electronics. The Gyricon beads are contained in individual spherical cavities in the rubber sheet. The cavity diameter is typically 25% larger than the bead diameter. The oil that fills the cavity is clear and dielectric. A photomicrograph of the edge of a Gyricon sheet (same aspect as Figure 20.1) is shown in Figure 20.2. The beads are called bichromal because they consist of two colors. Associated with each color is an electric charge; one hemisphere has a different magnitude of charge than the other, or a different polarity. Each bead has a dipole moment that is proportional to the charge difference between the two hemispheres. It also has a monopole moment that is proportional to the net charge on the bead. Both are important to the electro-optical behavior of the bead. A single bead in its cavity is illustrated in Figure 20" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002076_2008-01-1044-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002076_2008-01-1044-Figure2-1.png", + "caption": "Figure 2: Piston free body diagram, side view", + "texts": [], + "surrounding_texts": [ + "Computational models are very important for every aspect of engineering design, so the optimal operation of the piston, and the internal combustion engine as a whole, greatly depends on them. The piston, during its operation, apart from the axial reciprocating motion experiences small transverse oscillations. The identification of the magnitudes of these oscillations and the ability to control them is very crucial, as the piston performance depends on them. In the process of this identification computational models are utilized. Piston computational models were developed in the early 1980\u2019s. Li et al. [12] developed an automotive piston lubrication model to study the effects of piston pin location, piston-to-cylinder clearances and lubricant viscosities on piston dynamics and friction assuming a rigid piston. Li [11] considered the elastic deformation of the piston skirt; integrating this with hydrodynamic lubrication has formed the elastohydrodynamic lubrication analysis which is considered by most of the recent efforts [2, 3, 4, 5, 8, 9, 13, 22, 24, 25]. All these have contributed in better understanding the piston secondary dynamics problem. Zhu et al. [24, 25] were the first to consider the elastic deformation of the cylinder bore. In more recent years Duyar et al. [4] used the mass-conserving Reynolds equation to solve for the hydrodynamic pressures developed on the skirt. This method predicted lower hydrodynamic pressures which allowed for higher transverse motion of the piston and consequently higher contact forces. All these models however assumed that the effects of piston motion along the wrist-pin and land interactions with the cylinder bore are negligible. The authors have recently introduced [15] a new piston dynamics model that considers translation along the wrist-pin and rigid second land interactions with the cylinder bore. They modeled a conventional gasoline piston and demonstrated that motion along the wrist-pin becomes important in predicting piston wear, especially when the cylinder bore deformation and temperature distributions are asymmetric. In this work, the authors developed the above model further to consider an elastic second land. In the current model it is assumed that pressures due to lubrication or scuffing have no effect on the second land deformation, however, it deforms due to combustive, inertial and thermal loads. Only the second land was chosen to be modeled as it is, traditionally, the land with the larger diameter, thus it is expected to have the most dominant interactions with the cylinder bore. Also the piston rings are assumed to have negligible effects on the piston dynamics. The authors utilize a new generation gasoline piston with uneven thrust sides to demonstrate the differences on piston motion and piston wear between different modeling approaches." + ] + }, + { + "image_filename": "designv11_69_0002108_20080706-5-kr-1001.01051-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002108_20080706-5-kr-1001.01051-Figure1-1.png", + "caption": "Fig. 1 Intercept geometry", + "texts": [ + " The use of aerodynamic lift increases the divert capability of the missile up to 100%, while the use of divert thrusters provides a fast response to the guidance command. Model uncertainties created by the interactions between the airflow and the thruster-jets are taken into account and compensated for by SOSM-based autopilot. The integrated SOSM guidance-autopilot algorithm is tested via computer simulations against ballistic maneuvering targets. The following state model (Shtessel et. al., 2007) of missiletarget engagement kinematics (Fig. 1) is used 2 2 , / sin( ) , / , / cos( r V V V r V r V V rV \u03bb \u03b3 \u03bb \u03bb \u03b3 \u22a5 \u22a5 \u22a5\u22a5 \u22a5 =\u23a7 \u23aa \u23aa = + \u0393 \u2212 \u2212 \u0393\u23aa \u23a8 =\u23aa \u23aa ) ,= \u2212 + \u0393 \u2212 \u2212\u23aa\u23a9 \u0393 (1) 978-3-902661-00-5/08/$20.00 \u00a9 2008 IFAC 6226 10.3182/20080706-5-KR-1001.1344 where r is the range along line-of-sight (LOS), \u03bb is the LOS angle; M\u03b3 is a missile flight path angle, ( ) is LOS rate, V \u03bb\u03c9\u03bb = /rad s r \u03bb\u03c9\u22a5 = ( ) is a transversal component of relative velocity in the reference frame rotating with LOS, sm / \u0393 is missile normal acceleration, (disturbances, ) are projections of bounded target acceleration along and orthogonal to LOS , \u22a5\u0393 \u0393 2/ sm It is known (Shtessel at" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001787_cdc.2005.1582818-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001787_cdc.2005.1582818-Figure1-1.png", + "caption": "Fig. 1. Admissible range for the gradient vector", + "texts": [ + " The gradient is evaluated at the vertices of a simplex by projecting the gradient on the normal vectors of the bounding hyperplanes. This results into a linear inequality system for the input u(vi) in these vertices vi. First, the continuous inequalities are presented. The index e represents the number of the facet Fe where the trajectories should leave the simplex. The vertex ve is the only vertex that is not part of the facet Fe. For illustrating the following equations, a two dimensional problem is given in figure 1, where e = 3. \u2200 vi \u2208 V \\ {ve} : nT Fe x\u0307(vi, u(vi)) > 0 (7a) \u2200 vi \u2208 V \\ {ve} : \u2200Fj \u2208 F \\ {Fe, Fi} : nT Fj x\u0307(vi, u(vi)) \u2264 0 (7b) \u2200Fj \u2208 F \\ {Fe} : nT Fj x\u0307(ve, u(ve)) \u2264 0 (7c) n+1\u2211 j=1 j =e nT Fj x\u0307(ve, u(ve)) < 0 (7d) The first system of inequalities (7a) requires that the projection of the gradient on the normal vector of the facet Fe in the vertices point out of the simplex P . The second system of inequalities (7b) requires that the projection of the gradient on the normal vector of all other facets point into simplex P . E.g. in vertex v1 in figure 1, the projection of the gradient on the normal vector of F2 has to point into the simplex P . The admissible range of the directions of the gradients in the vertices are marked as dotted lines. The third condition (7c) forces the gradient in vertex ve to point into the simplex P . The last condition (7d) forces the gradient not to vanish in the vertex ve. In the next step these four conditions (7a-7d) from [4] are modified for discrete-time systems. The gradient x\u0307 has to be expressed as a difference xk+1 \u2212 xk. With xk+1 = \u03a6 xk + H uk + \u03c6 the expressions above modify to the following equations, In is the unity matrix of the dimension n: \u2200vi\u2208V \\{ve} : nT Fe ( (\u03a6 \u2212 In) vi + H u(vi)+\u03c6 ) >0 (8a) \u2200vi \u2208 V \\ {ve} : \u2200Fj \u2208 F \\ {Fe, Fi} : nT Fj ( (\u03a6 \u2212 In)vi + Hu(vi) + \u03c6 ) \u2264 0 (8b) \u2200Fj \u2208 F \\ {Fe} : nT Fj ( (\u03a6 \u2212 In)ve + Hu(ve) + \u03c6 ) \u2264 0 (8c) n+1\u2211 j=1 j =e nT Fj ( (\u03a6 \u2212 In)ve + Hu(ve) + \u03c6 ) < 0 (8d) As illustrated in figure 1 the next sample is not necessarily picked up in the neighbor simplex P\u0303 , if the conditions (8a8d) are met. I.e. a trajectory starts in xk in simplex P and in the next sample xk+1 the neighbor simplex P\u0303 is overleaped, although the conditions from [4] are met. Therefore, the discrete-time versions (8a-8d) of the continuous conditions (7a-7d) are only necessary to solve the problem. By adding the conditions (9- 11) the conditions turn out to be sufficient. The additional conditions \u2200vi \u2208 V \\ {ve} : \u2200F\u0303j \u2208 F\u0303 \\ {F\u0303e} : n\u0303T F\u0303j ( (\u03a6 \u2212 In) vi + H u(vi) + \u03d5 ) \u2264 0 (9) keep the direction of the difference with respect to the dynamic in simplex P in the neighbor simplex P\u0303 " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000876_iros.2003.1249272-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000876_iros.2003.1249272-Figure1-1.png", + "caption": "Fig. 1. Graspless Manipulation", + "texts": [ + ". INTRODUCTION Graspless manipulation [ 1 J (or nonprehensile manipulation [2]) is a method to manipulate objects without grasping. In this paper, we deal with graspless manipulation where the manipulated object is supported not only by robot fingers hut also by the environment (Fig. 1). Such contact tasks are usually performed by force-controlled robots to avoid excessive internal force. In some cases, however, force control is inappropriate in terms of manipulation stability; even minute disturbance could perturb the path of the manipulated object. Pushing operation on a plane is a typical example and therefore usually performed by a position-controlled pusher (for example, \u201cstable push [3J). Thus we have to use both position control and force control appropriately to achieve various robotic graspless manipulation" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003953_6.2009-2536-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003953_6.2009-2536-Figure1-1.png", + "caption": "Figure 1. Simplified Representation", + "texts": [ + " This model is useful since it is a simplified representation of a fully flexible wing. This base model simulates motion along a flight path, a flight path change, a change in pitch attitude, and an additional degree of freedom due to the torsion from the wing attachment. The degrees of freedom are the velocities of the center of mass (CM) of the system within a plane, u and w respectively, the change in pitch attitude, \u03b8, and a degree of freedom associated with the torsion of the wing, \u03c6. The simplified representation including reference frame locations is shown in fig. 1. Three degrees of freedom capture classical longitudinal rigid body motion. The fourth degree of freedom is associated with a rigid wing attached to the vehicle with a flexible support. A simple quasi-steady aerodynamic model is used. Nonlinear terms are retained which permit the onset of an LCO at specific flight conditions. The derivation of the 4DOF equations of motion was performed using both Newton-Euler Method and Gibbs-Appell Method to check validity of the development. For the derivation of the equations of motion some reference frames are defined" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003358_imece2010-37832-Figure7-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003358_imece2010-37832-Figure7-1.png", + "caption": "Fig. 7: ADAMS/Rail Model of the railway vehicle.", + "texts": [ + " Figure 6 shows eight worn hollow profiles categorized in two sets, profiles with false flange type of hollowing (No. 1 to 4) and profiles with rolling radius difference type of hollowing (No. 5 to 8). Theses profiles have been measured experimentally by Miniprof instrument. In Figure 6, to exhibit the manner of the profiles hollowing, they have been illustrated beside the S1002 standard profile. The profiles hollowing values are represented in Table 3. Finally, the modeled car body, bogies and wheelsets are assembled. The assembled model of the coach in ADAMS/Rail is shown in Figure 7. 32 ton Mass 568 kg.m2 Ixx 1.97e6 kg.m2 Iyy 1.97e6 kg.m2 Izz 24 m Body Length 2.2 m Body Width 3.0 m Body Height Tab. 1: Specification of Car Body. Izz (Kg.m2) Iyy (Kg.m2) Ixx (Kg.m2) Mass (kg) 3067 1476 1722 2615 Bogie frame 122 810 810 1503 wheelset Tab. 2: Specification of the bogie frame and wheelset. Profile No. HC (mm) Hm (mm) 1 0.9 1.1 2 1.5 1.2 3 2.0 2.6 4 2.7 3.8 5 0.3 3.0 6 0.6 4.3 7 0.6 5.3 8 0.8 6.2 Tab. 3:Hollowing values of the hollow profiles used in Downloaded From: http://proceedings" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001138_ramech.2004.1438973-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001138_ramech.2004.1438973-Figure1-1.png", + "caption": "Fig. 1. Dead\u00b7zone model", + "texts": [ + " For completeness, the system model is given as follows: r i=1 = bw(v) + d(t) (I) where Yi are known continuous linear or nonlinear func tions, Jet) denotes bounded external disturbances, parame ters ai are unknown constants and control gain b is unknown bounded constant, v is the control input, u( v) denotes dead zone nonlinearity described by { m(v(t) - br) u(t) = 0 m(v(t) - bd vet) br bl < vet) < br vet) :S bl (2) where br 0, bl :S 0 and m > 0 are constants, v is the input and u is the output. A graphical representation of the dead-zone is shown in Figure 1. For plant (1) with dead-zone nonlinearity, the u(t) can be expressed as u(t) = mv(t) - dl(v(t)) where -mbr -mv(t) -mbl It is clear that d( v{ t)) is bounded. vet) br bl < vet) < br vet) :S bl From the structure (3) of model (2), (I) becomes r x(n){t) + LaiYi(X(t), x(t), ... ,x{n l)(t)) i=l = (Jv(t) + d(t) (3) (4) (5) where (J = bm and d(t) = bdl(v(t)) + d(t). The effect of d(t) is due to both external disturbances and bd1(v{t)). We call d(t) a 'disturbance-like' term for simplicity of presentation and use D to denote its bound" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001205_pime_proc_1972_186_090_02-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001205_pime_proc_1972_186_090_02-Figure1-1.png", + "caption": "Fig. 1. Precessing rotor mechanisms", + "texts": [ + " Eulerian angle defining rotation of plane con- taining 0 about fixed (mainshaft) direction. Eulerian angle defining rotation of rotor about its polar axis relative to plane containing %, it is zero when principal axis m2 is normal to plane containing 8. In the discussion in Appendix 1 the three mechanisms outlined in Figs 1, 2 and 3 can be described as arrangements for connecting compound rotations of a \u2018near free body\u2019 type, i.e. motions involving low inertial forces, with simple rotation of a shaft. The reasons for this are discussed in Appendix 1. 2.1 Precession In Fig. 1 the motion of the rotor is a precession. Rotation of the mainshaft at a rate 6 causes simultaneous rotation of the rotor at a rate 4 relative to the tilted portion of the mainshaft. The ratio between the speeds of the rotor and shaft is controlled by the bevel and layshaft gears. The two examples show gear proportions appropriate to the frequency ratios of most practical interest. The case where = -24 leads to applications involving one cycle of volume variation for each rotor revolution; the case where 6 = -3$/2 leads to applications such as the four cycle engine in which two cycles of volume variation occur during each rotor revolution" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000853_rd.202.0109-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000853_rd.202.0109-Figure2-1.png", + "caption": "Figure 2 Example of lattice confinement by circular etched groove, A dc current applied to the circular conductor reduces \"friction\" forces of magnetostatic interaction with domains lying outside the groove,", + "texts": [ + " A result found by investigating a variety of garnet films is that mode stability and phase boundary reproducibility are greater in thicker films and films with smaller tilt of the [ 111] crystal axis from the film normal. The film used for Fig. 3 is relatively thick (12.6 /Lm compared to the value of the material length parameter, I 0.66 /Lm), has a tilt of 0.3\u00b0, and has composition Euo.nY2.23Ga, Fe 40o.2 and 41TMs = 1.75 x 10- 2 T (J 75 G). Observed responses of bubbles to pulsed currents indicate that a profile like that of Fig. 2(a) occurs when the coil plane is separated from the film by the thickness of a 0.15 mm coverglass, as shown in Fig. I (a). The field well in Fig. 1(c) results when the separation is increased by an additional 0.38 to 0.50 mrn, the substrate thickness with the sample turned OYer. To investigate the RBL phase in detail requires sup pressing other phases and preventing the bubble escape that can occur in the weak field-well confinement. Be cause of the resulting discontinuity in the domain wall energy, a strong, steep boundary potential is provided by the etched groove with steep sides [18], illustrated in Fig. 2. To avoid bubble interactions with defects at the imperfect groove, it is also desirable to repel the bubbles from the edge. This can be done by adjusting the width of the groove to accommodate the width of a stripe domain, thus allowing a trapped stripe to repel the bubbles from the groove. Alternatively, a de current may be applied to the conductor loop of Fig. 2 to collapse do mains inside the groove and shrink bubbles near the rim of the rotating lattice, thus reducing magnetic-dipole interactions with domains outside the lattice. Figure 1(c). (b) Streak photograph of rotating bubble lattice [RBL phase in (a i ]. 112 ARGYLE. SLONCZEWSK1, AND VOEGELI IBM J. RES. DEVELOP. The procedure for taking data consists of placing the sample, usually with film side away from the coil to pro vide the field well in Fig. I (c) ; adjusting the bias in the range of bubble stability; and \"chopping\" stripe domains using pulsed fields to produce a bubble lattice [ 19]", + " The groove configura tion provides confinement with a sharp potential dis continuity. In this study we (I) measured the angular rotation velocity in response to pulsed fields, (2) ob tained its velocity spectrum in response to rf sinusoidal excitation, (3) investigated the effect on the results in (2) of applying an in-plane de field, and (4) determined the bubble radial breathing mode response averaged over the lattice both when the lattice rotates freely and when it is \"clamped\" between a pair of long parallel grooves. The sample with the circular confinement groove in Fig. 2 has composition EUo.7Y2.3Gal.lFe3.9012' a thickness of 3.6 /Lm, a 47TMs value of 1.75 x 10-2T( 177 G), and a parameter I of 0.56 um. The response to pulsed bias given in Fig. 4(a) is a plot of angular rotation rate versus pulse amplitude. Data were taken up to the point at which lattice distortions, e.g., shearing, take place. The bubble size and lattice spacing are nominally 7 /Lm and II /Lm, respectively. The bias field (about 3.26 X 103 AIm, or 41 Oe), and pulse width (about 0.46 /Ls) where chosen optimally, i.e., for maximum rotation rate and minimum hesitation and lattice distortion. Interaction with outlying domains was eliminated by using 100 rnA de current in the concentric conductor strip line shown in Fig. 2 so as to collapse domains lying in the groove and shrink the circular layer of peripheral bubbles that accommodates the angular boundary of the hexagonal lattice to the circular shape of the confinement groove. The maximum bubble velocity, 1.5 /Lm per pulse [right-hand ordinates in Fig. 4(a)], observed at a rim having a radius of 125 /Lm, appeared limited by lattice shearing. Although this distortion might have been avoided by a more 113 MARCH 1976 BUBBLE LATTICE MOTIONS suitable choice of field-well, it is possible that the threshold for nonlinear velocity response had been exceeded" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002843_6.2007-1788-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002843_6.2007-1788-Figure2-1.png", + "caption": "Figure 2. Rib and kinematic mechanism.", + "texts": [ + " However, the advantages found with wing twisting can only be realized if the structural resistance is minimized. One of the obstacles is that in order to twist the wing, the skin must also be twisted. Many researchers have sought to develop a skin capable of twisting while retaining the ability to maintain the airfoil shape. The present work another idea is presented. Instead of forcing the skin to twist as the wing does, allow the support structure to twist underneath the skin. A. Initial Design Figures 2 and 3 show the kinematic mechanism and piezoelectric actuator used to apply torque to the ribs. Figure 2 shows the rib structure that directly supports the wing\u2019s skin. The two fixed hinges are shown as dots and the dotted lines represent moment arms attached to each portion of the rib. These moment arms rotate with the rib portions. A piezoelectric stack is attached to the end of each of the moment arms. In order to analyze the kinematic linkage and actuator in more detail, a schematic of the linkage circled in Fig. 2 is shown in Fig. 3. Points a and c in Fig. 3 are the hinges about which the front and rear portions of the ribs rotate, while links ad and ce are the moments arms that rotate with front and rear portion of the ribs, respectively. Application of an electric field causes the piezoelectric stack, link de, to extend. This makes the moment arms rotate away from each American Institute of Aeronautics and Astronautics other. Point g needs to be mechanically grounded in order to force equivalent rotation of the front and rear portions of the rib" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003361_icfcc.2010.5497776-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003361_icfcc.2010.5497776-Figure1-1.png", + "caption": "Fig. 1. The Pendubot system", + "texts": [ + " It is a simple underactuated mechanical system. Many studies have been reported for solving the pendubot problem, such as feedback stabilization of Spong[8]and output regulation of Sanposht[9] [10] et al. Compare with the previous method, the neural networks method can approximate the solution of the regulator equation up to an arbitrarily small error in any given compact subset and reduce tedious computation. We illustrate the approach by designing a tracking control law for the pendubot system as shown in Fig. 1. The continuous-time state space model of the Pendubot system is obtained from the dynamic model of the Pendubot such that 1 2 2 2 3 4 4 4 1 1 2 1 3cos( ) cos( )c c x x x Dx x x x Dx y l x l x x = = = = = + + (18) Where 1 2 3 4 1 1 2 2( , , , ) ( , , , )x x x x q q q q= , and 2 2 4 12 3 3 1 2 1 {[ cos( ) ( cos( )) Dx g x x \u03b8 \u03b8 \u03b8 \u03b8 \u03b8 = \u2212 \u2212 2 2 2 2 3 3 2 4 3 3 3 2sin( )( ) ( ) sin( )cos( )( )x x x x x x\u03b8 \u03b8 \u03b8+ \u2212 \u2212 33 25 1 3cos( ) cos( )] }g x x x u\u03b8 \u03b8 \u03b8+ \u2212 2 4 3 3 2 5 1 3 2 1 [ sin( )( ) cos( )Dx x x g x x\u03b8 \u03b8 \u03b8 = \u2212 + + + 2 3 3 2( cos( )) ]x Dx\u03b8 \u03b8+ where lc1 and lc2 are the length of the first and second links, respectively, with lc1 = 0:5 and lc2 = 1" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003757_20091021-3-jp-2009.00015-Figure4-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003757_20091021-3-jp-2009.00015-Figure4-1.png", + "caption": "Fig. 4. Mechanical system.", + "texts": [ + " The two following exercises deal with controller design in the frequency domain. The closed-loop system from Fig. 3 is applied again. This will make it possible for the students to compare the different design methods. The students will observe that the results using the Ziegler-Nichols methods can be improved quite a lot when a detailed model is available and is used in connection with the design. The last exercises are based on a number of real systems. This includes both a mechanical and an electricalmechanical system as shown in Fig. 4 and 5. The main issue here is to let the students go through all steps in controller design. Non-linear models must be formulated from the mechanical and electrical laws. Simulink models built based on the non-linear models. Linear models are derived and analyzed; controllers are designed and eventually implemented in the Simulink model. The controllers are validated by analysis as well as simulation on both linear and non-linear models. The practical exercises are an alternative to the Matlab and Simulink exercises, described in Section 4" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001777_oceanse.2005.1513234-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001777_oceanse.2005.1513234-Figure1-1.png", + "caption": "Figure 1 reproduces the 3C model in [SI, embodying communication, coordination and cooperation. The communication support consists in message systems. The cooperation support includes conversation contents and protocols. The coordination support manages the workflow carried out by the team. FiguTe 2 focus on the time axis, and the kind of know-how that may be called upon.", + "texts": [], + "surrounding_texts": [ + "Oceans - Europe 2005\nShips Confining an Oil Spill Over: A Scenario for Automatized Cooperation\nJ.F. Jimenez, J.M. Giron-Sierra, A. Do 'n ez,\nDep. ACYA, Fac. CC. Fisicas, Universidad Complutense de Madrid Ciudad Universitaria 28040 Madrid, Spain. gironsi@dacya.ucm.es\n(1) Canal de Experiencias Hidrodinamicas. CEHTPAR El Pardo. Madrid, Spain\nJ,M, De la Cruz, and J.M. Riola ?P\nAbsrract Cooperation between robots is an important contemporary issue. This can be translated to the marine environment, either using marine robots or introducing automatics in the operations of ships. A general research on this kind of problems has started in our group, after several years o f developing autonomous robotized ships. Several scenarios have been proposed for the study of cooperation details, This paper focus in a very interesting case, which is representative of other cases: several ships towing booms for oil spill over confinement. It turns out that the cooperation problem is not trivial. Along the operation several phases can be distinguished, and several coordination problems and needs appear. A computer simulation has been developed, after physics based analysis, and some initial coordinated control strategies have been proposed and tested. These strategies are supposed to be applied through verbal orders to captains. Along the operation phases the role of captains change, for an adaptive coordination. The paper introduces the research topic, then describes the scenario and its simulation, then focus on the cooperation probiems emerging from the operation phases and the control and coordination solutions that bave been proposed, and finally draws some conclusions. could be to include computers on board the cooperating ships, exchanging verbal messages and taking decisions under the supervision of the sfiips captains. A communication protocol should be established for such purpose. This research started by considering some different scenarios requiring cooperation of ships. First studies began with animated simulations, helping us to capture the specific needs along the operations, for a good cooperative work. In paralell a set of scaled ships are under construction, for experimental studies in the same scenarios considered by the simulations. This paper focuses on an interesting scenario. The general case will be several ships towing a boom, for oil spill over confmement. The study began with the simplest cooperation case: two ships towing the boom. The order in this paper will be the following. The next two sections will be devoted to a condensed review of cooperation in robot teams, and its translation to the marine environment. Then, the oil spill over confinement case will be studied. Then, physics based modeling will be introduced and the simulation will be briefly described. Finally, the operation will be studied, showing the specific cooperation needs that appear. The paper f ~ s h with some conclusions in view of the practical development of protocols.\nlI. ROBOTS AND COOPERATION I. INTRODUCTION\nAfter seven years of reseatcb on seakeeping control of ships by moving actuators such flaps and T-foils [ l J ] , a wider research on automatized cooperation of marine crafts has been launched. There are important reasons in favour of this initiative. Also, recent ecological catastrophes, Iike the Prestige sinking near the Spain coast, have motivated a particular attention to marine operations involving several ships and submergibles, to avoid as much as possible the effects of oil spill over..\nDuring the last part of our research on seakeeping control, a six degrees of fteedom approach has been followed This was a continuation of the frst studies in a towing tank with head seas. A scafed down replica o f a fast ferry has been built. This replica has on-board autonomous control, including speed, heading and seakeeping control. The replica has self-knowledge of its position.\nSince our field is automatic control and robotics, it was clear for us that the autonomous scaled ship is a kind of surface marine robot. Moreover, since part of our research concerns teams of mobile robots, the idea of taking the sea as an scenario of automatized cooperation between robots ~ t u r a l l y arised.\nIndeed, the daily contact with people having long experience in the sea, gave us the impression that a good realistic alternative\nThe initial research on mobile robots proposed hierarchical behaviour control architectures. A representative metaphor for the on-board control was a three person structure. A captain deciding where to go and the main traits of the path. Then, there is an officer to elicitate more planning details. And there is a pilot governing in real-time the ship, executing with some adaptations the plan. It was found that the hierarchical structure may lead to rigid and complicated control. A reactive paradigm was proposed in the form of a subsumption architecture [3], which offers a simpler way for robot behaviour building. However, the general problems of having a target and deciding how to reach it, requires some deliberative features. As [4] asserts, the consensus is now to combine deliberative and reactive paradigms.\nThere are two main axes for cooperation in a robot team. One axis is distribution of responsibilities. The other axis is time. There are coordination needs, and perhaps syncronization could also be needed. The book [ 5 ] provides examples of several terrestrial mobile robots pushing together and object, and other multiple physical agent cues with learning, adaptation, self-organisation and spatial distribution aspects. Protocols for robotic agents communication and interaction are more specifically studied in [6] . This reference says that the tasks must be assigned to agents along\n0-7803-9i03~9/051$20.0002005 IEEE 1226", + "time in function of the expertise and aptitude (it includes physical position) of the agents in each moment. The planning and decisions could be done in a CentTdized or distributed way. A blackboard may be used for the team decisions.\nThere are two main extremes. One is to bring full authority to a central coordinator. The other is all agents having full autonomy. There are intermediate formulas, combining central authority with some delegation in a certain degree. This degree may vary along an operation. Related with this is the issue of the combination of local knowledge and general knowledge about the operationd status that agents should have. Part of our specific experience, in a distributed avionic control system, is related with this matter [7]. In this case, smart components usually decide in function of local knowledge (for instance, a valve closing when a particular tank reaches a specified level), but they may have a different behaviour if they receive a message telling there is a fault mode in the system.\nComputer networking also leads to cooperation studies and developments. The book [SI on computer supported cooperative work includes interesting pages on multiagent systems, shared information, voting decision schemes, conflicts and reconfiguration, negotiation and contracts between agents, etc.\nFig. I . 3C model : communication, coordination and cooperation.\nFig. 2. Coopemtion issues related to h e .\nAn interesting contemporary topic, which is relevant for OUT research, is formation control. Reference [9] on following-the-leader is a good recent source of information on the topic.\nm. COOPERATION IN MARITIME SCENARIOS\nThere are m y maritime circumstances or operations involving several marine crafts. For instance, the maneuvering of ships in confined waters or channels, to avoid collisions. In certain cases there are attracting forces that cause difficulties [ 101. This is also the case of 6eight or person offshore transfers between two ships in motion. This example can be seen as a particular instance of a general rendez-vous operation categoly, which can be static or in motion. Some recent papers on this kind of problem are [11],[12] and [13].\nThe deployment of nets, sets of buoys, barriers, etc. is another kind of operation that usually requires cooperation of ships. A particular case is the confinement of oil spill overs. This operation may continue with recycling, towing, suction, etc.\nOther scenarios may include some sort of formation control, not only in the case of military or fishing operations. For instance, high speed transportation may require a temporary corridor to be respected by other ships.\nEmergency or disaster cases may involve special cooperative towing tasks, scanning for survivors in a certain zone using a team of ships and aircrafts, etc. Patrol and security operations can be of similar nature. This is connected with robotic topics about exploration and harvesting by, perhaps heterogeneous, robot teams.\nRecent scientific meetings are beginning to devote sessions to robot coordination in a marine context, with papers such [14] about coordinated motion control, or [I51 on towing.\nThe interest for cooperation also extends to the underwater medium. For example in servicing cases, with rendez-vous and docking aspects, In [16] platoons of underwater vehicles are considered.\nFrom the industrial point of view, it is interesting to notice that cooperative operational requirements may demand special working, maneuvering or dynamic positioning capabilities of the involved watercrafts.\nIV. SCENARIOS AND INITIAL FORMALIZATION\nFor the fxst steps of our research, an initial formalization of the cooperation scenarios has been introduced. The objective is to get a set of operation classes, with representative archetypes or examples for each class.\nLet us begin by considering different types of things that can be present in an scenario:\nPassive and mute. For instance a rock, a wreck, an oil\nPassive, speaking. For instance, a buoy with a light or\nActive, mute. For instance, A ship with a broken radio. Active and speaking. For instance, a ship with working\nmouth, a barge.\nradio.\nradio.\nThis classification uses the active or passive terms, referring to auto-propulsion or not. The different things can act as agents, obstacles, targets, or beacons (references). The classification can be extended, considering, for instance, whether the things have any sensor (for instance, human. vision, radar, etc.), and a decision system on board (for instance, a pilot).", + "Having several watercrafts in cooperation, the decisions can be\nA cenhal coordinator. He takes the main decisions, which can be based on information h m the team. Each individual, in silence. The individuals know what to do, only looking at the rest of the team. The operation is pre-planned. Speaking individuals. They give information about their own decisions. Hierarchical structure. For instance a fleet of smal l fleets.\ntaken according with the following organization alternatives:\nAlong an operation the organization may change. This includes the roles in the team; for instance, the coordinator may change. OrganiZatiOMl transitions can be notified to the team; or not, since they may obey to a pre-defined strategy.\nIn general, part of a operation can be planned in advance, and part must adapt to changing circumstances.\nThe operations can be characterized by the existence of certain virtual or physical constraints. For instance, channels or confined waters, positional formation of the ships, physical links between them (for instance, when towing or making transfers).\nAccording with the team working structure, in each moment, there are several communications schemes IO choose:\nCentralized polling. For iastance, with a coordinator. Turn-following. For instance, using tokens or time stamps. CSMA. For instance, with wireless ETHERNET. A strategy with priorities\nV. THE OIL SPILL OVER SCENARIO\nIt was noticed, during the response to the Prestige disaster , that booms are usually carried by only two ships. This limits tbe capability of the confinement operation. However, it may happen that there are difficulties in the coordination of ships, even for only NO.\nOur attention was attracted by this case. Trying to know more about what happens, a simulation study was initiated.\nIn a frst simplified analysis, three phases were distinguished in the operation. The next three figures show these phases, using the particular case of four ships handling the boom.\nThe target is an oil leak floating in a certain place. The leak must be confined, and, may be, carried to another place.\nThe fmt phase is leaving the harbour and towing the boom near the target. Resistance to motion should be minimized., and the distance between ships must be safe enough. Figure 4 shows an schematic view of this initial phase. The ships must advance with same speed and heading. It is good to know what ship is slower, to adapt the team speed to a reference, leaving some margin for individual adaptations." + ] + }, + { + "image_filename": "designv11_69_0001295_physreve.72.032701-Figure5-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001295_physreve.72.032701-Figure5-1.png", + "caption": "FIG. 5. Energy profiles \u2212W Q ,U for s=2.1, 4.0, 10, respectively. left to right . Positive values of W have been clipped.", + "texts": [ + " 3 We average these contributions over the x ,y -plane, using cos2 qxx =cos2 qyy =sin2 qyy = 1 2 , sin3 qyy =4/3 , sin4 qyy = 3 8 . The averaged specific energy is w = 1 8 Kqx 4u0 2 + 1 8 Bqy 2u0 2 + 3 256 qx 4u0 4 \u2212 3 Sqx 2u0 2 + 4 S2 . 4 All coefficients of this equation except one control parameter s can be set to unity by introducing the dimensionless quantities Q4 = K Bqy 2 qx 4, U2 = 3 32K u0 2, W = 3 4BKqy 2 w, and s = 8 3 BKqy S . One arrives at the equation W = Q4U4 + Q4U2 + U2 \u2212 sQ2U2 5 where a constant proportional to S2 has been dropped. Figure 5 shows the energy profiles for three selected values of s. The extrema of W are found from dW /dU=4Q4U3+2Q4U +2U2\u22122sQ2U=0, and dW /dQ=4Q3U4+4Q3U2\u22122sQU2=0. These equations are satisfied by the trivial solution U=0 chain compression, no buckling . In addition, there are the nontrivial solutions Q= \u00b11, U= \u00b1 s\u22122 /2 for s 2. The interpretation is as follows: The wave number of the equilibrium deformation is Q =1, independent of the excess saturation s. Expressed in actual spatial coordinates, qx 2 = B K 2y0 " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000147_thc-2003-11202-Figure3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000147_thc-2003-11202-Figure3-1.png", + "caption": "Fig. 3. Experimental setup from above (left) and a side view (right).", + "texts": [ + " (6) consist of the elastic contribution Fe(q) and direction dependent dissipative moments Fdsgn(q\u0307) also known as Coulomb friction [20]. The passive moments can be summed up as a time and angle dependent column vector \u03c4 p(q, t) or simply \u03c4p which was the focal point of this study: \u03c4p = [\u03c4p1 \u03c4p2 \u03c4p3] T (7) In the performed experiment a positionally controlled antrophomorphic 6-DOF industrial robot (Yaskawa c\u00a9 MOTOMANsk6) was used for imposing a slow linear movement on the human arm in the sagittal plane (Fig. 3). A JR3 c\u00a9 4 dimensional strain gauge force sensor was mounted on the manipulator end effector and used for force data collection. The maximum force for the specified output was \u00b1110 N, with an acquisition resolution of 12 bits. A bicycle-like circular rubber coated handle was mounted on top of the sensor in such a way, that rotation around the x axis was freely allowed. The next element in the system was a bus passenger seat, equipped with additional straps as evident from Fig. 3. The plane of motion was perpendicular to the ground and fully aligned with the sagittal plane of the subject. In the first part of the experiment, the subject was asked to keep his muscles relaxed while holding the handle. The handle was held gently, while still allowing the arm to stay in good contact during the movement. Before starting the real measurements, it was also inspected whether the slight muscle activation due to gripping had any significant effect on the passive torque identification process" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000178_tmech.2002.805622-Figure3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000178_tmech.2002.805622-Figure3-1.png", + "caption": "Fig. 3. Local structurization division of Steward platform using sensors.", + "texts": [ + " If represents the length vector of six legs structure and Jacobian matrix represents the relations between position and orientation changes of mobile platform and then [7] (10) where , , represents the line velocity and angle velocity of the platform, respectively. The Jacobian matrix is as follows: ... ... The decoupling matrix proposed in this paper was derived from this Jacobian matrix. III. INTRODUCTION OF LOCAL STRUCTURIZATION METHOD (LSM) LSM is used to solve the forward kinematic problem (FKP) of parallel structure. It is a positive kinematics solution to the problem. LSM for six-leg virtual-axis NC machine uses G/P mechanism with two extra sensors as shown in Fig. 3. G/P mechanism means a mechanism that is composed of the globular and prismatic joints, its basic concept is based on the following. 1) If parallel structure can be split into several minimal structures (G/P mechanism) such as tetrahedral and/or triangle, and relations between them are known, one can solve FKP of parallel structure. 2) If parallel structure cannot be divided into minimal structures, it can be split through using passive link and/or virtual link with sensors. 3) If the positions of several sides can be specified with respect to a coordinate system. The tetrahedron (in space) or triangle (in plane) can be constructed and all relations between them in coordinate can be obtained. 4) If one G/P mechanism can be solved, then other G/P mechanisms, which share one or more sides with the solved G/P mechanism, can also be solved. For six-leg virtual-axis machine, two sensors are located at one of the base joints, for example, at point as shown in Fig. 3. In this case, the G/P mechanism can be described by the triangle and the vector can be fully defined. The points and are known. The lengths of , , are known. Assuming is , is and is ( ), then the G/P mechanism can be solved through and the position and orientation of mobile platform can be obtained. Additional details can be found in [8]. IV. DEDUCE OF DECOUPLING MATRIX It will then be the choice of the designer to determine the number of sensors necessary to get the best compromise between the supplementary cost and complexity of the manipulator (increasing with the number of sensors) and the computation time for solving the direct kinematics problem (decreasing with the number of sensors) [9]. So the number of sensors used in this paper is the necessarily minimum number. From formula (10), note that the Jacobian matrix is very complex. In order to analyze conveniently, the supported Jacobian matrix is then formula (10) becomes (11) where . . . . . . ... ... . . . . . . , ... ... . As shown in Fig. 3, vector for ( ) can be represented in two loop equations as follows: (12) where is a constant vector from to and the velocity relation of becomes (13) From , a new velocity relation can be derived (14) Therefore, the second term in the right-hand side is , since rigid body motion occurs along . Then, the prevoius velocity equation, with appropriate relations of (12) and (13), becomes (15) For further analysis, we established a sensor Descartes coordinate system\u2014leg-coordinate system . Select the direction of as the direction of axis, axis lies in the plane and in the outward direction from B, axis is then decided by the right-hand side rule" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000183_ultsym.1998.762239-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000183_ultsym.1998.762239-Figure2-1.png", + "caption": "Fig. 2 Rayleigh wave generation.", + "texts": [ + " PRINCIPLE When the Rayleigh wave propagates on elastic material surface, particles on the surface move along an elliptical locus as hown in Fig. 1 . A slider arranged on the elastic substrate isdriven by frictional force. The slider is pre-loaded so that friction force is enough to drive. A piezoelectric material of 128\" Y-cut LiNbO, 0-7803-4095-7/98/$10.00 0 1998 IEEE 1998 IEEE ULTRASONICS SYMPOSIUM - 679 was used for the SAW motor. When high frequency voltage is input to an interdigital transducer(1DT) on the piezoelectric substrate, the Rayleigh wave is generated and propagates, as shown in Fig. 2. 2 0.4 ~ The slider made of silicon has many projections ~ o,3. on its surface in order to control contact conditions. Figure 3 is a photograph of the slider surface. The 3 0 . 2 ~ diameter of the projections is 10 to 50 pm. v - I PREVIOUS MOTOR We have also updated 10 MHz SAW motor performance. The size of the transducer was 15 x 60 x 1 mm3. The operating frequency was 9.6 MHz. The Fig. 6 Step driving. Time [msec] ness @ Silicon Slider LiNbO, Substrate Fig. 4 10 MHz motor experimental setup. Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002347_15397730701404684-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002347_15397730701404684-Figure1-1.png", + "caption": "Figure 1. Components of turbocharger.", + "texts": [ + " A locomotive turbocharger is a device in which heat energy from engine exhaust gas turns a radial inflow turbine along with a centrifugal compressor on the same shaft and supported by two inboard-mounted fully floating shaft sleeve, such that inflow air is pressurized by the compressor and supplied to the engine, thereby improving the engine\u2019s combustion efficiency (Toshimitsu, 2002). Thanks to this relatively simple principle, nearly 75% of the total output of an engine is the result of turbocharging. So the running stability of turbocharger is a precondition for other components to run safely. The components of turbocharger are illustrated in Fig. 1. The centrifugal compressor is perhaps the most common example of high-speed turbomachinery (Mcame and Robinson, 1999). In the long-distance transport of material, centrifugal compressors are playing more important roles (Yamaguchi, 2002). The high rotational speed, high centrifugal force, and high dependability are basic characteristic of compressor. The exhaust flow from the engine is directed over the blades of the turbine to provide the force to turn the shaft and compressor (Bhope and Padole, 2004)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001455_s00542-006-0293-x-Figure11-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001455_s00542-006-0293-x-Figure11-1.png", + "caption": "Fig. 11 Traction torque transmitted from roller to web", + "texts": [ + " (8), it is recognized that the entrained air film thickness decreases drastically in the web moving direction as shown in Fig. 10 and the minimum air film thickness hmin is given by the following equation. hmin \u00bc 0:589R 6gU T 2=3 2kTB gtwU \u00f012\u00de When the minimum air film thickness hmin is larger than 3r, there is no contact between the web and roller, and then the slippage will occur. Under such a condition, the slip onset velocity Uslip is determined from the following equation. 0:589R 6gUslip T 2=3 2kTB gtwUslip \u00bc 3r \u00f013\u00de 3.2 Formulation of model 2 Figure 11 shows the interface model between the web and roller. In the figure, when the traction torque transmitted from rotating roller to web, or from web to roller, is less than the brake torque at the bearing, the slippage will occur at the interface between the web and roller. Then, the slippage condition is expressed as follows. FR\\fr \u00bcMb \u00f014\u00de Based on Euler\u2019s belt theory, the traction force F is given by: F \u00bc TL elB 1 \u00f015\u00de where the traction coefficient l is estimated from Eqs. (10), (11), which is a function of velocity U" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003157_iccas.2010.5670141-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003157_iccas.2010.5670141-Figure1-1.png", + "caption": "Fig. 1 Geometry of single-link flexible manipulator", + "texts": [ + " The suppression of the residual vibration can be accomplished by rotating the joint along the generated trajectory, that is, the proposed control scheme is an feedforward control that doesn\u2019t require sensors to measure unwanted vibrations. By performing the simulations as well as the experiments, the effectiveness of the proposed trajectory planning method for suppressing the residual vibration and saving the operating energy is demonstrated. We consider a single-link flexible manipulator with a tip mass as shown in Fig. 1. The variable s is the coordinate along the deformed configuration of the flexible link. Since the manipulator is constrained to move only in the horizontal plane, the influence of gravity on the movement of the manipulator is negligible. In order to accurately obtain a mathematical model of the flexible manipulator, the effects of geometric nonlinearities are taken into account [5, 7]. The position vector r of the flexible manipulator relative to the inertial frame at an arbitrary location s can be expressed as \u23a5 \u23a6 \u23a4 \u23a2 \u23a3 \u23a1 +++ \u2212++ ==\u23a5 \u23a6 \u23a4 \u23a2 \u23a3 \u23a1 \u03b8\u03b8 \u03b8\u03b8 cossin)( sincos)( wusa wusa Y X r , (1) where a is the radius of a rigid hub, \u03b8 is the joint angle, and u and w are the axial displacement and the transverse displacement, respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003398_ijmr.2009.026577-Figure9-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003398_ijmr.2009.026577-Figure9-1.png", + "caption": "Figure 9 3D plot for a section of Section A-A", + "texts": [ + " The highest temperature on the tool-chip interface is equal to the average value estimated through the calculations. The isotherm of temperature closest to the point of measurement (1.25 mm from the cutting edge on the top rake face) is higher for Section B-B (Figure 8) compared with Section A-A (Figure 7). The isothermal contours give a good indication of how the temperatures change with distance from the tool-chip interface. The 3D plot for temperature distribution of section A-A is presented in Figure 9. It is imperative therefore to emphasise here the possible comparison with the work of Hong and Ding. It is a case where the same Titanium alloy has been investigated. Their model predicted temperatures between 872\u00b0C at a speed of 120 m/min, feed of 0.2 mm/rev and a depth of cut of 1.28 mm. Our model predicted an average temperature of 910\u00b0C at 100 m/min and 1169\u00b0C at 150 m/min with a feed of 0.2 mm/rev and depth of cut of 0.25 mm. This near similarity ends with the greater temperature differences with changes in process parameters" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003895_isaf.2009.5307548-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003895_isaf.2009.5307548-Figure1-1.png", + "caption": "Fig. 1 Structure of longitudinal to torsional vibration converter with diagonal slits", + "texts": [ + " Nakamura\u2019s work, we propose a double-rotor ultrasonic motor with a symmetrical longitudinal-torsional sonotrodes to achieve double sides' driving. In the symmetrical construction, the first longitudinal and the second torsional vibration modes of the converter are used. This paper deals with both the simulation and the experimental investigations of a prototype ultrasonic motor. As a result of the experiments with the prototype motor, we have achieved a maximum torque of 1.8N\u00b7m. The structure of the cylindrical longitudinal to torsional vibration converter proposed here in this paper is shown in Fig. 1. The cylindrical converter has a silted part along its circumference adjacent to the nodal position of longitudinal vibration. The slit length required for changing the wave progressing direction is very short, only in the case where the slits are positioned at a vibration node [4]. The diagonal slits are only partially cut at about 45 degrees at the circumference of the cylinder with the inner cylindrical part remaining unstilted. The converter is driven only by a singular longitudinal vibration source" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002778_tmag.2007.893864-Figure6-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002778_tmag.2007.893864-Figure6-1.png", + "caption": "Fig. 6. Outline of the system.", + "texts": [ + " 5 shows the distribution of the parameter Lmax and Smax at the location of the crack in simulations. In the distribution line, the diameter of the line was 9.2 mm. The current of the line was 50 A at 50 Hz. Six magnetic sensors were placed along the ring at 60 apart from each other. The ring diameter was set at 25 mm. At the location of the crack, the changes of the parameter Smax was more sensitive than that of the parameter Lmax. It was confirmed that the improvement of the algorithm for detecting cracks occured by using the new parameter Smax. Fig. 6 shows the outline of the experimental system. In the machine, six magnetic sensors were placed on the ring at 60 apart from each other. The sensor ring diameter was set at 25 mm. The machine measured the strength of magnetic field around the distribution line by running on it. The speed rate of the machine was set at 10 mm/s. The measured data were transmitted to the personal computer, which calculated the parameter Smax. We experimented by using the distribution line, 60 mm in sectional area. Crack depth was 1 mm, and crack width was 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001685_wcica.2006.1712736-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001685_wcica.2006.1712736-Figure1-1.png", + "caption": "Fig. 1. Phase portrait for the second-order SMC system with matched uncertainties", + "texts": [ + " Since v > 0, r1 > 0 and R1 < \u03b1r1, vz(k) \u2212 \u03b1r1 + w1(k) < \u2212\u03b1r1 + R1 < 0, which suggests that the value of x1(k) decreases towards s = 0 and then cross it. However, when x1(k+1) enters the region s < 0 from region s > 0, the following motion of z in region s < 0 will move into the sector of 0 < z(k) < (\u03b1r2 +R2)/(1\u2212|d|). There are two cases when (x, z) jumps into this sector at step k: one is sk < 0 and the other, sk > 0. For the first case, (x1(k + 1), z(k + 1)) will increase towards s = 0 and then cross it due to the fact x1(k+1) = x1(k)+vz(k)+\u03b1r1+w1(k) (see the sequence a \u2192 d \u2192 b in Fig. 1). For the second case (see the sequence a \u2192 b in Fig. 1), (x1(k + 1), z(k + 1)) will move towards the sector \u2212(\u03b1r2 + R2)/(1 \u2212 |d|) \u2264 z(k) < 0. There are similarly two cases two cases of sk > 0 and sk < 0 when (x, z) jumps into this sector at step k. For the first case, (x1(k + 1), z(k + 1)) will go towards s = 0 due to the fact x1(k +1) = x1(k)+ vz(k)+\u03b1r1 +w1(k) as analyzed before (see the sequence b \u2192 e \u2192 c in Fig. 1). For the second case, (see the sequence b \u2192 c in Fig. 1), the state will cross the switching line. Repeating the above motion, (x1, z) will cross s = 0 repeatedly. To find the bound of x1, consider the limiting case, let us say that the trajectory at step k just lands on s = 0\u2212 and step into s > 0 at the next step (see the sequence p \u2192 q in Fig. 1). Hence x1(k) = \u2212c\u22121z(k). Then x1(k +1) = \u2212c\u22121z(k)+vz(k)+\u03b1r1 +w1(k). Because 0 < z(k) \u2264 (\u03b1r2 + R2)/(1 \u2212 |d|), the upper bound for x1 is x1 < \u03b1r1 + |v \u2212 c\u22121|(\u03b1r2 + R2)/(1 \u2212 |d|) + R1 The same reasoning applies and we can show that the lower bound for x1 is x1 < \u2212\u03b1r1 \u2212 |v \u2212 c\u22121|(\u03b1r2 + R2)/(1 \u2212 |d|) \u2212 R1 Overall, the bound of x1 is |x1| < \u03b1r1 + |v \u2212 c\u22121|(\u03b1r2 + R2)/(1 \u2212 |d|) + R1 Section III has fully explored the bounded conditions of discretization behaviors of second-order SMC systems. These results can be extended to higher-order systems" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000210_iros.1996.571043-Figure3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000210_iros.1996.571043-Figure3-1.png", + "caption": "Figure 3: Frame CO", + "texts": [ + " Also, similar results can be seen in other cases that the friction distribution is not uniform, the contact point is slightly away from the foot of the perpendicular, and the object shape is another convex polygon. Considering these results, we assume that the contact point in the trajectory planning is chosen so as the center of friction to be in the velocity cone, and the contact point in the tracking control is settled at the same point. 3 Pseudo Center This section presents the relation among the linear and rotational velocities of the object at the pseudo center, the velocity at the contact point, and the position of the pseudo center. As shown in Figure 3, the object frame CO is fixed to the object base with its origin at the center of friction, its X axis parallel to the object edge including the contact point or the line tangent to the object at the contact point. The position of the contact point and that of the pseudo center expressed in CO are denoted as \"p, and \"p,, respectively. Further, the velocity at the contact point expressed in CO and the angle of the velocity with respect to the X axis of CO are denoted as Ow, and a, and cos cy Ow, = 8, I sina J Now, we know that the object translates without rotation when Ow, is toward the center of friction" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002928_978-3-540-70534-5_16-Figure16.4-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002928_978-3-540-70534-5_16-Figure16.4-1.png", + "caption": "Figure 16.4: Dead reckoning", + "texts": [ + " Dead reckoning In many cases, driving robots have to rely on their wheel encoders alone for short-term localization, and can update their position and orientation from time to time, for example when reaching a certain waypoint. So-called \u201cdead reckoning\u201d is the standard localization method under these circumstances. Dead reckoning is a nautical term from the 1700s when ships did not have modern navigation equipment and had to rely on vector-adding their course segments to establish their current position. Dead reckoning can be described as local polar coordinates, or more practically as turtle graphics geometry. As can be seen in Figure 16.4, it is required to know the robot\u2019s starting position and orientation. For all subsequent driving actions (for example straight sections or rotations on the spot or curves), Figure 16.2: Beacon measurements green beacon red beacon 45\u00b0 165\u00b0 Localization and Navigation 16 the robot\u2019s current position is updated as per the feedback provided from the wheel encoders. Obviously this method has severe limitations when applied for a longer time. All inaccuracies due to sensor error or wheel slippage will add up over time" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000597_cbo9780511547126.019-Figure17.4-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000597_cbo9780511547126.019-Figure17.4-1.png", + "caption": "Figure 17.4.1: Generation of profile-crowned tooth surfaces by application of rack-cutters: (a) for pinion generation by rack-cutter c ; (b) for gear generation by rack-cutter t .", + "texts": [ + "4 PROFILE-CROWNED PINION AND GEAR TOOTH SURFACES The profile-crowned pinion and gear tooth surfaces are designated as \u03c3 and 2, respectively, whereas 1 indicates the double-crowned pinion tooth surfaces. Profile-crowned pinion tooth surface \u03c3 is generated as the envelope to the pinion rack-cutter surface c . The derivation of \u03c3 is based on the following considerations: (i) Movable coordinate systems Sc (xc , yc ) and S\u03c3 (x\u03c3 , y\u03c3 ) are rigidly connected to the pinion rack-cutter and the pinion, respectively [Fig. 17.4.1(a)]. The fixed coordinate system Sm is rigidly connected to the cutting machine. (ii) The rack-cutter and the pinion perform related motions, as shown in Fig. 17.4.1(a), where sc = r p1\u03c8\u03c3 is the displacement of the rack-cutter in its translational motion, and \u03c8\u03c3 is the angle of rotation of the pinion. (iii) A family of rack-cutter surfaces is generated in coordinate system S\u03c3 and is determined by the matrix equation r\u03c3 (uc , \u03b8c , \u03c8\u03c3 ) = M\u03c3c (\u03c8\u03c3 )rc (uc , \u03b8c ). (17.4.1) Cambridge Books Online \u00a9 Cambridge University Press, 2009available at https://www.cambridge.org/core/t rms. https://doi.org/10.1017/CBO9780511547126.019 Downloaded from https://www.cambridge", + "4.2) Equation fc\u03c3 = 0 may be determined applying one of two approaches (see Section 6.1): (a) The common normal to surfaces c and \u03c3 at their line of tangency must pass through the instantaneous axis of rotation P1\u2013P2 [Fig. 17.2.1(a)]. (b) The second approach is based on the following equation of meshing: N(c) c \u00b7 v(c\u03c3 ) c = 0. (17.4.3) Here, N(c) c is the normal to c represented in Sc , and v(c\u03c3 ) c is the relative velocity represented in Sc . The schematic of generation of 2 is represented in Fig. 17.4.1(b). Surface 2 is represented by the following two equations considered simultaneously: r2(ut , \u03b8t , \u03c82) = M2t (\u03c82)rt (ut , \u03b8t ) (17.4.4) ft2(ut , \u03b8t , \u03c82) = 0. (17.4.5) Here, vector equation rt (ut , \u03b8t ) represents the gear rack-cutter surface t ; (ut , \u03b8t ) are the surface parameters of t ; matrix M2t (\u03c82) represents the coordinate transformation Cambridge Books Online \u00a9 Cambridge University Press, 2009available at https://www.cambridge.org/core/t rms. https://doi.org/10.1017/CBO9780511547126" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000881_robot.2004.1308085-Figure3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000881_robot.2004.1308085-Figure3-1.png", + "caption": "Fig. 3. RRRobot on a plane.", + "texts": [ + " [4] for mechanical systems with constraints. Kinematic reduction is useful because the kinematic reduced system is easier to control using velocity inputs than the unreduced dynamic system using , acceleration inputs (see [4] for details on controllability properties for reducible systems). 11. LEGLESS LOCOMOTION MODELS A. RRRobot on a plane We begin studying legless locomotion by exploring RRRobot on a plane. The RRRobot-on-a-plane model is a hemispherical shell with two short actuated legs (see Fig. 3). The massless shell has radius r, and the massless legs have length 1. There are five masses on the robot indicated by black dots: a mass at the distal end of each leg (Afi) ,~a mass where each leg is pinned (Ms) , and a mass at the bottom of the shell ( M b ) . Torques T~ and i-2 may be applied at the leg joints, and the shell rolls on the plane without slip. The configuration of RRRobot on a plane q consists of the sphere's position and orientation (z ,y ,R) with respect to a spatial frame and the internal configuration of its legs (+1,+2) " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003893_iceee.2010.5661528-Figure2.2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003893_iceee.2010.5661528-Figure2.2-1.png", + "caption": "Figure 2.2. Winding element figure after imposing current", + "texts": [ + " This motor\u2019s structure is double stators and single rotor. Because the magnet path of this motor is symmetrical, so the model is one half of this motor for simplifying calculation. The three dimensional model of stator, winding, and permanent magnet (PM) is shown in Fig.2.1. And the winding element figure which is This work was supported in part by Outstanding Young Teachers Project of Shanghai Municipal Education Commission(sdj09014 ,sdj09A111) 978-1-4244-7161-4/10/$26.00 \u00a92010 IEEE imposed the electric current density is showed in Fig.2.2. And the magnet density distribution under the same pole and different radius is shown in Fig.2.3. And the magnet density circumference distribution on the stator teeth is also showed in Fig.2.4. The number and magnitude of basic wave and some harmonic waves are shown in Table ,which change alone with time and position in period, and which cause the vibration and noise of motor. The electromagnetic vibration and noise of this motor are calculated by using FEM. And the acoustic field analysis of motor is a coupling field of fluent and structure" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002065_ecc.2007.7068339-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002065_ecc.2007.7068339-Figure2-1.png", + "caption": "Fig. 2. Joint frame F j and Plate frame F p.", + "texts": [ + " Furthermore, xe is pointing toward the satellite and ye and ze are such vectors perpendicular to xe. They can also be aligned to yb and zb with two rotations of Fe around yb and zb axes. More details are shown in Fig. 1. \u2022 Joint frame F j describes the movement of the azimuth motor with respect to the frame Fb. The origin of F j is placed on the antenna body. The vector z j is always aligned with zb. The vectors x j and y j are initially aligned with the vectors xb and yb respectively, but they change due to the rotation of the azimuth motor which rotate on zb axis. Fig. 2 illustrates this frame on the antenna body. \u2022 Plate frame F p describes the movement of the elevation motor with respect to F j. The origin of F p is placed on the plate of the antenna. The vector yp is always aligned with the vector y j and F p is rotating around yp due to the rotation of elevation motor. The axis xp is perpendicular to the plate of the antenna and yp. Finally, zp is perpendicular to both xp and yp as shown in Fig. 2. Now, we should understand how these frames move with respect to each other. First of all, we cancel all translative motions and distances between the origins of frames and only analyze the rotations between the frames. It\u2019s due to the fact that we assume the movement of the ship is negligible in respect of the distance between the satellite and the ship. Therefore, we begin from Earth frame Fe which is fixed on earth and then link other frames to each other one after another. \u2022 Rotation between frame Fb and frame Fe is originated from the waves and winds effecting the dynamics of ship motion", + " This rotation matrix has the property of R\u0307eb = RebSkew(\u03c9b eb), (1) where the entries of \u03c9b eb are the coordinates of the angular velocity vector of Fb relative to Fe, resolved in Fb [10]. \u03c9b eb = [ p q r ]T , (2) where p, q, and r are pitch, roll, and yaw angular velocities respectively. Also, the function Skew is defined as Skew(\u03c9b eb) = \u23a1 \u23a3 0 \u2212r q r 0 \u2212p \u2212q p 0 \u23a4 \u23a6 . (3) These disturbances are illustrated in Fig. 1. \u2022 The only change between Fb and F j is caused by the azimuth motor which is installed on the base. This rotation can be computed directly from the rotation angle of the motor. Consider \u03b8 j b j as the rotation angle of the azimuth motor as shown in Fig. 2. We can represent this rotation matrix by Rb j = \u23a1 \u23a2\u23a3 cos(\u03b8 j b j) \u2212sin(\u03b8 j b j) 0 sin(\u03b8 j b j) cos(\u03b8 j b j) 0 0 0 1 \u23a4 \u23a5\u23a6 . (4) \u2022 The rotation between F j and F p is caused by the elevation motor which is installed on the plate around y j axis. This rotation can be computed directly from the rotation angle of the elevation motor as illustrated in Fig. 2. Consider \u03b8 p jp as the rotation angle of the rotor of elevation motor. We can represent the rotation matrix as R jp = \u23a1 \u23a3 cos(\u03b8 p jp) 0 \u2212sin(\u03b8 p jp) 0 1 0 sin(\u03b8 p jp) 0 cos(\u03b8 p jp) \u23a4 \u23a6 . (5) The dynamics of the motors and kinematics of the antenna has been analyzed in [9]. Due to the inherent characteristics of the employed motors, which are step-motors, their dynamics are represented by simple equations \u03b8\u0307 j b j = u j b j (6) and \u03b8\u0307 p jp = up jp, (7) where u j b j and up jp are the inputs of the azimuth and elevation motors respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002985_978-0-8176-4558-8_14-Figure14.3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002985_978-0-8176-4558-8_14-Figure14.3-1.png", + "caption": "Fig. 14.3. Mechanical collision scheme. Interaction between two entities can lead to alignment in both models, while the resulting moving direction can be different in the models.", + "texts": [ + " Starting from a random initial condition and provided bacteria are sufficiently elongated, the system evolves, in both models, towards a steady state in which the microorganisms move in swarms. The length-to-width aspect ratio of bacteria \u03ba turns out to be a key parameter that controls the level of clustering in the system for a given density of cells.8 In what follows we discuss in more detail how the \u201cmicroscopic\u201d rules in both models lead to the emergent \u201cmacroscopic\u201d patterns observed in the simulations. Fig. 14.3 illustrates how in both models the local interaction between two individuals can cause local alignment. The active directed movement of the two cells plus volume exclusion force bacteria to become locally aligned and to point in the same direction. This local arrangement of cells lasts for some characteristic time which depends mainly on the magnitude of the active force and the length of the particles. During this period a two-bacteria cluster can eventually incorporate a third bacterium upon a similar collision process" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000391_2000-01-0920-Figure8-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000391_2000-01-0920-Figure8-1.png", + "caption": "Figure 8. Coordinate transformation", + "texts": [ + " The parameter ring twist and piston tilt greatly affect the shape of the piston ring during breaking in. Figure 7 displays a laterally buckled piston ring due to piston tilt. Since the piston ring is moving relative to the piston, every motion of the piston directly effects the motion of the piston rings. Assuming that the piston tilts around its piston pin with the angle \u03d5, one can calculate the angle that the piston ring is tilted against the cylinder liner by a simple coordinate transformation as shown in Figure 8. Since the program uses a moving Cartesian coordinate system around the piston ring, the spin vector (8) can be described in the piston ring coordinate system as function of the circumferential position, \u03b3 response. (9) (10) since \u03d5v and \u03d5u must apply for the relationship (11) because it is a cartesian coordinate system. Rz Rz1 Rz2+= \u03b4V ktot L H --- \u03b4s\u22c5 \u22c5= ktot kadhesive kabrasive kchemical+ += \u03d5 \u03d5v \u03d5 \u03d5sin\u22c5= \u03d5v \u03d5 \u03d5cos\u22c5= \u03d5v 2 \u03d5u 2 \u03d52 =+ 5 Ultimately the flow chart of the wear program can be explained as in Figure 10" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000544_jjap.40.4626-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000544_jjap.40.4626-Figure1-1.png", + "caption": "Fig. 1. Schematic of the experimental setup. P: polarizer, PBS: pellicle beamsplitter, EOX: EO crystal (ZnTe), QWP: quarter-wave plate, WoP: Wollaston prism, and PD: photodiode.", + "texts": [ + " Aperture dimensions of this size can significantly effect the THz pulse because of the large portion of low-frequency components in the THz pulse. In this study, a conductive waveguide model was constructed to simulate the propagated waveforms. The features in the experimentally observed waveforms were reproduced in simulations using this model. Experimentally, almost half-cycle pulses were generated by carefully avoiding the waveguide effect. Temporal field waveforms of focused THz pulses were observed by the electrooptic (EO) sampling method9) with the experimental setup shown in Fig. 1. In our experiment, THz pulses were generated by a large-aperture photocon- E-mail address: tukamoto@laserlab.bk.tsukuba.ac.jp yE-mail address: hattori@bk.tsukuba.ac.jp Jpn. J. Appl. Phys. Vol. 42 (2003) pp. 1609\u20131613 Part 1, No. 4A, April 2003 #2003 The Japan Society of Applied Physics 1609 ductive antenna. The antenna was constructed by directly placing two aluminum electrodes onto a h100i nondoped semi-insulating GaAs wafer surface with an intergap spacing of 30mm. The thickness of the GaAs wafer was 350 mm, and the diameter was 50mm" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003299_12.859448-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003299_12.859448-Figure2-1.png", + "caption": "Figure 2. Biaxial shape memory effect of a monodomain SmC* elastomer [21]. (a-1) a photograph and (a-2) an X-ray pattern were taken at room temperature. (b-1) a photograph and (b-2) an X-ray pattern were taken in SmA at 90\u00b0C. Corresponding models were illustrated in (a-3) and (b-3).", + "texts": [ + " The elastomer obtained shows the following phase sequence; g -6 SmX* (SmF*) 32 SmC* 80 SmA 115 I (in \u00b0C) The transition temperatures listed above were confirmed by differential calorimetry (DSC) and temperature-dependent X-ray investigations [16]. To investigate deformation behavior of the monodomain SmC* elastomer during the successive phase transitions, the shape change of the elastomer film is observed in a cooling and heating process. While the shape of the elastomer is rhomboid at room temperature in the tilted phase (Figure 2(a-1)), it transforms into a nearly rectangular-like shape in the temperature region (90\u00baC) of the SmA phase (Figure 2(b-1)). It is noteworthy that the reverse deformation takes place in cooling from the SmA to SmC* phase. Namely, the rectangular-like film spontaneously transforms into a rhombic one on decreasing the temperature from the SmA phase to the SmC* phase. To analyze whether the macroscopic shape changes directly correlate with molecular re-alignment processes, X-ray investigations are carried out. The X-ray pattern of the tilted smectic phase and that of the SmA phase are shown in Figure 2(a-2) and (b-2), respectively. In the X-ray Proc. of SPIE Vol. 7775 777509-2 Downloaded From: http://proceedings.spiedigitallibrary.org/ on 06/21/2016 Terms of Use: http://spiedigitallibrary.org/ss/TermsOfUse.aspx pattern of the tilted smectic phase at room temperature shown in Figure 2(a-2), reflection located near the meridian (arrow 1) indicates a uniform alignment of smectic layers, while reflection at wide angle (arrow 2) indicates that the mesogenic groups are aligned uniformly in the direction inclined as \u03b8X with respect to the layer normal. On the other hand, an orthogonal molecular alignment associated with the SmA phase is confirmed in Figure 2(b-2). Comparing both X-ray patterns of Figure 2(a-2) and (b-2), we recognize that the location of wide-angle reflection rotates clockwise with increasing temperature, while the layer reflection remains near meridian. Namely, the mesogens inclined with respect to the layer normal in Figure 2(a-2) are re-arranged in the direction parallel to the layer normal in Figure 2(b-2), although the direction of the layer normal is almost independent of temperature. In the cooling process the reverse change of the molecular alignment is confirmed; the location of the wide-angle reflection rotates anti-clockwise with decreasing temperature, and then it reverts to the original position at room temperature. Since the residual tilted order originating from cross-links commands mesogens to tilt in the direction of the original orientation, they accordingly tilt in the same direction during the SmA-SmC* phase transformation. As the internal-shear-stress originating from cross-links induces the macroscopically uniform molecular-tilting [26], the monodomain SmC* elastomer is to hold the memory of its original alignment. According to the correspondence of the sample observation to the X-ray scattering pattern, we are able to schematize the relationship between the shape change of the elastomer film and the molecular re-arrangement during SmC*-SmA phase transformation in Figure 2(a-3) and (b-3) where the shape of elastomer film is closely associated with the local symmetry Reversible transformation \u03b8E (a-1) LE (a-2) 1 2 90 - \u03b8X (b-1) (b-2) Proc. of SPIE Vol. 7775 777509-3 Downloaded From: http://proceedings.spiedigitallibrary.org/ on 06/21/2016 Terms of Use: http://spiedigitallibrary.org/ss/TermsOfUse.aspx of the liquid-crystalline phase. It has to be emphasized that the tilt angle of the elastomer film \u03b8E, which is defined as the angle between of the edge of the film and the direction of the first uniaxial deformation (see Figure 2(a-1)), approximately agrees with the molecular tilt angle \u03b8X characterized in the X-ray patterns (see Figure 2(a-2)). Namely, the macroscopic symmetry defined by the shape of the SmC* elastomer film corresponds directly with the microscopic local symmetry due to the molecular alignment of the liquid-crystalline phase of SmC*, as depicted in the schematic model of Figure 2(a-3) and (b-3). We can conclude that cross-linking between polymer backbones under mechanical shear field produces the monodomain SmC* elastomer in which a memory of the biaxial molecular alignment of the SmC* state is perpetuated. Moreover, we have confirmed the biaxial memory is perpetuated even after exposure to the isotropic state [21]. Since it is expected that the monodomain SmC* elastomer exhibits physical properties, such as ferroelectricity and piezoelectricity, due to the macroscopic C2 symmetry of the unwound SmC* state, in following subsections, we present experimental results of the piezoelectric and inverse-piezoelectric properties of the SmC* elastomers" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002034_2007-01-2860-Figure18-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002034_2007-01-2860-Figure18-1.png", + "caption": "Figure 18 \u2013 Moment (Mtz) generated by resulting forces F1 and F2, due to size difference of the profile borders", + "texts": [ + " The twist movement causes the crosspiece to bend around point Cc, if the inertia of the inferior profile border relative to point Cc is very low when compared to the inertia of the superior border, as in Figure 17. This flexibility difference generates bigger Ydisplacement of the inferior border, forcing positive toe during compression. Therefore, it can be concluded that the greater the proportion between border inertias is, the greater will be the trend to produce toe during compression. 7 3 \u2013 Relation between C-profile border size: As illustrated in Figure 18, the size difference between the borders generates a moment that will direct the trailing arm to toe or divergence movement during compression. e) Dynamic variation of Iz inertia It is important to evaluate variation of Iz inertia in function of profile twist because with profile deformation there can be alteration in Ix and Iz values. Low Iz values produce greater toe variation. f) Force component in Y direction The greater the camber angle is, the greater will be the Y-force component of the ground, i" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002819_6.2007-3813-Figure27-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002819_6.2007-3813-Figure27-1.png", + "caption": "Figure 27. Rectangular pin in Tap 14.", + "texts": [ + " At some point the jet becomes too strong to remain attached to the body and thus the benefits of tangential blowing begins to diminish as is evidenced by the pressure contours in Figure 26 and the data in Figure 25. The fact that there were only limited changes in the force at lower tunnel Mach numbers may also be due to the fact that the separation was not as strong at these lower Mach numbers. As described above, both a round and a rectangular pin were manufactured that could be inserted into any of the pressure tap locations. The pressures generated on the model were measured when the rectangular pin was inserted into Tap 14 for various tunnel Mach numbers as shown in Figure 27. As anticipated when the pins were located in Page 14 the cavity region of the projectile model, the results were similar to that of the normal jets which included an upstream increase in pressure and downstream regions of decreasing pressure immediately behind the pin followed further downstream by increases in pressure. These pressure changes were found to increase in both magnitude and in spatial separation with increasing tunnel Mach number as shown in Figure 27. Upon comparing the round and the rectangular pin and the normal jet at the Tap 14 location, Figure 28, it is apparent that the rectangular pin most strongly influences the pressure distribution on the projectile body. For the round pin this is due to the fact that the flow is disturbed more by the rectangular pin as the round pin provides a 3-D relieving effect. Both of the pins provide a stronger pressure change than the jet as they provide more flow blockage as the pin diameter is roughly three times that of the jet diameter" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003904_6.2009-4887-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003904_6.2009-4887-Figure2-1.png", + "caption": "Figure 2. Finite Element Model Component Layout", + "texts": [ + " Both techniques have been used with comparable results. The results presented here use the more rigorous mesh matrix development. The epicyclic stage of the transmission is more complex due to its multiple components and the orbital motion of the planets. For convenience, the system model implements an existing lumped parameter epicyclic model from the literature45, 46. Each subcomponent is placed at its appropriate shaft node in the model as discussed previously. American Institute of Aeronautics and Astronautics 4 Figure 2 presents a comprehensive illustration of the transmission model by subcomponent. The nodes on the three shafts each have the conventional six dof\u2019s. The nodes of each shaft were placed at key locations as dictated by shaft geometries, gear-mesh, and bearing locations. Shafts A and B each have eight nodes. Shaft C consists of 13 nodes. The total number of dof\u2019s for the shafts is 174. The epicyclic components each have three dof\u2019s \u2013 two in translation and one in rotation (torsional). The ring and planets therefore total 3+3n dof\u2019s, where n is the number of planet gears" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001183_20050703-6-cz-1902.01279-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001183_20050703-6-cz-1902.01279-Figure1-1.png", + "caption": "Fig. 1. Correlation between the robot\u2019s position and the point in the configuration space", + "texts": [ + " THE CONFIGURATION SPACE The main idea of the configuration space is to represent the robot as a single point in an appropriate space and to map the obstacles in this space. By this the problem of motion planning for a dimensioned body is transformed into the problem of planning the motion of a single point. A position of a manipulator robot can be completely described using the values of all joints of the robot at a particular robot position. This list of joint values is called a configuration. All possible configurations build the configuration space. The number of joints of the robot is equal to the dimension of the configuration space. Figure 1 demonstrates the correlation between a robot\u2019s position and the corresponding point in the configuration space for a planar manipulator robot with two degrees of freedom. In the configuration space all configurations which lead to a collision between the robot and at least one obstacle in its environment can be marked. This can be done pointwise by moving the robot in the simulation to every possible configuration and performing a collision check at this position. Thereby the obstacles are mapped from the workspace into the configuration space" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002979_978-3-540-70534-5_7-Figure8.1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002979_978-3-540-70534-5_7-Figure8.1-1.png", + "caption": "Figure 8.1: Driving and rotation of single wheel drive", + "texts": [ + "1 Single Wheel Drive Having a single wheel that is both driven and steered is the simplest conceptual design for a mobile robot. This design also requires two passive caster wheels in the back, since three contact points are always required. Linear velocity and angular velocity of the robot are completely decoupled. So for driving straight, the front wheel is positioned in the middle position and driven at the desired speed. For driving in a curve, the wheel is positioned at an angle matching the desired curve. U Driving Robots 132 8 Figure 8.1 shows the driving action for different steering settings. Curve driving is following the arc of a circle; however, this robot design cannot turn on the spot. With the front wheel set to 90\u00b0 the robot will rotate about the midpoint between the two caster wheels (see Figure 8.1, right). So the minimum turning radius is the distance between the front wheel and midpoint of the back wheels. 8.2 Differential Drive The differential drive design has two motors mounted in fixed positions on the left and right side of the robot, independently driving one wheel each. Since three ground contact points are necessary, this design requires one or two additional passive caster wheels or sliders, depending on the location of the driven wheels. Differential drive is mechanically simpler than the single wheel drive, because it does not require rotation of a driven axis" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003745_15502280902939486-Figure5-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003745_15502280902939486-Figure5-1.png", + "caption": "FIG. 5. Residual strains for different stress-strain curves.", + "texts": [ + " However, the behavior of low E/Y materials keeping more residual strain while unloading goes inline with the known fact that stress-strain relation during unloading is straight line and approximately parallel to the initial part of the stress-strain relation [20]. Figure 4 shows the stress-strain relation in a non-dimensional way provided by the authors in their previous work [17]. It gives the variation of the ratio between mean contact pressure and yield stress against the dimensionless interference during loading of a sphere of low E/Y value against a rigid flat. Figure 5 gives residual strains that could have been obtained from different stress-strain curves (like Figure 4) using the above reality that stress-strain relationship is linear while unloading. In this schema, AB and MN are parallel to initial elastic part of the curve. If the unloading process started from a particular strain level (point O), the material corresponding to curve 1 (low E/Y) gives a residual strain of EP1 and elastic strain Ee1. But material FIG. 4. Mean contact pressure/yield stress vs" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002080_ijcnn.2007.4371377-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002080_ijcnn.2007.4371377-Figure1-1.png", + "caption": "Fig. 1. The relationship between COG, ZMP and reaction force", + "texts": [ + " Its advantage is that the IPM could generate various patterns of humanoid motion. Goswami[8] had investigated different stability criteria for humanoid robots. It is shown that ZMP and COP (center of pressure) are equivalent while keeping motion stability. ZMP can be 1-4244-1 380-X/07/$25.00 \u00a92007 IEEE considered as the center of ground reaction force. Under assumption 3, the level reaction force is linear to the position of COG. Therefore, according to the definition of ZMP, the ZMP is linear to COG. The relationship between ZMP, COG and reaction force is depicted in Fig. 1. (Fr, F ) is reaction force. The position of COG is(X, H), the position of ZMP is (aX + b, 0), and the normal vector of ground force is parallel to vector (cX + d, H). The dynamics of the IPM can be given by FB :eFu zH,w+g) = (c+d):tH (1) Because z = H, we can get z = g(cx+ d)/H (2) A. IPM for Humanoid Motion in the Sagittal Plane One step of humanoid walking consists of two phases, single support and double support phase. The dynamics equation of the IPM in the sagittal plane is given by gs: (Zgs + g) Xgd (Zd +g) 5gs(c - ala) : H Xgd(Cxd a d) H where xgs and Xgd are the trajectories of the IPM's COG during the single support phase and double support phase respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003608_ipin.2010.5647715-Figure4-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003608_ipin.2010.5647715-Figure4-1.png", + "caption": "Figure 4. Layout of the measurement setup.", + "texts": [ + " The center of the profile was determined by measuring a cone around the profile and calculating its axis. The routine has to be repeated for each robot individually. For determining the synchronization (i.e. \u2013 possibly variable \u2013 latency) between the two robots triggered measurements to both robots at the same time were carried out. Each tracker observed one robot. The robots\u2019 tools moved along a 500 mm long straight line parallel to the Y-axis of the coordinate system. The measurement setup is depicted in Fig. 4 with R1 and R2 being the tools of the two robots. One measurement cycle contained the movement from 0 to 500 mm, a short stop and the movement back to 0 mm. To ensure fully synchronized data, the robot control unit provided a trigger signal for the trackers with a frequency of 500 Hz, i.e. the trackers measured a point every two milliseconds. The trigger signal ensured the trackers to start the measurement at the same point of time without any delay. This is crucial for investigating the synchronization of two objects, in this case, the two robots" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001166_1-84628-559-3_7-Figure3.2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001166_1-84628-559-3_7-Figure3.2-1.png", + "caption": "Fig. 3.2. Geometry measurement using fringe projection.", + "texts": [ + "1) offer the potential to fulfil all these demands. Especially methods based on the triangulation principle promise a good feasibility concerning accessibility of gear flanks, speed and accuracy, whereas interferometric measurements are supposed to fail, mainly due to a minor accessibility [21, 22]. Namely stripe pattern projection gives access to both, a high and complete information content (areal recording of a whole flank at once) and a sufficient accuracy (coded light approach [4, 23-26, 32, 33] and phase shift procedure [24, 27-29]). Figure 3.2 illustrates the triangulation principle and the basic set-up, where a well-defined black and white pattern (mostly parallel stripes) illuminates the curved measuring object. A camera (CCD or CMOS) observes the projected pattern at a defined (and previously verified) triangulation angle, extracting the object\u2019s 3D-surface coordinates from the distortion of the recorded pattern, caused by the sculptured surface. The main advantages of this measuring principle are simple and fast implementation, stripe density and phase shift adaptable to the measuring problem, measuring area from mm2 to m2, depth resolution down to 1/10.000 of diagonal of measuring area possible, improved cameras and image processing directly applicable One major difficulty occurs on evaluating the registered pattern, caused by the so-called \u201corder problem\u201d [23, 24, 33]: one certain camera pixel cannot identify the number (order) of an observed stripe period, leading to an ambiguity in the detected distance d between the inspected surface point and the corresponding camera pixel (Figure 3.2). An elegant procedure labelled \u201ccoded light approach\u201d avoids this problem [24], where a series of pattern (e.g. 8) with varying period lengths illuminate the (static) object surface. Thus, each pixel \u201csees\u201d a sequence of dark and light spot intensities, transformed into a binary code. Moreover, phase shifting of the \u201cfinest\u201d stripe pattern further increases the depth resolution of this measuring principle. [30] reports another method for identifying the order of a certain stripe by varying the pattern" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001758_detc2005-84336-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001758_detc2005-84336-Figure2-1.png", + "caption": "Figure 2. Experimental Set-up", + "texts": [ + "1 shows the stresses for the strain of the upper row in Fig.1 for 0 q 1, c = 1; = 1. Two type experiments have been conducted. (1) Type1 experiment ; A viscoelastic test specimen is subjected to Type1displacement in this experiment. (2) Type2 experiment ; The viscoelastic test specimen is subjected to Type2-displacement in this experiment. Some results of the experiment were already reported by Nasuno-Shimizu (Nasuno and Shimizu, 2004), but for readers\u2019 convenience, brief description of the experiment will be given below in this chapter. Figure 2 shows a hydraulic-servo type actuator which was used for the experiment. Multi-layered cylindrical acrylic viscoelastic body (Type SD112 of Sumitomo 3M Co.) which is shown in the figure was used as the test specimen. The size of the specimen is = 60mm in diameter and h = 27mm(1mm 27 layers) in height. Adhesion of the upper and the lower surfaces of the viscoelastic body to the attachments of the test apparatus was firmly done by using the ability of self adhesion of the material after processing the surface treatment of the attachments so that the test specimen does not peal off or slip from the surfaces of the attachments" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000363_0094-114x(78)90058-7-Figure4-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000363_0094-114x(78)90058-7-Figure4-1.png", + "caption": "Figure 4.", + "texts": [ + " For related reasons, we shall also use the notation J = J to indicate that two joints are parallel; in such a case, it may be inferred that the joints act in their normal manner, and that nothing more than parallelism is indicated. Both of these notations were employed to advantage in [4]. Some weight is added to the consideration of the H =H combination as a single joint by comparing the independent closure equations of, for example, the P-'P-, H = H ~ - P - , H = H \u00b1 H = H - sequence of linkages. The following equations are easily obtained by substituting the appropriate dimensional constraints into the 2-, 3- and 4-bar sets of closure equations. The symbols employed in the equations are illustrated in Fig. 4, which depicts a hypothetical, general 4-bar linkage. P~-P - : rl + r2 = 0 (i) with cOj = c02 = - 1 H = H ~ - P - : R l + R2 + hlOl + h202 + r3 = O (i) 01 + 05 = (2k + l)rr (ii) with c03 = - 1 H = H ~ - H = H - : RI + R2+ R3+ R4+ hlO! + h202+ h303+ h404 = 0 01 + 02 = (2k + 1)zr 03+ 04 = (2/+ l)lr (i) (ii) (iii) The number of independent closure equations in each case is, of course, one less than the connectivity sum of the corresponding linkage. For the latter two linkages, however, we can restructure the closure equations and dimensional constraints so that they are directly comparable with those of the first linkage" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001915_0284185172012s31332-Figure4-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001915_0284185172012s31332-Figure4-1.png", + "caption": "Fig. 4. Elevation of detectors in system 2.", + "texts": [], + "surrounding_texts": [ + "Ventilation of the steel rooms is accomplished by a fan capable of eight com plete air changes per hour. The incoming air is filtered through a so-called abso lute filter removing all particles above 1 / 1 6ih 3 = 4r > 3. When Xi and Mi satisfying LMIs (17) and (18) are found, the feedback gain F 5 \u03c4 6 is given by F 5 \u03c4 6i=kj r \u2211 i K 1 \u00b5i 5 \u03c4 6 Mi l j r \u2211 i K 1 \u00b5i 5 \u03c4 6 Xi l 9 1 (19) This section presents a numerical simulation result where the proposed GS-SF design technique was applied to a tracking problem of a two-link robot arm shown in Fig. 2 (Palm, et al., 1997; Hsieh, et al., 2001). The masses of the two links were concentrated at the ends and the motor inertia are neglected. The equation of motion of the two-link robot arm was described as M m x n x\u0307 m t n/o N m x n p Lu m t n (20) where x qpsr q1 q2 q\u03071 q\u03072 t T was the state vector including the angles and the angular velocities of the two links, u qpkr u1 u u2 t T was a torque vector. M m x n included the inertia matrix and N m x n was a vector including centrifugal, Coriolis, gravitational forces and damping" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001809_2005-01-0384-Figure3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001809_2005-01-0384-Figure3-1.png", + "caption": "Fig. 3 FE Model of a Sedan Type Vehicle", + "texts": [ + " 2 that the calculated coefficient of friction of the cornering force 2005-01-0384 Vehicle Cornering and Braking Behavior Simulation Using a Finite Element Method Tatsuya Fukushima, Hitoshi Shimonishi and Toshikazu Torigaki Nissan Motor Co., Ltd. Takahiko Miyachi and Yasuyoshi Umezu The Japan Research Institute, Limited showed good agreement with the measured data for both small and large slip angles and for both small and large belt loads. A sedan type vehicle was modeled with finite elements as shown in Fig. 3. The link parts of the front and rear suspensions, suspension members, as shown in Fig. 4, and the vehicle body were modeled as rigid bodies, because deformation of a running vehicle was regarded as being smaller than that of the tires and kinetic motion of the suspension systems. The beam elements in Fig. 3 were grouped into one rigid body to express the total inertia of the body, engine, power-train unit, parts in the cabin and the other parts mounted on the body suspended by the suspension systems. To the vehicle model shown in Fig. 3 the outer skin panels of the body were mounted to the group of beam elements to visualize the behavior of the vehicle. The outer skin panels had no inertia in the simulation. The suspension springs, shock absorbers and bump rubbers were modeled with discrete elements and the bushings of the suspensions and insulators were modeled with beam elements having six local degrees of freedom. The properties of stiffness and damping of these parts were defined as functions of the measured values. Before attempting to simulate the dynamic behavior of the vehicle, we compared the properties of the suspension systems with the measured results", + " Figure 5 shows that the calculated load properties of the suspension stroke agreed well with the measured results. Figure 6 shows that the calculated load properties of body roll also coincided with the measured data. The rack of the rack and pinion steering system was modeled with a translational motor joint for steering the front wheels in cornering behavior simulations. Simulation of vehicle cornering behavior Vehicle cornering behavior was simulated on a flat, dry asphalt road surface with the vehicle model shown in Fig. 3. The vehicle speed before cornering was 19.44 m/sec (70 km/h). The transmission was shifted to the neutral position just before steering the vehicle so as to cut off the transfer of driving torque to the wheels. Three steering conditions were simulated and are denoted as Test A, Test B and Test C in Fig. 7, which shows the displacement of the rack. In these three tests, driving torque and braking torque were not transmitted to the wheels. However, it is not easy for a test driver to reproduce the steering profiles shown in Fig", + " The three test conditions A, B and C were simulated using the measured vehicle speed as the initial velocity and the measured steering profile as the boundary condition. Table 1 gives the measured initial velocity in each test, and Fig. 8 shows the measured rack displacement. The data in Table 1 and Fig. 8 were the inputs used in the vehicle cornering simulations conducted in this study. Table 1 Initial Velocity in Cornering Simulations Fig. 8 Measured Rack Displacement Simulation of vehicle braking behavior Vehicle braking behavior was also simulated on a flat, dry asphalt road surface using the FE tire model shown in Fig. 1 and the vehicle model shown in Fig. 3. The vehicle speed and braking torque applied to the wheels were first measured in a driving test. The measured vehicle speed just before braking was input as the initial velocity and the braking torque as the load condition of the four wheels. Figure 9 shows the measured braking torque used in the braking simulation. In this measurement, the driver applied a steady force to the brake pedal. However, braking torque fluctuations at the wheels are observed in Fig. 9. Those fluctuations were caused by the operation of the anti-lock braking system, and they indicate that tire slip occurred intermittently", + " The maximum acceleration in Test C was close to the upper limit of acceleration of the vehicle on the dry asphalt road surface used in the tests. The maximum acceleration calculated under both conditions, including the extreme condition, showed good agreement with the driving test data. Figure 19 shows the peak angular velocity at the vehicle\u2019s center of gravity in Tests A, B and C. It is seen in the figure that the calculated maximum angular velocity agreed well with the measured values. Simulation of vehicle braking behavior Vehicle braking behavior was simulated with the FE tire model shown in Fig. 1, the vehicle model shown in Fig. 3 and the measured braking torque applied to the wheels in Fig. 9. Figure 20 shows simulated vehicle attitudes before braking and during braking, and it is seen that the nose of the body dives due to braking. Figure 21 shows the calculated and measured longitudinal acceleration at the vehicle\u2019s center of gravity. The calculated time history of longitudinal acceleration shows good agreement with the driving test results. A sedan type vehicle was modeled using finite elements to represent the tires and kinematics of the front and rear suspension systems in order to simulate the dynamic behavior of the vehicle" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000694_b:tels.0000029042.75697.f0-Figure6-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000694_b:tels.0000029042.75697.f0-Figure6-1.png", + "caption": "Figure 6. (a) Principal point, image and sensor axes. (b) Observation axis relative to sensor and world axis.", + "texts": [ + " Denoting the position the sensor in the body frame as xb s , the directional cosine matrices relating the body to earth frame as Ce b and the sensor to body frame as Cb s , the position of the target in the earth frame is given by ze t = xe b + Ce bxb s + Ce bCb s zs t . (14) The associated uncertainty of the observation is required in the form of the covariance matrix R. An approximation converted to Cartesian space relative to the axis defined by the observation itself is Rt t = \u03c3 2 r 0 0 0 r2\u03c3 2 \u03c8 0 0 0 r2\u03c3 2 \u03b8 . (15) Figure 6(a) shows more explicitly how this approximation relates geometrically in 2D to the world and sensor axes. The covariance of the target in the earth frame can be calculated using the direction cosine matrices relating the target to sensor, sensor to body and body to earth frames. Re t (sensor) = [ Ce bCb s Cs t ] Rt t [ Ce bCb s Cs t ]T . (16) Additional errors induced by the uncertainty in the vehicle body location, roll, pitch and yaw must also be accounted for. Following flight tests it became apparent that these errors are the major contributors to the uncertainty of the observation" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001600_j.2042-7158.1971.tb08790.x-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001600_j.2042-7158.1971.tb08790.x-Figure1-1.png", + "caption": "FIG. 1 . Light passes through the condenser lens and is polarized by the Polaroid filter A. It passes through the die and the second Polaroid B to the camera. Upper and lower punch pressures are measured by the instrumented pressure rollers, the values being recorded on the oscilloscope or the chart recorder. The camera and oscilloscope are triggered by the contact wire, with timing marks being derived from the stroboscope and an initiating flash from the flash bulb.", + "texts": [ + " We have been able to make similar calculations, and have used the energy input to calculate the expected temperature rise for comparison with the results of Travers & Merriman (1970). Their method, implanting a thermocouple in a tablet during compression, cannot be extended easily for measurements on a rotary machine. The work done by a punch during compression is M A T E R I A L S A N D M E T H O D S A Manesty D3 rotary tabletting machine, capable of producing 500 tablets per minute in normal operation, was modified to enable the required measurements to be made. The general arrangement of the measuring equipment fitted to the machine is shown in Fig. 1 (for fuller details of technique and results, cf. Rosser, 1970). Fifteen of the stations were blanked off, and the Perspex die fitted at the remaining station on the die table with its associated shortened upper punch and lengthened lower punch. Brackets were bolted to the machine frame to carry the equipment for photoelastic measurement. Light from an Atlas 100 W projector bulb was passed through a 10cm diameter condenser lens, focal length 25 cm, to give a parallel beam. This beam then passed through a Polaroid filter, a first quarter-wave plate, the Perspex die, second quarter-wave plate, analysing Polaroid and into the camera, a Hycam (Red Lake Laboratories, Inc" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002184_12.715863-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002184_12.715863-Figure1-1.png", + "caption": "Fig. 1. The test apparatus.", + "texts": [ + " of SPIE Vol. 6523, 652316, (2007) \u00b7 0277-786X/07/$18 \u00b7 doi: 10.1117/12.715863 Proc. of SPIE Vol. 6523 652316-1 Downloaded From: http://proceedings.spiedigitallibrary.org/ on 06/16/2016 Terms of Use: http://spiedigitallibrary.org/ss/TermsOfUse.aspx experimental data is used to identify the model parameters. The model is evaluated and a modification for more accuracy is proposed. The modified model is compared with the original model. 2. EXPERIMENTAL SETUP In this paper, the test apparatus shown in Fig. 1 is used to evaluate the load-dependent hysteresis model experimentally. Terfenol-D, an alloy of iron, terbium, and dysprosium is a magnetostrictive material. A Terfenol-D rod is used in the apparatus and hysteresis curves for different loads are obtained. In the actuation unit, a Terfenol-D rod is surrounded by an electrical magnet. The electrical magnet is powered by a programmable current source controlled by a PC computer. When there is an electrical current in the magnet, magnetic field (H) is produced. As a result, Terfenol-D rod becomes slightly longer. The vertical rod shown in Fig. 1 moves upward. An optical encoder measures the displacement. A pickup coil inside the actuation unit is used to measure flux density B. Two thermistors are included to measure temperature to avoid overheating. The measurements are fed to the PC computer running MATLAB Real-Time Workshop\u00ae. The results are recorded there. In this setup, a set of washer springs provide the force on the actuator. The springs are soft enough so that the force can be assumed constant. The force can be adjusted with a bolt on top of the setup and is measured with a load cell" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003730_20091021-3-jp-2009.00033-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003730_20091021-3-jp-2009.00033-Figure2-1.png", + "caption": "Fig. 2. Plant scheme", + "texts": [ + " The position of the ball acts when rolling as a wiper similar to a potentiometer. The angle of the servo motor is recorded by the position of the encoder. In order to obtain the mathematical model of the plant, let\u2019s begin by examining the forces acting on the ball. Those are a translational force due to gravity and a rotational force caused by the rotational acceleration of the ball. Then, if the Second Newton law is applied: . 5 2sinsin 2 xmmg R xJmgFFxm rxtx (1) Where it is been considered the angle alpha as shown in Figure 2. From (1), it is obtained: .sin 7 5 gx (2) When small oscillations are considered, sin . Then: 27 5 )( )( s g s sX (3) As far as the motor is concerned, the transfer function from the motor voltage to output angle is: sKKRBsRJ KK sV s gmmgmeqmeq gtmg )()( )( 22 (4) The motor parameters are defined in Table 1. Finally, the relationship between \u03b8 and \u03b1 can be obtained as follows: Expressed as a transfer function: r L s s )( )( (6) The goal of the system is to control the position of the ball to a desired set point along the beam" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001541_1.2362952-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001541_1.2362952-Figure2-1.png", + "caption": "Fig. 2. The side view of the apparatus. The arrow shows the direction of current. Current flows along the wire, ball, recording paper, and clip.", + "texts": [ + " The maximum voltage between electrodes is 110 V, and it reaches this value in 1/1000 s or less. The current is 0.1 A. There is no danger even if someone touches the surface of the recording paper that\u2019s in contact with an electrode. The wooden plate must be carefully leveled. Two ramps are set up at the end of the plate. Two identical steel balls (3 cm in diameter) are launched down the ramps simultaneously so that they collide with each other. When a ball moves on the paper, it always touches at least one wire (as shown in Fig. 2). The top layer of the recording paper is an insulator. By increasing the dc voltage, this layer is destroyed and the lower carbon layer is exposed. As the carbon layer is a conductor, current flows through the paper from the bottom of the ball to the clip. Then we can get easily seen black marks. The very small force that the wire exerts on the ball must be negligible. Theory The black marks show that each ball follows a slightly curved path for a very short time after the collision and then moves in a straight line, as shown in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002979_978-3-540-70534-5_7-Figure10.6-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002979_978-3-540-70534-5_7-Figure10.6-1.png", + "caption": "Figure 10.6: Double inverted pendulum robot", + "texts": [ + " This gives us the equivalent of a double inverted pendulum; however, with two independent legs controlled by two motors each, we can do more than balancing \u2013 we can walk. Dingo The double inverted pendulum robot Dingo is very close to a walking robot, but its movements are constrained in a 2D plane. All sideways motions can be ignored, since the robot has long, bar-shaped feet, which it must lift over each other. Since each foot has only a minimal contact area with the ground, the robot has to be constantly in motion to maintain balance. References 163 Figure 10.6 shows the robot schematics and the physical robot. The robot uses the same sensor equipment as BallyBot, namely an inclinometer and a gyroscope. 10.4 References CAUX, S., MATEO, E., ZAPATA, R. Balance of biped robots: special double-in- verted pendulum, IEEE International Conference on Systems, Man, and Cybernetics, 1998, pp. 3691-3696 (6) DEL GOBBO, D., NAPOLITANO, M., FAMOURI, P., INNOCENTI, M., Experimental application of extended Kalman filtering for sensor validation, IEEE Transactions on Control Systems Technology, vol" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000149_icsmc.1997.638144-Figure8-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000149_icsmc.1997.638144-Figure8-1.png", + "caption": "Figure 8 Secondary Motion of Proximal Phalanx", + "texts": [ + " The displacement of node P to P\u2019 is given by the weighted sum: , = I where s is the number of proximal phalanx flexion acting on node P and f, is a flexion displacement factor. After all the web tissue node displacements are computed, the new orientation of the other phalanx affected is given by direction of the vector B,, where: , = I P\u2019(e), is the displaced node at the edge of the mesh connected to the proximal phalanx, Ab, is the corresponding point of attachment at the base of the phalanx to the mesh and n is the number of those nodes. In Figure 8 on next page, the proximal phalanx in the right side is flexed outwards which causes secondary motion to the left proximal phalanx. Whenever proximal phalanx motion is involved at any simulation time step, the above computation is performed t o simulate the secondary motion appropriately. Figure 9 on next page shows the secondary motion of index and ring finger caused by flexion of proximal phalanx of middle finger. We used a hand tree to represent the computation association of each segment of the hand model" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001393_kem.291-292.483-Figure9-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001393_kem.291-292.483-Figure9-1.png", + "caption": "Fig. 9 Theoretical normal tooth section", + "texts": [ + " 8 shows the axial tooth profile of by theoretical analysis, enveloping machining method and NC machining method. It is clearly that the NC manufacturing method has higher machining accuracy than that of the enveloping method. The equation of normal tooth profile of the stationary internal gear can be denoted as [9], ( ) ( )T 333 3 3 T 333 1M1 zyxzyx p ppp = . (8) Combining equations (4) and (5), let 03 =pz in equation (8), the normal tooth section can be produced by giving the planet worm-gear rotating angle 2\u03d5 . Fig. 9 shows the theoretical results of the tooth profile normal section for the stationary internal toroidal gear with abovementioned parameters. This shape is proved to be a circle at the centre of the meshing rollers and to have the same radius with the rollers. Similarly, by using 3D-Measuring machine, the normal section tooth profile of the internal gear manufactured by enveloping method and by NC method can be also obtained. Fig. 10 shows the normal tooth profile of by theoretical analysis, enveloping machining method and NC machining method" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000875_amc.2004.1297953-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000875_amc.2004.1297953-Figure1-1.png", + "caption": "Fig. 1: Graphical interpretation of sliding surface for ' a second-order system", + "texts": [], + "surrounding_texts": [ + "I. INTRODUCTION\nShip handling in harbours is one of the most complicated and sophisticated types of ship manoeuvring and navigation. Low speed conditions, non-linear dynamics of ships and complicated harbour environment have all made the safety of manoeuvring in harbour areas of large ships an increasingly urgent issue. Therefore, from safety and economy points of view, it is an essential task to construct a controller to guide the ships in harbour manoeuvres.\nTo develop such a controller, several problems must be solved. The dynamics of ships in harbour manoeuvres are fundamentally non-linear in nature; therefore, to be able to describe ship manoeuvring motions in harbours, a suitable mathematical model is necessary. Several mathematical models for ships motions at low speed range have been proposed, however unfortunately, until now none of them have been seen as an ideal model. In this study, a multi-term mathematical model of ship motions is adopted in order to be able to use for a wide range of ship harbour manoeuvres. The model was originally presented by K. Kose el ai. [ I ] and has further been developed by Le and Kose [21, Le and Nguyen [3]. To control such a non-linear system, a robust controller is always desirable. The Sliding mode control is known as a non-linear robust control method [41 and has been successhlly applied in the control of Underwater Vehicles. Yoerger and Slotine [SI, [6], proposed the use of Sliding Methodology for non-linear and robust trajectory control of Underwater Vehicles. Healey and David [7] reported the successful implementation of the MSMC for Autonomous Diving and Steering of Unmanned Underwater Vehicles. Fossen [8] has furthermore extended the applications of The Sliding mode control into a wide range of Advanced Autopilots Design for Remotely Operated Vehicles. However, comparing to those objects, controlling of large ships in harbors is much more complicated and concerning studies\n0-7803-8300- 1/04/$20.00 02004 IEEE. 695\nshould be made in conditions similar to real aspects of harbor maneuvers.\nRecently, several studies concerning automatic control systems for ships' harbour manoeuvres have been carried out [8], [9], [IO] or [IO]. In most of those studies, bow and stem thrusters were used as the means to provide controlling forces and moment, while control of large ships in harbour areas usually involves the use of tugboats. This study applies the Sliding Mode Control to control ship position and anitude in harbour manoeuvres through the use of tugboats. This paper proves that a MSMC based on state feedback is effective to deal with the problems of controlling large ships in harbour manoeuvres. The influence of slow speed, modelling non-linearity and environmental disturbances can be compensated effectively.\nThis paper is organized as follow. Section II gives the general basics of the Sliding Mode Control. Application of the Sliding Mode Control for ship harbour manoeuvring, given a non-linear model of ships motions, is presented and discussed in Section 111. Using a non-linear ship model for a VLCC, the 180 deg. turning manoeuvre was carried out under various strong wind conditions. The simulation results are shown in Section IV to prove the effectiveness and the robustness of the control method in dealing with the slow speed, modelling non-linearity and environmental disturbances. In Section V. some main conclusions are summarized and future works are listed. Section VI is Acknowledgments and Section VI1 (Appendices) contains information relating to the VLCC used in this study.\nII. THE SLIDING MODE CONCEPT\nSliding mode control is started from the idea that a 1\" order system is easier to deal with than an n *order system. For a system of order n, design of the control system can be simplified by choosing a priori well-behaved system of order n-I, and that idea gives the way to deal only with 1\" order system. A robust control system usually consists of two parts: one is for compensating the nominal plant and the other is for the model uncertainty. To illustrate the sliding mode concept let consider a SlSO system:\nwhere X = [ x X i _._ x '\" - \" ] ' , f ( X ) and b( X ) are not exactly known, but the extent of the imprecision is upper bounded by a known continuous of X ,\nx i\" ' = f ( X ) + b ( X ) L i (1)\nAMC 2004 - Kawasaki, Japan", + "If define a tracking error ?as:\n. .. I=X-X, =[? ? 2 ... ?\"-\"I' (2)\nwith X, is the desired state, slidingsurjace is defined as:\n(3) d df s(X,r)=(-+I) '\"-\" P\nhere, I is a strictly positive constant and can be interpreted as the control bandwidth.\nThis dynamics may be visualized in the state space as a time-varying surface called sliding surjace. Then the problem of perfect tracking the state X = X, is defined as remaining or \"sliding\" along the surfaces(X,t)=O, or equivalently, as keeping the scalar quality s a t zero for all t > 0 . This does not mean that all elements of ? are zero, but the prescribed dynamics (3) are achieved. Since I is a strictly positive constant, the remaining of the system on the sliding surface s (X, t )=O implies that\n? + Oorcondition X = X, is guaranteed. For example, the sliding surface for a second-order\nsystem is given as:\ns ( i , f ) = + I P (4)\nThis corresponds to a line (with the slop -a ) moves with\nthe point ( x d , . i d ) , as illustrated in Fig. I . Fors(X,r)=O,\nthe control surface (4) describes a sliding surface with\nexponential dynamics:\ni ( r ) = exp[-I(r -fo)]P(to ) ( 5 )\nwhich ensures that the tracking error ?( t ) exponentially converges to zero in finite time along the sliding mode ( s (X, t )=O ) for any initial condition ?(to) . Thus the control objective is reduced to constructing a control law, which ensures that lim s=o . This condition can be\nguaranteed by simply choosing a control law U that satisfies the sliding condition:\nt -1-\n> yo\nIn this study, a well-known and widely used model, known as the MMG model [I21 is adopted to expresses the ship's surge, sway, and yaw motions. The model shown in formula (7) (non-dimensional form) consists of the open-water characteristics of hull(s), propeller(s) and rudder@) individually and interaction terms among them:\n(7) m'(< - v'i ' - xLr.') =Xi + Xi + Xi + X;\nI;?' + m'xi(6' + u'r') = N; + N: + N; + N ; m'(e' + u'r' + r i p ' ) = Y,' + Y; + Yn' + Y;\nwhere: - u ' , ~ ' , r ' a r e the ship's surge, sway and yaw velocities, respectively and i',~',+' are their corresponding derivatives with respect to time; - m',I; are ship mass and moment of inertia; - .r; is distance from mid-ship to the ship's centre of\n- X , Y , N terms with subscripts H , P , R , E respectively are forces in longitudinal and lateral directions and moment induced by ship hull(s), propeller(s), rudder@) and external effects, respectively. All the terms in the above mathematical model are expressed in the ship-fixed coordinate system XYZ with the origin at the centre of symmetry of the hull, and the Earth-fixed coordinate system isq%, as shown in Fig. 2.\nThe forces and moment induced by ship hull(s) are described by a multi-terns mathematical model originally presented by Kose et a/ . [ I ] . Its form is given by:\npravity;", + "X; =-m:lj'+X:.u\" + ~ : . v \" +(x: +m:,)v.r* + X;4wl< I\".r'/u'+X:,\".I\"'/U'+XI:,. 1Y. I\" . fi; = -m;.v.+ r:\"' +T:l\" I Y. I ; +x.,.' + (8) (Y ' , -m: )v ' r '+Yy +Y~v\"r'u'iU'' N, = - J L i ' + ( ~ : v ' + ~ ~ \" I v ' i v ' ) ( L ; + L ; B ) +\n+ N,Lr I Y' I Gr' +N>Gr'\nHere: u ' = m and t a n p = - ( v ' / u ' ) , m:,m;, ,JL\nare added mass and moment of inertia in the forward, transverse, and yaw directions, respectively. Non-dimensional forms are calculated using ship length (L), gavity acceleration (g), and water density ( p ) as described in [I].\nThe main particulars (hydrodynamic coefficients) of the VLCC are given in Table 1 of Appendices (Section VII). The coefficients were estimated by Free-running method for a model ship, as described in detail in [Z], [3].\n5. Environmental efects\nthe following system of equations: Forces and moments generated by the wind are given by\nx, = (l/WAC,(~A)A$'A*\nN, = ( I / ~ ) P , C , ( ~ ' , ) ~ , V , ~\nY. = ( I~W,C, (~ , )A ,V , ' (9)\nHere, p, is air density; A,,A, are transverse and longitudinal projected areas, respectively; 8, is wind relative direction; V, is relative wind speed and Cr,C>,,Ce are forces and moment coefficients in X, Y, N\ndirections, respectively. The relative direction of wind V, in the ship-fixed coordinate system is defined as in Fig. 2, the relation between wind relative direction and forces and moment coefficients C, , C, , Ca are shown in Fig. 3.\nC. Application of the Sliding Mode Control to Control Large Ships in Harbor Maneuvers\nApplying the Sliding Control method, a position and attitude tracking controller can be designed for large ships\nin harbour manoeuvres. System of following equations (10) and ( I 1) expressing ship motions at low speed range can be rewritten in the following form of non-linear equations:\nMli + N ( V , q ) = T (10) and i=J(q)v (1 1)\n, where q = [ x y yll'and v=[u v rl'are the vectors that express ship position (and Euler angle) and velocity in the horizontal plane (surge, sway, yaw), respectively. Both q and v are usually assumed to be measured. M is a matrix that expresses the influence of inertia, N is a (non-linear) matrix expressing the damping part, J is the kinematic transformation matrix that expresses the relationship between the earth-fixed reference frame xK and the body fixed reference frame X I Z (see Fig. 2),T is a matrix that expresses control forces and moment (from thrusters, tugboats, rudders and so on), as well as environment effects. The kinematic transformation matrix J is given by following formula:\ncosyl -shy 0\nJ = sinyl cosy 0 (12)\n[ o o d Suppose that the desired position and anitude in the Earth-fixed coordinate system are described by a smooth time-varying reference trajectory, which expressed by vectorq, = [ xd , y , , y, ] ', let denote the tracking error by vector = q - q,, the control law is designed to ensure that the following measure of tracking error convergences to zero:\n= $+nq (13)\nhere, A is a positive constant. The convergence of s to zero implies that the tracking error i j converges to zero.\nDefine a virtual reference trajectory q, :\ns=l ) -q ,\nor i ,=i-s=i-(ii+nii)=i,-nii (14)\nThen similar to ( l l ) , i, can be transformed to the body-fixed reference frame as i,=J(qjv, , and the body-fixed virtual reference vectors v, and v, can be computed from (1 1) as:\nV , = J - ' (q ) i , (15)\ne, = J' CtlXi i , -J(q) J-' (Q)i, I (16)\nThe control law is chosen as follows:\nT = MV, + N(v,,q)-J'(q)K,s (17) An important issue in implementation of the Decoupling Controller is how to design the manoeuvring trajectory (ship path). Letrd be the steady-state reference vector of ship position, that is q, (=-) =rd, then the ship kinematics can be used to calculate desired position and attitude [SI:" + ] + }, + { + "image_filename": "designv11_69_0001072_acc.1994.751711-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001072_acc.1994.751711-Figure1-1.png", + "caption": "Figure 1: Planar multibody interconnection", + "texts": [ + " During the first phase (the finite time transient maneuver) the multilink is transferred from an arbitrary initial state to a specified intermediate state which satisfies periodicity conditions consistent with the tracking objective. During the second phase (the exact tracking phase) internal motions are made periodic in a way to guarantee that starting at the intermediate state the orientation of the multilink tracks exactly a specified tracking objective. In Section 5 we apply the theory developed in the paper to a specific spacecraft tracking problem. We fix the origin of a non-rotating reference frame z,y) at the center of mass of the multilink. Re- I errin to Fig. 1, we specify the absolute orientation ofa planar multilink by an angle 8 ( t ) between the x-axis and the line which passes through the joint of the first link and the center of mass of the first link. We denote the joint angle between the ith and ( i + 1)st link by &. We assume that the angular momentum of the multilink about the multilink\u2019s center of mass is conserved. We denote the constant value of the angular momentum by H. Conservation of the angular momentum relates the joint angle veloc- ities &(t) , i = 1, - - e , (N - l), to the time rate of change of the orientation of the multilink 6 ( t ) as where the functions mee(4) > 0, mei(4) are 27rperiodic in each of the 4i\u2019s" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001788_j.compstruct.2006.02.029-Figure6-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001788_j.compstruct.2006.02.029-Figure6-1.png", + "caption": "Fig. 6. Energy density distributions along critical \u2018\u2018boundaries\u2019\u2019 and resulting symmetry reactions on part of the orthotropic disc. Also displacements (1:1) of contact boundary is shown together with the actual contact force distribution. Four cases: (e) radial elliptic, (f) radial + 10% tangential elliptic, (g) radial super elliptic with power = 1.7, and (h) radial + 10% tangential super elliptic with power = 1.7.", + "texts": [ + " The analytical distribution according to simple Hertz theory is elliptic and this case is shown in (b) that closely agrees with case (a), but as seen the maximum energy density are higher for the calculated distribution (about 11%). We use super ellipse distributions ((x/a)g + (y/b)g = 1) with powers g different from 2 to study further the influence from changed distributions. Cases (c) and (d) shows super-elliptic distributions, with a more rectangular distribution (super ellipse power g = 3) in (c) and a more triangular distribution (g = 1.7) in (d). Best agreement with the calculated distribution (a) is case (d), that also gives almost identical maximum elastic energy density. Fig. 6 shows the influence from changed directions for the distributions along the contact areas. For all four cases the total force in the y-direction is the same as for the cases in Fig. 5. First of all we notice a change in the reactions at the symmetry boundary with pure tension stress for the pure radial directions, cases (e) and (g). With added 10% tangential forces (rough model for friction) we get the results in case (f) where the radial distribution is elliptic and in case (h) where the radial distribution is super elliptic (g = 1.7) (as in case (d) in Fig. 5). To be noticed in Fig. 6(e) and (g) for pure radial pressure is that the maximum energy density appears at the boundary of the symmetry line, i.e., not in the interior. The displacement of the contact boundary is almost the same for all eight cases in Figs. 5 and 6, indicating that the assumption as to the load distribution is not essential for the contact problem. Also the modeling of the pin we discuss, although this is more simple and is based on isotropic material, here taken as aluminum with modulus of elasticity E = 7 \u00b7 109 Pa and Poisson\u2019s ratio m = 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001818_gt2005-68973-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001818_gt2005-68973-Figure1-1.png", + "caption": "Figure 1. Schematic of a brush seal", + "texts": [ + " NOMENCLATURE Subscripts ax Axial circ,c Circumferential n Normal to the bristle surface rotor,r Rotor Surface BR Backing Ring D Diameter of Bristle d Bristle Deflection d * Normalized Bristle Deflection ( D d= ) H Backing Ring Overhang Height K Number of Bristle Elements L Actual Bristle Length R Total Reaction Force UDF User Defined Function x,y,z Circumferential, Axial, Radial direction respectively \u00b5 Friction Coefficient \u03b4 Initial gap clearance between bristle and backing ring \u0424 Lay Angle (=0 degrees for radially aligned bristle) \u2206P Pressure Difference Across Seal 1 Copyright \u00a9 2005 by ASME url=/data/conferences/gt2005/72542/ on 07/07/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use Do Brush seals are a relatively new type of aerodynamic seal for turbomachinery applications. They have found increasing industrial use in gas turbine aero and power generation engines over the last decade. A brush seal is a set of fine diameter fibers densely packed between retaining and backing plates. There are three main components: bristle pack, front plate and backing plate (Fig. 1). It is convenient to divide the bristle pack into two regions: the overhanging region, as indicated, and the outer portion. The backing plate is positioned downstream of the bristles to provide mechanical support for the bristles under the differential pressure loads. The inherent flexibility enables the seal to survive large rotor excursions, resulting from vibration, thermal and centrifugal growth, and rotor imbalance or eccentricity during normal operation, without sustaining appreciable permanent damage" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000781_bf02644033-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000781_bf02644033-Figure1-1.png", + "caption": "Fig. 1 8 - A v e r a g e M s vs r c u r v e s fo r {a) model DS s t r u c t u r e and (b) DS s t r u c t u r e with s c a t t e r i n g in ga-owth d i r ec t i ons of the c o l u m n a r g r a i n s a s shown by the in se r t .", + "texts": [ + " l(a) and l(b) show, respect ive ly , typical c a s t - t o - s i ze DS and monocrysta l gas - tu rb ine blades that have been tes ted in advanced a i r c ra f t jet engines. Both the axis of the monocrysta l blade and the growth d i r e c - t ions of the columnar grains a re approximate ly p a r a l - lel to the C001) direct ion. Each gra in in the DS s t ruc - ture has a rota t ional degree of f reedom about i ts growth axis . As in the case of a monocrystal , the DS s t ruc ture is C. C. LAW is Assistant Materials Project Engineer, Pratt & Whitney Fig. 1-Modern air-cooled gas turbine blades directionally Aircraft, Middletown, CT 06457. A. F. GIAMEI is Senior Materials solidified to obtain (a) columnar-grained, and (b) monoerysProject Engineer, Pratt & Whitney Aircraft, Manchester, CT 06040. tal structures. Blades are macroetched and approximately Manuscript submitted November 12, 1974. 10 cm in length. expected to posses s mechanical anisotropy. The e l a s - tic anisotropy of a model DS s t ruc ture in which the growth d i rec t ion of al l the columnar gra ins is pa ra l l e l to (001> and the gra ins a re randomly rota ted about this d i rect ion has been analyzed by Wells " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002258_1.3005982-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002258_1.3005982-Figure1-1.png", + "caption": "FIG. 1. Schematic representation of the surface, of the nematic n , and tangent to the surface t directors.", + "texts": [ + " In our approach, the influence of the variations in the azimuthal angle is neglected, while the correct boundary conditions are derived using a variational analysis. A. Variational analysis Let us consider a nematic liquid crystal limited by a grooved profile, z0 x = A 1 + cos qx . 1 The direction of the grooves is assumed to be parallel to the y-axis of a Cartesian reference frame and we assume the system as two dimensional, when the x ,z -plane contains the nematic director n x ,z =ux cos x ,z +uz sin x ,z , where is the tilt angle see Fig. 1 and ux and uz are the unit vectors along the x- and z-axes, respectively. We assume the cell to be infinite in the x- and z-directions. The nematic deformation is periodic with spatial period . We limit our investigation to 0 x . Deformations are expected to decay rapidly when moving away from the surface. As we will show 0 z 2A+2 is sufficient to approximate correctly an infinite medium. Let us consider the region R, limited by thea Electronic mail: marco.scalerandi@infm.polito.it. 0021-8979/2008/104 9 /094903/9/$23" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003881_j.precisioneng.2010.10.003-Figure6-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003881_j.precisioneng.2010.10.003-Figure6-1.png", + "caption": "Fig. 6. Spindle with torque measurement equipment.", + "texts": [ + " We devised a mechanism in which the beam is fixed between the carrier and the stopper, and a piezofilm is placed on the beam, as shown in Fig. 5. In a previous study [6], a strain gage was used to detect the strain; however, it was changed to a piezofilm to improve sensitivity. The magnitude of the strain changes according to the cutting torque applied on the sun roller during end milling. The piezofilm used here is 16 mm \u00d7 41 mm and 0.028 mm in thickness. The developed speed-increasing spindle with measurement equipment was set in a vertical MC as shown in Fig. 6. The output time response at the piezofilm during idling is shown in Fig. 7(a), where the input rotary speed of the tool spindle of the MC is 2000 min\u22121 (the rotary speed of the tool is 10,000 min\u22121). The spectrum shown in Fig. 7(b) has the characteristics of the frequency response. The vertical axis indicates the half amplitude of the time response at each frequency. A strong peak appears at 33.3 Hz. This is considered to be due to the torque irregularity occurring at the spindle motor because the frequency agrees with the input rotary frequency fm" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001405_j.topol.2006.07.004-Figure11-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001405_j.topol.2006.07.004-Figure11-1.png", + "caption": "Fig. 11. \u03b51 = +, \u03b52 = \u2212. The case when n = 2.", + "texts": [ + " It occurs if and only if the body of the spider is located on an interior of an edge of D (Fig. 8). (3) Exactly two arms are stretched-out. It occurs if and only if the body of the spider is located at a vertex of D (Fig. 9). (4) The body of the spider is located at Bk . The kth arm, which is folded, can rotate around Bk (Fig. 10). It can occur only when 0 < R < Rn. Definition 3.7. Let \u03b8k (\u2212\u03c0 < \u03b8k \u03c0) be the angle from \u2212\u2212\u2212\u2192 BkJk to \u2212\u2212\u2212\u2192 JkC. The index of the kth arm, \u03b5k \u2208 {+,\u2212,0,\u221e}, is given by the signature of tan \u03b8k 2 , where \u2212\u221e is identified with \u221e (Fig. 11). We say that the kth arm is positively bended (or negatively bended) if its index \u03b5k is + (or respectively, \u2212), bended if it is either positively or negatively bended. We note that it is stretched-out if \u03b5k = 0, and folded if \u03b5k = \u221e. Definition 3.8. (1) We call D \u2229 {C = (x, y): |CBk| = 2} the kth edge of D. (2) Define D\u030a by D\u030a = { C = (x, y): 0 < |CBk| < 2 (1 k n) } , namely, D\u030a = { IntD \\ {B1, . . . ,Bn} when 0 < R < Rn, IntD when Rn < R < 2, and call it the open body domain. It is the domain where the body of a spider can be located whose arms are all bended" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000656_pesc.2003.1218325-FigureI-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000656_pesc.2003.1218325-FigureI-1.png", + "caption": "Fig. I(a)CmsssectionofCphase SRh4.", + "texts": [ + " PID controller [XI has heen used in speed reduced speed overshoot, low starting time and reduced speed loop for predicting the drive performance. oscillation. Such a performance is not easy to meet, as it requires to accommodate a large number of system nonlinearities. The electromagnetic torque developed by the SRM is a nonlinear function of stator current and rotor position. For This section describes the control requirements of a typical 816 forward motoring, the appropriate stator phase winding must pole configuration SRM as shown in Fig. I(a). It is a fourremain excited only during the period when rate of change of phase machine excited by a semiconductor-controlled midphase inductance is positive. Else, the motor would develop point inverter. Ideal inductance profile of the four phase braking or no torque at all. The inductance of a stator phase is windings of motor versus rotor position is s h o w in Fig. l(b) maximum when its pole is directly opposite the rotor pole and where the different zones of inductance, namely minimum is minimum when the inter-polar rotor region is opposite it", + " The the excited phase has attained its maximum value and remains detailed specification of the motor used in this investigation is at this value for some time depending on the overlap of stator given in Appendix. and rotor pole widths. Later the inductance begins to fall as the rotor pole moves away from the stator pole. The concept of tun-on and turn-off angles for any phase winding of switched reluctance motor is shown in Fig. l(c). The practical range of tun-on angle and tum-off angle depends on the inductance profile and therefore on the configuration and pole geometry of the particular switched reluctance motor. respectively 01 e2 e3 04 Fig. I@) Ideal inductance profile of SRM. SimofTID A . Control Philosophy Gem\" cc&dla 0. + The schematic of the closed loop drive system of a typical 4 phase SRM is shown in Fig. 2. It consists of outer speed loop comprising motor with rotor position sensor, speed controller COUQOlia ~comnuuuon Lo@@ . and inverter. The inner current loop consists of current sensors, reference current generator, current controller and commutation logic. The working of the system is briefly discussed here so as to develop the control algorithm" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002034_2007-01-2860-Figure3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002034_2007-01-2860-Figure3-1.png", + "caption": "Figure 3: Degrees of Freedom and Reference Axes ISO 4130 & DIN 70000", + "texts": [ + " A perceptible disadvantage of this solution is the impossibility of rear transmission. This type of suspension is used only in frontwheel drive vehicles due to its own natural geometry. The greater functional limitation is the fact that the wheels follow the body roll in curves, generating positive camber angles in outer wheel, which can cause oversteering. During the development of this study, ISO 4130 and DIN 70000 norms were used. These norms establish vehicle reference axes, as well as the degrees-of- freedom, which are listed as follows, and presented in Figure 3: - yaw angle: Z-axis rotation angle; - pitch angle: Y-axis rotation angle; - roll angle: X-axis rotation angle; For kinematic validation of the suspension, two basic movements are studied: first the symmetrical vertical movement of the wheels, caused by pitch for instance, as illustrated in Figure 4. The kinematic behavior influenced by the body roll, as in a curve maneuver, is also studied, generating an asymmetrical vertical movement of the wheels, as illustrated in Figure 5. 3 The total dimension of both symmetrical and asymmetrical movements depends on project parameters such as unsprung mass, in-roll mass transference, spring stiffness and end-restraints characteristics" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002353_tac.1966.1098435-Figure4-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002353_tac.1966.1098435-Figure4-1.png", + "caption": "Fig. 4. Spquist locus of ItOo) ai Example 2 .", + "texts": [ + " 748 IEEE TRANSACTIONS ON AUTOMATIC C TROLC BER Example 2 Let the linear subs>-stem have the transfer function Plants of this type are encountered in missile roll control systems, for example. TV(s) satisfies the conditions for stability in the limit. Let r ( t ) = a + bt (12) where a and b are real numbers. Then all conditions of Theorem 2 are satisfied and from the Xyquist plot of IT-(ju) (see Fig. -I), it can be concluded that the output c ( t ) of the system will follow the input r ( t ) with a zero steady-state error for all 4(u, t ) in the sector [0.595, 5.01. I t should be noted that the circle in Fig. 4 is not unique (the same applies to the degenerate circle of Fig. 3), and other permissible sectors for @(u, t ) can be found. The circles would become smaller in diameter if the center were selected closer to the jw-axis, until the limiting case is reached at -0.044, with a circle of zero diameter corresponding to a gain of 22.75 which can have no time variation (the system is on the verge of instability). On the other hand, if the center of the circles were selected progressively towards the left, the diameter of the circles would become larger" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000347_isatp.1997.615388-Figure4-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000347_isatp.1997.615388-Figure4-1.png", + "caption": "Figure 4: Graspless manipulation for insertion task", + "texts": [ + " Motion of the robot doesn\u2019t bring the same motion of the object. Then even if we control a robot hand in perfect; we may not be able to control the motion of an object. But the other way, the posture of the object can be changed larger than that of the robot hand. So if you can control the position and posture of the object, it is a big advantage. Using these advantages, yoii can achieve various dexterous tasks which cannot be done by pick-andplace. An Insertion task is shown as an example in Figure 4. If a robot grasps a part rigidly, it cannot make the part insert into a hole because a side of a wall becomes an obstacle against a finger. To assemble the part, the robot needs to remove one finger from the object and makes the part slide down into the hole using an edge of the hole as drawn in the lower figures in Figure 4. The orientation of the part may change even if that of the robot is fixed. This can be achieved easily by human operators but not by robots. 3 Reactive Motion for Manipulation Since graspless manipulation is able to be regarded as an operation which changes a state of contact between an object and floors, we can make an off-line planning method of graspless manipulation that uses a graph theory, which generates a contact state transition graph with nodes as state of contact and arcs as manipulation method to translate a state to another" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000215_icsmc.1995.538489-Figure3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000215_icsmc.1995.538489-Figure3-1.png", + "caption": "Figure 3: Configuration of the boundary singularity for GRYPHON manipulator.", + "texts": [ + " Therefore, define where 'Ro is the rotation matrix which transforms the components expressed in coordinate 0 into those expressed in coordinate 3. Then, 0 .*S. - &Cr -&er -4C. 0 E.C. + E , - Irs. EB - kS4 -&S, 0 1 1 0 0 0 0 1 at +a&% + asCa - d & . m _ _ _ _ _ _ _ _ _ _ _ _ - _ _ _ _ _ _ _ - Obviously, when the interior row of 'JE will become a null row. Thus, the interior singular direction is in parallel with the ZQ axis. As for the boundary singularity, it occurs when 83 = 0 or R. However, as far ae the working space of GRYPHON manipulator is concerned, 83 = R is unreachable. Referring to Fig. 3, it can be seen that the boundary singularity happens when the Wrist is located on the x p axis or the x g axis is in parallel with the x p axis. This key point can also be identified by viewing the Jacobian matrix in coordinate 3. When the boundary singularity occurs, 8 3 = 0, and 3 M ~ becomes -d& -&Cr -d&r 0 et +a* - d& a. - drS. -d& 0 _ _ _ _ _ _ _ _ _ _ _ _ - 1 1 1 0 0 0 0 1 , it is clear that the first row and the fourt dependent. Therefore, the boundary sin- gular direction is in the x3 direction. This viewpoint can also be observed from Fig. 3 where the End-effector linear velocity component in x 3 direction is linearly dependent with 4234 = 8 2 + 83 + 84. After analyzing all of the singularities of GRYPHON manipulator, considerations and various resolution methods for the Inverse Kinematics singularity problem will be presented in the next section. 3 Singularity Considerations and Resolutions Then, the proposed Singularity Isolation plus Compact QP (SICQP) method will be presented. 3.1 The Pseudoinverse Method The PI solution is obtained by the following quadratic programming: (14) 1 " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001205_pime_proc_1972_186_090_02-Figure18-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001205_pime_proc_1972_186_090_02-Figure18-1.png", + "caption": "Fig. 18. Gimbals mechanism and Euler\u2019s angles", + "texts": [], + "surrounding_texts": [ + "Attention is drawn to a new class of rotary piston machines for the exchange of fluid energy with a rotating shaft. The machines are characterized by the near free body motion (precession in some instances) of their working elements and a compact arrangement with several chambers on each rotor. They offer a unique combination of freedom from inertial loads, the ability to seal by continuous sliding contact and freedom from valve mechanisms with a stationary casing containing-inlet and outlet ports. One of these machines has been built and tested as a two-cycle engine and extensive effort has been necessary to improve the sealing effectiveness. However, recent results have pointed the way to a solution of the seal effectiveness problem for certain compressor duties and a high speed engine. A much more extensive programme of work is necessary to improve seal life and to satisfy the effectiveness requirements of low speed engines but it seems probable that improvements can be made if the development effort is forthcoming. There are similarities with the Wankel engine but there are also significant differences many of which favour the new machines." + ] + }, + { + "image_filename": "designv11_69_0003275_detc2009-86970-Figure3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003275_detc2009-86970-Figure3-1.png", + "caption": "Figure 3. Geometry of master gear tooth space.", + "texts": [ + " Then, the process of double flank gear rolling testing is simulated by applying the concept of tooth contact analysis (TCA) [10-11]. Tooth contact types including surface-to-surface contact and tip-to-surface interference are considered, and three possible combinations of these two contact types occurring on each tooth flanks are discussed as well. The results of this study can provide the industry a significant process to establish the analysis and capacities for double flank testing. 2 MATHEMATICAL MODEL OF GEARS 2.1 Mathematical Model of Master Gear M\u03a3 Figure 3 illustrates a single tooth space of the master gear M\u03a3 which comprises the left and right side tooth flanks ML and the MR in coordinate system ),( MiMiMi YXS . Herein, bMr , pMr and tMr are the radii of the base circle, pitch circle and addendum circle, respectively. LM _\u03be and RM _\u03be represent the involute profile parameters (i.e. the involute polar angles), which determines the position of point on involute curves ML and MR. Otherwise, M\u03b1 represents the pressure angle, while MN denotes the tooth numbers" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003591_016918610x538471-Figure5-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003591_016918610x538471-Figure5-1.png", + "caption": "Figure 5. Geometry of the simulated hand.", + "texts": [ + " If calculated paths are colliding (or not valid), the configuration halfway along the GSn portion is generated (noted qinter) and the algorithm is recursively applied to two subpaths connecting this intermediate configuration to the initial and the final ones (transform_path function).When all the necessary subdivisions are completed, the concatenation of all elementary subpaths is collision-free and respects the manipulation constraints. The process is guaranteed to converge thanks to the reduction property. This section presents results obtained from computer simulations, for four different planning problems. We have developed a planner written in C++ that uses the PQP library [25] for collision detection. The simulated hand has four 3-d.o.f. fingers (Fig. 5). As the fingers have only 3 d.o.f., the inverse kinematics of the fingers have only one solution. If it is not the case, a solution to the inverse kinematics problem is randomly chosen during the sampling phase. The contact model needed for force closure test is the PCWF (point contact with friction) model and the chosen friction coefficient is 0.8. The graph is built using the visibility-PRM method [26]. The computation times correspond to experiments conducted on a PC equipped with an Intel Core2Duo processor (two 2" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001118_0471758159.ch5-Figure5.16-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001118_0471758159.ch5-Figure5.16-1.png", + "caption": "FIGURE 5.16 Modeling of a physical capacitor with an equivalent circuit.", + "texts": [ + " Thus ceramic capacitors are typically used for suppression in the radiated emission frequency range, whereas electrolytic capacitors, by virtue of their much larger values, are typically used for suppression in the conducted emission band and also for providing bulk charge storage on printed circuit boards as we will see. For a more complete discussion of capacitor types see [6]. Both types of capacitors have similar equivalent circuits, but the model element values differ substantially. This accounts for their different behavior over different frequency bands. Both types of capacitor can be viewed as a pair of parallel plates separated by a dielectric, as illustrated in Fig. 5.16. The loss (polarization and ohmic) in the dielectric is represented as a parallel resistance Rdiel [1]. Usually this is a large value, as one would expect (hope). The resistance of the plates is represented by Rplate. For small ceramic capacitors, this is usually small enough in relation to the other elements to be neglected. Once again, the leads attached to the capacitor have a certain inductance represented by Llead and capacitance Clead. Again, these parasitic element values depend on the configuration of the two leads" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000314_lfa.1988.24952-Figure4.2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000314_lfa.1988.24952-Figure4.2-1.png", + "caption": "Figure 4.2. Explanation of Definition 4.1", + "texts": [ + "1: Let the Path, be P i , N , P f , and the P a t h , be Pi,N,Pf,, where N, and N, are the corners of p a t h , and p a t h , respectively. T o simplify the notations, we define 8 ,=LN,Pi ,Pf l , Q,=LN,Pf,Pi,, Q,=LN,Pi,Pf,, and 8,==LN,Pf,Pi2. If N, is in D,, 0, and 8, are greater than zero. When N, is on the ray P i , P f , , 0, equals to zero. When N, is on the ray P f ,P i , , 8, equals to zero. Otherwise, 8, and 8, are less than zero. If N, is in S,, O3 and 8, are greater than zero. If N, is on the ray Pi,Pf,, 0, equals t o zero. If N, is on the ray .Pf,Pi,, 8, equals t o zero. Otherwise, 19, and 8, are less than zero (see Figure 4.2). Definition 4.2: The circle with center a t P i , and radius v l t + r l + r 2 may intersect RC( t ,P i , ) a t points I l ( t ) and Iz(t). The line Ii(t)Pi, (i=1,2) intersects RC(t, P i , ) a t points idck(B1) where k=1,2. When t changes, points idck(B1) define a pair of curves named initial discriminant curves, IDCk(B,) (k=1,2) (see Figure 4.3). The curves IDC,(B,) and IDC,(B,) are symmetric t o the line Pi,Pi,. U Figure 4.3: Initial Discriminant Curves The line P i ,P i , separate the whole plane into two subplanes, then one branch of IDC(B,) which locates in the same subplane with the point Pf, is denoted as IDC1(Bl)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001293_rob.20099-Figure6-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001293_rob.20099-Figure6-1.png", + "caption": "Figure 6. Step 3 of guidance example: a step 3.1 and b step 3.2, respectively.", + "texts": [ + "1 Measure the new offset, d2, along the same detector considered in step 1. Calculate the actual orientation angle of the platform from the x axis, a, and determine the difference = d\u2212 a, where d is the desired platform orientation. 2.2 Translate, by b, and, then, rotate, by , the platform about the center of the fixed side of the platform to compensate for d2 and . This movement should result in the LOS considered in step 1 to keep hitting the center of its detector and, furthermore, the platform achieving its final desired orientation. Step 3 Figure 6 : 3.1 Determine the largest offset among the two other PSDs, d3, and determine the corresponding distance, l, that the platform needs to move along the slope of LOS #1. 3.2 Translate the platform by l along the slope of LOS #1. In theory, once the two LOS, #1 and #2, hit their respective detectors\u2019 centers, the third laser beam LOS #3 should automatically hit the center of its detector as well. Since systematic motion errors can never be completely eliminated, corrective vehicle motions would only decrease these errors" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003622_s0026-0657(09)70201-2-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003622_s0026-0657(09)70201-2-Figure1-1.png", + "caption": "Figure 1. Conventional heating/cooling channels (left), Conformal heating/cooling channels with DMLS (right).", + "texts": [ + " \u201cIt is suitable for making very complicated parts to a very high quality in a much faster time than traditional moulding,\u201d agrees Mayer. \u201cFor example, a gear wheel requires special milling tools to be created, and can take up to four months using existing technology. Using DMLS, it is possible to make a prototype in just one week.\u201d Benefits of conformal cooling According to an EOS white paper, conformal cooling works by creating a suitable cooling channel at a well defined distance to the cavity, which is impossible using a conventional drilled cooling mechanism (Figure 1). Cooling channel cross sections can take almost any shape. Turbulence of the coolant (the desired high Reynolds number) within the system can thus be controlled by actively choosing different cross sections and by switching between different cross sections. As a consequence, turbulence inside the coolant stream is generated close to the mould cavity along the whole path of the channels. Changing the cross sections or forking the cooling channel can be done without splitting up the form. This allows for additional heat/cooling advantages in areas that cannot be reached by conventional methods" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001979_50037-5-Figure7.20-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001979_50037-5-Figure7.20-1.png", + "caption": "Figure 7.20 Model after meshing process.", + "texts": [ + "18) click button [A] Set in the Size Controls \u2192 Lines option and pick the curved line on the front of the arm. Click OK afterward. The frame shown in Figure 7.17 appears. In the box No. of element divisions type 4 and press [C] OK button. In the frame MeshTool (see Figure 7.18) pull down [B] Volumes in the option Mesh. Check [C] Hex/Wedge and [D] Sweep options. This is shown in Figure 7.18. Pressing [E] Sweep button brings another frame asking to pick the pin and the arm volumes (see Figure 7.19). Pressing [A] OK button initiates meshing process. The model after meshing looks like the image in Figure 7.20. Pressing [F] Close button on MeshTool frame (see Figure 7.18) ends mesh generation stage. After meshing is completed, it is usually necessary to smooth element edges in order to improve graphic display. It can be accomplished using PlotCtrls facility in the Utility Menu. From Utility Menu select PlotCtrls \u2192 Style \u2192 Size and Shape. The frame shown in Figure 7.21 appears. From option [A] Facets/element edge select 2 facets/edge, which is shown in Figure 7.21. In solving the problem of contact between two elements, it is necessary to create contact pair" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002979_978-3-540-70534-5_7-Figure26.16-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002979_978-3-540-70534-5_7-Figure26.16-1.png", + "caption": "Figure 26.16: Camera-based automatic parking [Br\u00e4unl, Franke 2001]", + "texts": [ + " V v1 \u2026 vheight, ,( ) SobelH xi y1,( ) \u2026 SobelH xi yheight,( ) i 1= width , , i 1= width = = W w1 \u2026 wwidth, ,( ) SobelV x1 yj,( ) \u2026 SobelV xwidth yj,( ) j 1= height , , j 1= height = = \u0398v 1 2 -- max vi 1 i width\u2264 \u2264{ }\u22c5= \u0398w 1 2 -- max wj 1 j height\u2264 \u2264{ }\u22c5= Automatic Parking 433 26.7 Automatic Parking Parking aids have been introduced in luxury vehicles a number of years ago and have since made their way into smaller cars and after-market systems. While simpler systems only measure the distance between the front and rear bumper of a vehicle to an obstacle, more complex systems can automatically park a car on the press of a button (see Figure 26.16). Sonar sensors have been the choice for most commercial parking aid systems, while radar and laser sensors have been used mainly in research applications. In the patent application [Br\u00e4unl, Franke 2001] a camera-based approach is described that could serve as an alternative low-cost sensor, providing significantly more accurate and detailed information than sonar sensors. Motion Stereo The principle used in this patent application is to apply stereo processing to subsequent images of a moving monocular camera (motion stereo)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002400_s12239-008-0037-2-Figure22-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002400_s12239-008-0037-2-Figure22-1.png", + "caption": "Figure 22. Comparison of steering characteristics of a ordinary car and a Formula car.", + "texts": [ + " The critical lateral acceleration of the rear wheel for each weight distribution (for a vehicle weight of 1500 kg) was calculated using equations (5)~(9). The critical lateral acceleration was greatest for the weight distribution ratios 50:50 and 55:45 at 0.86 G. It was lower for a weight distribution ratio of 60:40, and even less for 40:60. The friction circle of the rear wheel is a maximum for the weight distribution ratio of 40:60; the total cornering force was reduced because the load movement of the inside and outside wheels increases as the vertical wheel load increases. Figure 22 shows the difference in the weight distributions of an ordinary car and a formula car. Because a ordinary car is heavy, the rear wheel previously entered a state of saturation while in the nonlinear area of the tire when the weight distribution of the rear wheel was increased, and this increased the propensity of the car to spin. The influence of the front and rear weight distribution on circuit maximum-speed cornering in a Formula car was examined. As a result, the following conclusions were made: (1) Through maximum speed cornering simulations, it was found that a vehicle weight distribution ratio of 40:60 was best for maximum speed cornering" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001410_0005-2795(72)90018-9-Figure3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001410_0005-2795(72)90018-9-Figure3-1.png", + "caption": "Fig . 3- P l o t s o f t h e e x t i n c t i o n a n g l e s (Z) a g a i n s t t h e s h e a r r a t e s (G). IS], r - - 5 ; ~ , r = IO a n d O , r 2o a n d 30. C o n c e n t r a t i o n : 4 m g / m l ; i n c u b a t i o n : 25 \u00b0C f o r 24 h . A r a b i c n u m e r a l s in t h e f i g u r e m e a n a p p a r e n t l e n g t h s o f f l a g e l l a in /~m.", + "texts": [ + " The re la t ive values of the difference of An, (Anr_ 2 0 - - A n r - 3o)/Anr = 20, became small as the shear ra te increased. This was undoub ted ly due to the po lyd ispers i ty of the po lymers Biochim. Biophys. Acta, 278 (1972) 585-588 in solution. The dashed lines in Fig. 2 ind ica ted the expec ted behav iour of the An values at G = 9 \u00b0 s -1 and 0.3 s -1, p rov ided t ha t the po lymers are monodisperse and their length is cons tan t above r = 20. The po lyd i spers i ty of po lymers in solut ion also affected the ex t inc t ion angle, g (Fig. 3). The appa ren t mean length of flagella a t r ~ 20 and 30 e s t ima ted from the observed ex t inc t ion angles var ied from 3 # m (at G = 9 \u00b0 s -t) to I I # m (at G = 0. 3 s- l ) . According to Donne t \u00b0, i t holds t h a t L o a = Iq\u00b0(x)x~d'v/ I ~(x)xadx (3) where L 0 is the appa ren t mean length of po lymers at G = o and go(x) is the volume fract ion of po lymers having ' )ength x. The present resul t suggests t ha t go(x) was very a b road funct ion of x. Fo r example , if a Poisson d i s t r ibu t ion is assumed, i" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001572_1.6868-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001572_1.6868-Figure2-1.png", + "caption": "Fig. 2 Sensing truss.", + "texts": [ + " 3 Block diagram for direct and nearest-neighbor vehicle coordination: V(i) , vehicle dynamics; T(i) , tracking controller; R(i) , regulation controller; P(t), temporal functional; rT(P), spatial functional; \u2014\u2014, hardwired signals; - - - -, wireless communication; \u2014\u00b7\u2014, wireless mea- surement; , receiver; and , transmitter. measurements of aircraft are received from GPS satellite transmitters that are well separated. If relative measurements are needed, the transmitters on the aircraft need to be well separated. In a planar system, two transmitters and one receiver on each of the i th and j th aircraft fully determine the relative position and orientation of the j th aircraft relative to the i th aircraft and the i th aircraft relative to the j th aircraft (Fig. 2). In a three-dimensional system, three transmitters and one receiver on each of the i th and j th aircraft fully determine the relative position and orientation of the j th aircraft relative to the i th aircraft and the i th aircraft relative to the j th aircraft (communications with V. Perez, System Programmer, Mechanical and Aerospace Engineering, North Carolina State University, Raleigh, North Carolina). The receiving signals and the transmitting signals for the measurement process make up a sensing truss, which possesses stability characteristics that are analogous to the stability characteristics of the corresponding structural truss (Fig. 2). The complexity of the network is important, particularly in the event that failures occur. The methods and procedures for accommodating for failures grow in complexity with the complexity of the network. The question arises to what extent the complexity of the network can be minimized. The two simplest networks are direct networks and nearest-neighbor networks. Direct networks use inertial measurements on each aircraft, eliminating the need for relative measurements, all together. Nearest-neighbor networks designate at least one aircraft as the leader and the others as the followers" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000215_icsmc.1995.538489-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000215_icsmc.1995.538489-Figure2-1.png", + "caption": "Figure 2: Configuration of the interior singularity for GRYPHON manipulator.", + "texts": [ + " According to the above definition, 'ME (which ie the reduced tool-configuration Jacobian matrix expressed in coordinate 0 with the End-effector as the velocity reference point) can be expressed as M11sx8 t M 1 2 s ~ i _ _ _ + _ _ _ M21ix3 I M22ix2 Therefore, the conditions for GRYPHON's singularities can be obtained by equating the determinant of 'ME to zero as follows det('ME) = a2a3~3(a l + a 2 ~ 2 +a3C23 - d5S234) = o (7) where Ci, Si , Cij, Sij , Ci'k, and s i j k represent coe(8j), s i n Bi), cos(8; + Bj), sin(di + 8j), cos(8i + 8 j j- e!), and sin[& + 8 j + 8k), respectively. From (7), it is clear that, there are two conditions, denoted as 7i and 7 b for GRYPHON singularities. One is the interior singularity with and the other is the boundary singularity with (8 ) (9) A vi = a1 + a2C2 + a3C23 - d5S234 = 0 A 7 s = ss = 0. Referring to Fig. 2(a), the interior singularity happens when the End-effector ie located on the g o axis. A special case is also shown in Fig. 2(b). In this case, both o1+a2C2+a3C~s and d 5 S 2 3 4 are equal to zero. When the interior singularity occurs, one of the axes in XI, z 2 , or z3 may be chosen aa the singular direction. Here, the direction in $3 is chosen. Therefore, define where 'Ro is the rotation matrix which transforms the components expressed in coordinate 0 into those expressed in coordinate 3. Then, 0 .*S. - &Cr -&er -4C. 0 E.C. + E , - Irs. EB - kS4 -&S, 0 1 1 0 0 0 0 1 at +a&% + asCa - d & . m _ _ _ _ _ _ _ _ _ _ _ _ - _ _ _ _ _ _ _ - Obviously, when the interior row of 'JE will become a null row" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000662_s00366-004-0269-3-Figure10-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000662_s00366-004-0269-3-Figure10-1.png", + "caption": "Fig. 10 A compound truss", + "texts": [ + " According to the linkage-truss duality [8], the truss corresponding to a linkage consisting of dyads only is a simple truss, namely, a truss constructed by starting with a basic triangular element, and adding two connected rods at a time [11]. On the other hand, linkages composed of higher order modular groups correspond to compound trusses in which all the analysis equations are to be solved simultaneously. Thus, in accordance with the results of the section Transferring the method to linkages, the transformed Willis method in trusses enables to replace a compound truss with a simple one using the proposed technique. To clarify this idea, the compound truss given in Fig. 10 is analysed using the transformed Willis method. The truss of Fig. 10 is compound, thus it is reasonable to attempt solving it using the proposed procedure. The linkage dual to this truss is shown in Fig. 11a. The linkage of Fig. 11a, which actually is the same as the one treated in the section \u2018\u2018An example for the analysis of linkage by means of the transformed Willis method\u2019\u2019 (Fig. 8), is composed of a driving link 1\u2019 and a tetrad, thus according to the previous section it would be efficient to fix link 3\u2019 and change the driving link to 4\u2019, as shown in Fig. 11b", + " One of such methods is the wellknown graphical method, called Maxwell-Cremona diagram [12], shown in Fig. 12b. One can see that consistently with the results reported in [2], the Maxwell Cremona diagram of the truss of Fig. 12a is identical to the image velocity diagram of its dual mechanism, as appears in Fig. 8. The solution of the transformed truss can now be substituted into Eq. 7 to yield the ratio between the weighted value of the external force and the weighted forces in the rods of the original truss. The algebraic manipulations needed to find the force in rod 6 of the original truss (Fig. 10) appear in Fig. 12c. The paper has introduced an approach for transforming methods between engineering fields through graph representations. It employs the fact that planetary gear trains and linkages have been represented by the same graph representation, to transform the Willis method from planetary gear trains to the terminology of linkages. Since the Willis method is applicable to solve compound gear trains, upon transformation, it became a method suitable for analysis of compound linkages, such as those containing tetrads" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001329_j.jmatprotec.2005.02.261-Figure3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001329_j.jmatprotec.2005.02.261-Figure3-1.png", + "caption": "Fig. 3. Description of the problem domain and its parameters.", + "texts": [ + "261 When the thermal resistance of the interstitial medium is high and the one of the asperities is negligible (which is often the case in several practice configurations) the thermal contact resistanceRc is essentially due to the thermal constriction resistances, Rcs, as: Rc = R(1) cs + R(2) cs (2) where R (1) cs and R (2) cs are the total thermal constriction resistance of solid (1) and solid (2), respectively. 3. Problem formulation Taking into account of Eq. (2) it is sufficient to calculate Rcs for each cylinder in order to determine Rc. For this purpose, we consider a semi-infinite cylinder, as shown in Fig. 3, with a radius b and thermal conductivity k. The surface z = 0 is subjected to N uniform annular heat sources qj regularly spaced (which simulate the heat flux from the other cylinder). The remainder of the surface z = 0 and the surface r = b are assumed to be insulated. The temperature at z \u2192 \u221e is zero. The steady-state heat conduction into the cylinder is twodimensional, T (r, z), and can be described by 1 r \u2202 \u2202r ( r \u2202T \u2202r ) + \u22022T \u2202z2 = 0, (0 \u2264 r \u2264 b; z \u2265 0) (3) with the boundary conditions( \u2202T \u2202r ) r=0 = 0 (4) ( \u2202T \u2202r ) r=b = 0 (5) \u2212k ( \u2202T \u2202z ) z=0 = { qj : at contact j 0 :elsewhere (6) Tz\u2192\u221e = 0 (7) 4" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001164_1-84628-559-3_20-Figure8-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001164_1-84628-559-3_20-Figure8-1.png", + "caption": "Fig. 8. Photograph of the assembly 3-DOF positioning device", + "texts": [ + " The friction adjusting mechanism set on the top surface of the stage is used for providing a suitable holding fricitonal force for the rotaional stage. Figure 6 shows a photograph of the modualrized springmounted PZT actuator, in which the PZT actuator has the dimension of 5 5 10 mm (Tokin). The stiffness of the spring is 0.023 N/mm. Six actuating units are symmetrically mounted to the positioning stage having a radius on the bottom side. The positioning stage made of stainless steel having a mass of 2.6 kg is then set on the base having the same radius on the contace surfaces. Figure 7 shows the positioning stage and the base. Figure 8 shows the 3-DOF positioning stage. The waveform of applied voltage was generated by LabVIEW, which is a Windows supported graphical programming language. A 16-bit DA/AD converter was used to transform the pulse driving waveform to the power amplifier and then to the PZT actuator. Due to the number limitation of measuring probes, two kinds of gap sensors were used to detect the simultaneous motion behaviors of the positioning stage along three rotation axes. They were the capacitive gap sensor (ADE 5300-5504) and the fiber optic displacement sensors (Fotonic MTI-2100)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002018_004051756903900209-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002018_004051756903900209-Figure1-1.png", + "caption": "Fig. 1. Transducer geometry.", + "texts": [ + " For these reasons, cord forces which are in directly determined from strain measurements on loaded tires are more likely to show qualitative trends rather than quantitative results. The main purpose of this paper is to describe a tirecord tension trans ducer which measures cord forces directly without intermediate cord strain measurements. Experimental Transducer Description \u2022 The t~record tension transducer is basically a 0.375 111. aluminum bar with a 0.100 X 0.050 in. rectangular ~ross section (Fig. 1). A 0.040 in-diameter hole IS located near each end of the bar. Through each at WESTERN OREGON UNIVERSITY on June 3, 2015trj.sagepub.comDownloaded from Tire Parameters The data reported in this study were obtained from three, identically constructed 8.25-14 two-ply rayon tires. Each ply had a cord count of 19 ends/in. be fore expansion. In the finished tire, the angle between the cord at the crown and the circu1l1feren_ tial line at tread center was 36\u00b0. The location of the transducers and the meridional geometry of the tires at 24 lb/in" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003387_c0ay00382d-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003387_c0ay00382d-Figure1-1.png", + "caption": "Fig. 1 Chemical structures of sodium risedronate (A) and its specific contaminant (B).", + "texts": [ + " The potentiometric response towards a highly hydrated risedronate species was strongly promoted by the presence of the PANI layer. In contrast to uncoated membranes, the PANI-coated membrane exhibited a sensitivity of 39.9 0.9 mV decade 1 and the detection limit reached 0.45 mg mL 1. The results of the potentiometric analysis of risedronate agreed with those obtained with a validated chromatographic method. The TDDMACl-based membrane coated with the PANI layer differentiated 0.94 wt % and 3.44 wt % of a specific contaminant in 1.6 mg mL 1 risedronate. Risedronate sodium (Fig. 1A) belongs to a class of drugs called bisphosphonates that are widely used in the treatment and prevention of diseases associated with bone loss.1 The pharmacological function of these active compounds is determined by their P\u2013C\u2013P fragments, where two phosphate groups are covalently linked to a carbon atom. Analyses of bisphosphonates are usually done using high-performance liquid chromatography (HPLC) coupled with a variety of spectroscopic methods.2 Direct detection techniques with UV, fluorescence, or electrochemical detectors require the pre- and post-column derivatization of bisphosphonates", + "6 The observed phenomenon resulted from stronger PANI hydration in the presence of highly hydrated anions, which facilitated their transport through the inter-particle space of the PANI layer.7 Here, PANI as a hydration-promoting medium is tested for the potentiometric detection of highly hydrated risedronate. This study evaluates the influence of PANI on the potentiometric determination of risedronate in both the model and in real samples. Moreover, we discuss the applicability of the proposed sensor to detect a specific contaminant (Fig. 1B) in risedronate. High-molecular-weight poly(vinyl chloride) (PVC), 2-nitrophenyl octyl ether (NPOE), tridodecylmethylammonium chloride (TDDMACl) and tetrahydrofuran (THF; stored over a molecular sieve) were purchased from Fluka (Selectophore, Switzerland). Aniline (>99%) was supplied by Sigma-Aldrich (Germany). Acids and various inorganic salts used in potentiometric measurements were of analytical grade from Lachema (Czech Republic). Ammonium dihydrogen phosphate ($99.0%), tetrabutylammonium bromide ($99.0%) and ethylenediaminetetraacetic acid disodium salt (Reag. Ph. Eur., 99\u2013101%) were supplied by This journal is \u00aa The Royal Society of Chemistry 2010 Pu bl is he d on 1 8 A ug us t 2 01 0. D ow nl oa de d on 0 4/ 01 /2 01 4 01 :1 6: 24 . Fluka (Germany). Sodium hydroxide (98%) was purchased from Lach. Ner. (Czech Republic). HPLC-grade methanol was produced by J. T. Baker (Holland). The working standard of sodium risedronate (Fig. 1A), a specific contaminant of sodium risedronate (Fig. 1B) and tablets were obtained from Zentiva k.s. (Prague, Czech Republic). The analyzed tablets were composed of: microcrystalline cellulose, crospovidone, magnesium stearate, hypromellose, talc, macrogol 6000, titanium dioxide, ferric oxide red, ferric oxide yellow. 1 mg of TDDMACl, 33 mg of NPOE and 66 mg of PVC were dissolved in 1.0 mL THF at ambient temperature. To prepare 0.5 mm thick membranes, 1.0 mL of the solution was deposited onto a metallic ring with an internal diameter of 16 mm resting on a glass plate, and dried in air", + " The intra-day potentiometric precision (expressed as relative standard deviation, RSD) was evaluated by determination of risedronate in working standard solutions at two concentrations (1.41 and 14.1 mg mL 1) three times with experimental electrode within one day. The stability of the potentiometric parameters was tested with prepared membranes over a period of one year. When not used for measurements, the membranes were stored in distilled water. A 2 mg mL 1 solution containing a specific contaminant (Fig. 1 B) was mixed with pure sodium risedronate (1.6 mg mL 1) to yield various levels of contaminated risedronate samples. A program found at http://www.molinspiration.com was used to calculate the log P values of anionic species. This journal is \u00aa The Royal Society of Chemistry 2010 HPLC experiments were run on an Agilent 1100 Series (Santa Clara, USA) with an Inertsil ODS 2 column (250 mm 3.2 mm 5 mm). The mobile phase consisted of 0.005 M ammonium dihydrogen phosphate, 0.002 M tetrabutylammonium bromide and 0", + " The intra-day precision was tested with repeated measurements of working standard solutions at two concentrations (Table 1). Recoveries of 103.0\u2013100.4 and a relative standard deviation ranging between 3.0 and 0.8% demonstrated that the electrodes are suitable for risedronate analysis. The potential variability did not exceed a few millivolts. Potentiometric detection of a contaminant in risedronate The ability of a TDDMACl-based membrane coated with a PANI layer to discriminate a contaminant in risedronate was tested. The specific contaminant of risedronate used was a dimer of risedronic acid (Fig. 1B). With the proposed potentiometric sensor, the discrimination should be based upon differences in their hydrophilicity. According to the calculated log P values, the active pharmaceutical ingredient and its contaminant have highly comparable and simultaneously poor lipophilicity, i.e. risedronate (log P \u00bc 4.829) and contaminant (log P \u00bc 5.151). As a point of comparison, phosphate, known to be a very hydrophilic anion has log P \u00bc 4.53, whereas chloride, known to be a lipophilic anion has log P \u00bc 2" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002979_978-3-540-70534-5_7-Figure9.4-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002979_978-3-540-70534-5_7-Figure9.4-1.png", + "caption": "Figure 9.4: Mecanum principle, vector decomposition", + "texts": [ + " For the construction of a robot with four Mecanum wheels, two left-handed wheels (rollers at +45\u00b0 to the wheel axis) and two right-handed wheels (rollers at \u201345\u00b0 to the wheel axis) are required (see Figure 9.3). Omni-Directional Drive 149 Although the rollers are freely rotating, this does not mean the robot is spinning its wheels and not moving. This would only be the case if the rollers were placed parallel to the wheel axis. However, our Mecanum wheels have the rollers placed at an angle (45\u00b0 in Figure 9.1). Looking at an individual wheel (Figure 9.4, view from the bottom through a \u201cglass floor\u201d), the force generated by the wheel rotation acts on the ground through the one roller that has ground contact. At this roller, the force can be split in a vector parallel to the roller axis and a vector perpendicular to the roller axis. The force perpendicular to the roller axis will result in a small roller rotation, while the force parallel to the roller axis will exert a force on the wheel and thereby on the vehicle. Since Mecanum wheels do not appear individually, but e.g. in a four wheel assembly, the resulting wheel forces at 45\u00b0 from each wheel have to be combined to determine the overall vehicle motion. If the two wheels shown in Figure 9.4 are the robot\u2019s front wheels and both are rotated forward, then each of the two resulting 45\u00b0 force vectors can be split into a forward and a sideways force. The two forward forces add up, while the two sideways forces (one to the left and one to the right) cancel each other out. 9.2 Omni-Directional Drive Figure 9.5, left, shows the situation for the full robot with four independently driven Mecanum wheels. In the same situation as before, i.e. all four wheels being driven forward, we now have four vectors pointing forward that are added up and four vectors pointing sideways, two to the left and two to the right, that cancel each other out" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000881_robot.2004.1308085-Figure6-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000881_robot.2004.1308085-Figure6-1.png", + "caption": "Fig. 6. The simple pendulum performs harmonic oscillations for", + "texts": [ + "n of the mass and the inertia of the system is determined by the weight distribution on the robot. If T is the wheel radius, &I is the lumped mass of the system, and p is the radius of gyration of the system with respect to an axis passing through the contact point and perpendicular to the plane, the time-period for small amplitudes~ is where g is gravity. Note that T, decreases as T increases, and T, increases as p increases. If the Pivoting Dynamics model is restricted to oscillate about the X or Y axis, then the Pivoting Dynamics model is similar to a simple pendulum (see Fig. 6). whose time-period is Tap = 2i8, where p is the radius of gyration, and g is gravity. The time-period Tsp decreases as p decreases. Note that to get similar oscillatory behavior between the eccentric mass wheel and the simple pendulum, a rearrangement of masses may he required. Table I shows the time-periods for X and Y rotations for the RRRoboton-a-plane model and the Pivoting Dynamics model. Since the time periods of the two models are close to each other, we do not rearrange the masses. MODEL A N D THE PIVOTING DYNAMICS MODEL I X Rotations (sec) I Y Rotations (sec) RRRobot-on-a-olane I 1" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001632_pac.2005.1591644-Figure4-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001632_pac.2005.1591644-Figure4-1.png", + "caption": "Figure 4: World x-y plane section at beam entrance to yoke.", + "texts": [ + " This introduces a small dipole component that corrects for a slight angular offset from the horizontal (\u03b2). To avoid compound error in 400 circulation passes per extraction cycle, we optimize electromagnetics for the circulating beam in the ring (CB). To this end, at beam entrance and exit the yoke design features iron squared to the CB path rather than the extracted beam (EB) path, in order to minimize quadrupole affects on the CB. Provision is also made to capture the CB within a shield assembly consisting of the \"septum plate\" and the \"shield plate\" (Fig.4). \u201cClamshelled\" between these plates the CB vacuum chamber is wrapped in a thin non-magnetic shimstock to provide an effective impedance to flux to which the CB would otherwise be exposed. A narrow \"shield cap\" runs along the CB axis flush with the shield assembly. The total iron thickness (minus CB pipe) yields a flux density within iron saturation limits, both at the entrance in the thin septum, and near the exit despite return path asymmetry from the CB outboard of the ELS centerline. Along the beam direction the iron overhang shields the \u201csandwiched\u201d CB beyond the ELS fringe field" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003570_cefc.2010.5480331-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003570_cefc.2010.5480331-Figure1-1.png", + "caption": "Fig. 1 Topology of 12-stator-slot/10-pole HEFS machine", + "texts": [ + " INTRODUCTION Flux-switching permanent magnet (FSPM) motor is a relatively novel brushless machine. However, since the opencircuit air-gap flux is produced by the stator-magnets solely, it is difficult to be regulated. Hence, recently a novel HEFS machine, which introduces field-excitation windings into FSPM motor, is proposed [1]-[3]. In this paper, the fluxregulation capability of the HEFS machine with three typical types of magnets materials, namely NdFeB, SmCo and Ferrite, are compared based on finite element (FE) analysis. Fig. 1 shows the topology of a 12-stator-slot/10-rotor-pole HEFS machine, in which the magnets, concentrated armature and field windings are located in the stator with a simple and robust rotor. According to the value and polarity of the fieldcurrents in the field-excitation windings, the air-gap flux can be modulated to be strengthened or weakened. The topology and operation principle can be found in [3] in detail. To investigate the flux regulation capability of the HEFS motor, Fig. 2 compares the open-circuit phase fluxes due to magnets or field windings alone, respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002532_n02-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002532_n02-Figure1-1.png", + "caption": "Figure 1. The NISE system.", + "texts": [ + " It is useful if the implantation can be checked both from ultrasound 0031-9155/08/040035+05$30.00 \u00a9 2008 Institute of Physics and Enginering in Medicine Printed in the UK N35 images and visually at the same time to understand what really takes place at each stage of the implantation process. It is also a good idea to be able to see the final seed distribution and to compare it with the plan. For these testing and training purposes we developed and built an implantation ultrasound phantom. The NISE system (figure 1) consists of two similar templates connected by rails through which the needles are set. The rear template is moveable and has a locking mechanism for needles. The system is connected to a stepper unit and a TRUS transducer. After dose planning, the needles are set through the templates to the selected coordinates (X, Y ) not penetrating the patient\u2019s skin. The z-coordinate (depth) for each needle is taken into account so that the needles are set relatively to each other. This can be done for example by measuring the distances between the needle butt (or a specific mark on the needle) and rear grid with a ruler" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003833_ijsurfse.2010.030488-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003833_ijsurfse.2010.030488-Figure1-1.png", + "caption": "Figure 1 Gear device (Flender Graffenstaden)", + "texts": [ + " In the same time, the fluid viscosity is largely modified by pressure and the behaviour of the fluid is not Newtonian. In that case the applied load could be less than one hundred Newton (100 N) but the contact surface is very small. The distinction between distributed contacts and hertzian contacts is only formal; for example, the elastic displacements should be taken into account to calculate the behaviour of connecting-rod journal bearings. The same discussion can be carried on for elastomeric lip seals where the elastic displacements are of the same order of magnitude as the film thickness. Figure 1 shows a gear box in which the different types of contacts are present. Journal and thrust bearings are representative for distributed contacts and the contacts between gear teeth correspond to the hertzian model. Stribeck (1902) presented experimental results on friction in plain and rolling-element bearings. His results were summarised on curves, which show the variation of friction vs. relative speed. These curves are very general and can be used to present the different phenomena that occur in lubricated contacts" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003779_gt2010-22877-Figure11-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003779_gt2010-22877-Figure11-1.png", + "caption": "Fig. 11 Static Pressure contours and streamline distribution in the first bristle pack of the brush seal with 0.3mm clearance", + "texts": [ + " Brush seal 2 with shim installed shows no influence on the leakage flow in the bristle pack by comparison with brush seal 1 at zero clearance size. At the 0.1mm sealing clearance, the similar flow pattern is observed for brush seals 1 and 2 as shown in fig. 10. In addition, similar flow fields are shown for brush seals 1 and 2 at zero and 0.1mm sealing gaps. As shown in Figs 11 and 12, a different flow pattern is observed for a clearance of 0.3mm and 0.5mm for brush seals 1 and 2. For brush seal 1, the leakage jets flow upwards at first and then rapidly downwards inside the bristle pack. Figure 11(b) shows that the leakage flow leaves the bristle pack at the tip and mixes with the clearance leakage flow. For brush seal 2, as shown in Fig. 11(b), part of leakage flows upward at first and then rapidly downward and leaves the bristle pack. The other part of leakage flow forms an approximately symmetrical parabolic line distribution. The shim installed between the front plate and bristle pack of brush seal 2 changes the flow pattern in the bristle pack. The corresponding pressure distribution along the bristle pack is also changed. At the 0.5mm sealing clearance, the different flow pattern is observed for brush seals 1 and 2(Fig. 12). Two approximately symmetrical parabolic lines in the bristle pack are observed for brush seal 1" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001157_vetecs.2006.1683417-Figure3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001157_vetecs.2006.1683417-Figure3-1.png", + "caption": "Fig. 3. Path between nodes D and A at metrical level", + "texts": [ + " All geometrical data are available at topological level by means of the link structure. Topological maps can be generated only once or updated like an evidence grid [9]. At topological level, the number of nodes in the map basically depends on the size and structure of the environment. Thus, any path planning algorithm like the A* can quickly process a topological route [6]. This route is propagated into regions at metrical level by means of the existing link structure. Then, a Potential Fields method is used within these regions to obtain a minimum cost path to the goal. Fig.3 shows a planning example. To travel from the departure to the arrival location, it is detected in the structure which nodes they are linked to (D and A respectively). Using a A* algorithm the best route to move between both nodes is computed in the topological map. This route is propagated down to metrical level, and a Potential Fields planner computes the final path over the bounding box at metric level enclosing the regions linked to the nodes of the topological route. (b) Fig.2: a) DGPS metric map of a parking lot; b) associated topological map and regions linked to each node Since slippage, sonar and DGPS errors and unexpected obstacles make strict path tracking unfeasible, every inflection point in the resulting path becomes a partial goal at metric level. The trajectory in Fig. 3, for example, presents an unique partial goal, plus destination. We let the reactive layer reach each goal sequentially, so it is modulated by the deliberative one, which, taking into account all available data on the environment, avoids local traps and provide efficient paths. Sometimes unexpected obstacles or external factors (e.g. traffic jam, accidents) may make a global route unfeasible, but if the map is modified on-line, the deliberative layer can be retriggered to calculate a new trajectory from the current position to the goal, as proposed in [6]", + " Our mutation is derived from the PFs: the output direction and repulsion vectors corresponding to all obstacles within the reaction range are combined to calculate a new solution. The casebase is pruned from time to time to remove all cases whose below a certain efficiency threshold. Usually, a few trajectories are enough to seed the database with knowledge about the working environment. Human/machine shared control is as simple as adding the control vectors from both sources, weighted if necessary depending on the degree of confidence on each source. Fig. 5 shows how the system tracks the trajectory in Fig. 3 both using the proposed CBR based technique and a conventional Potential Fields approach. In our experiments, 6 trajectories were supervisedly trained a priori by non-expert guiders, resulting in 624 cases in the casebase. The gray square in the Fig. 5.a represents a car parked within detection range when the robot tracks the path that was obviously not included in the processed metrical map. It can be noted that the proposed approach presents less oscillations and gets closer to the obstacle than the PFs based one" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000118_s0043-1648(03)00330-2-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000118_s0043-1648(03)00330-2-Figure1-1.png", + "caption": "Fig. 1. A cylinder between two planes.", + "texts": [ + " The friction model used is a two-dimensional brush model. This model, although further developed in this work, is general and has previously been used to simulate the friction in gears and cam-follower mechanisms (see [6]). \u2217 Corresponding author. Tel.: +46-8-790-78-38; fax: +46-8-20-22-87. E-mail address: chriss@md.kth.se (C. Spiegelberg). A simple configuration with a cylinder between two planes is studied. The three bodies in the system are denoted by A (upper plane), B (cylinder) and C (lower plane), as shown in Fig. 1. There are two contacts in the system. The contact between the upper plane (A) and the cylinder (B) is referred to as the upper contact, and the contact between the cylinder (B) and the lower plane (C) is referred to as the lower contact. The aim is to investigate the friction in the contacts and the motion of the roller when the planes are in motion. The friction in the contacts is modelled with a brush model. As can be seen from the equations of motion (Eq. (1)), the friction in the contacts together with the inertia determine the motion of the cylinder. The three coordinate systems shown in Fig. 1 are used to analyse the system. The n frame is global and fixed. The a frame is fixed to and translates with the centre of the cylinder, but it does not rotate relative to the n frame. The b frame is fixed to the cylinder and thus translates and rotates relative to the n frame. Two generalised coordinates s1 and s2 are used: s1 represents the position of the centre of the cylinder 0043-1648/$ \u2013 see front matter \u00a9 2003 Elsevier Science B.V. All rights reserved. doi:10.1016/S0043-1648(03)00330-2 Nomenclature a oscillation amplitude of upper plane (m) aH semi-Hertzian contact width (m) A dimensionless oscillation amplitude of upper plane d \u00af A dimensionless movement of upper plane d \u00af B,1 dimensionless movement of peripheral point on the cylinder E\u2217 combined elasticity of surfaces in contact (N/m2) f \u00af 1 dimensionless translatory natural frequency of cylinder f \u00af 2 dimensionless torsional natural frequency of cylinder F 1, F 2 friction force in upper and lower contact, respectively (N) F \u00af 1,F\u00af 2 dimensionless friction force in the upper and lower contact, respectively J inertia of cylinder (kg m2) K total stiffness of one contact (N/m) Kt representation of the tangential stiffness of the surface (N/m3) l cylinder length (m) m cylinder mass (kg) n1, n2 unit vectors N1, N2 normal load in upper and lower contact, respectively (N) N \u00af 1,N\u00af 2 dimensionless normal load in the upper and lower contact, respectively P normal load of contact (N) qx traction (N/m2) Qx friction force calculated with the brush model (N) Qx,C friction force acc" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001809_2005-01-0384-Figure4-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001809_2005-01-0384-Figure4-1.png", + "caption": "Fig. 4 FE Models of Suspensions", + "texts": [ + " 2 that the calculated coefficient of friction of the cornering force 2005-01-0384 Vehicle Cornering and Braking Behavior Simulation Using a Finite Element Method Tatsuya Fukushima, Hitoshi Shimonishi and Toshikazu Torigaki Nissan Motor Co., Ltd. Takahiko Miyachi and Yasuyoshi Umezu The Japan Research Institute, Limited showed good agreement with the measured data for both small and large slip angles and for both small and large belt loads. A sedan type vehicle was modeled with finite elements as shown in Fig. 3. The link parts of the front and rear suspensions, suspension members, as shown in Fig. 4, and the vehicle body were modeled as rigid bodies, because deformation of a running vehicle was regarded as being smaller than that of the tires and kinetic motion of the suspension systems. The beam elements in Fig. 3 were grouped into one rigid body to express the total inertia of the body, engine, power-train unit, parts in the cabin and the other parts mounted on the body suspended by the suspension systems. To the vehicle model shown in Fig. 3 the outer skin panels of the body were mounted to the group of beam elements to visualize the behavior of the vehicle" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003592_1.4001258-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003592_1.4001258-Figure1-1.png", + "caption": "Fig. 1 Geometry of the manifold", + "texts": [ + " CFD-FE coupled simulations Contributed by the Heat Transfer Division of ASME for publication in the JOURAL OF THERMAL SCIENCE AND ENGINEERING APPLICATIONS. Manuscript received January 5, 2009; final manuscript received November 5, 2009; published online April 23, 010. Assoc. Editor: Thomas Trabold. ournal of Thermal Science and Engineering Applications Copyright \u00a9 20 om: http://thermalscienceapplication.asmedigitalcollection.asme.org/ on 0 of exhaust manifold were performed for investigating the transient thermal stress behavior during the entire manifold run cycle. The current model is shown in Fig. 1. It is a liquid-cooled exhaust manifold, which is used for a 12.5 l six-cylinder diesel marine engine. This type of the engine is typically installed in tag boats and fishing boats. 2.1 Load Analysis. In reality, three kinds of loads are applied on the manifold. The manifold suffers from relatively high operational temperatures, which can lead to significant thermal expansion. Since the inner structure of the manifold is strictly constrained, expansion causes remarkable thermal stress. The manifold is also subjected to mechanical stresses induced by the vibration of the exhaust system" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001579_j.jappmathmech.2006.11.007-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001579_j.jappmathmech.2006.11.007-Figure1-1.png", + "caption": "Fig. 1.", + "texts": [ + " One algorithm is obtained by synthesizing a traditional two-pulse optimal control and a local bang-bang stabilization algorithm.10 The other algorithm provides non-local stabilization of the system.11 Similar problems were considered previously in Refs. 10\u201314. Prikl. Mat. Mekh. Vol. 70, No. 5, pp. 801\u2013812, 2006. E-mail address: mefremov@relex.ru (M.S. Yefremov). 0021-8928/$ \u2013 see front matter \u00a9 2006 Elsevier Ltd. All rights reserved. doi:10.1016/j.jappmathmech.2006.11.007 Consider a spacecraft, represented by a rigid body with two dynamic elastic elements (rods) (Fig. 1). We will assume that identical viscoelastic rods are fastened symmetrically and that the oscillations of the elastic structure are small. We will only consider the controlled motion around the longitudinal axis of the spacecraft. Suppose the rods execute antisymmetric oscillations. We can therefore confine ourselves to considering only one rod, and the effect of their interaction on the main body is doubled. A similar problem has already been considered in the literature (see, for example, Ref" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002979_978-3-540-70534-5_7-Figure14.4-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002979_978-3-540-70534-5_7-Figure14.4-1.png", + "caption": "Figure 14.4: Puma 560 simulation and conceptual drawing", + "texts": [ + " We start by introducing the standard manipulator joints and their unambiguous drawing norm. There are three basic types or manipulator joints (see Fig. 14.3): rotational joints with the rotation axis along the link, rotational joints with the rotation axis perpendicular to the link (hinge joints), and prismatic joints (telescopic joints). All other joints, e.g., a more complex ball joint, can be described as a combination of these basic types. The Puma 560 from Unimation/St\u00e4ubli is a standard 6-dof manipulator that is frequently used as a model in textbooks. Figure 14.4 shows a simulation screenshot in RoboSim [Br\u00e4unl 1999] of this robot and its conceptual drawing, labeling its joints \u03b81, .., \u03b86. Manipulator kinematics deals with answering the following two basic questions for a specified manipulator geometry: This will be answered by forward kinematics. This will be answered by inverse kinematics. Kinematics 209 There are of course a number of additional and more complex questions in manipulator kinematics. Below are a few examples, but these are beyond the scope of this chapter" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003303_j.ejc.2009.09.007-Figure12-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003303_j.ejc.2009.09.007-Figure12-1.png", + "caption": "Fig. 12. Proof of Example 7.", + "texts": [ + " So, Theorem 3 seemingly contradicts their result, but it does not, since our origami-deformations \u2018fold-out\u2019 and \u2018fold-in\u2019 are not immersions. Corollary 3. From a pyramid, remove the bottom face and cut off the remaining convex point. Then the resulting polyhedral surface is s-reversible. Proof. By making many \u2018pleats\u2019, we can change the shape of the surface into a (part of) rectangular tube. By s-reversing this tube, and then by unfolding the pleats, we can s-reverse the original surface. There are s-reversible polyhedral surfaces that are not tube-like. Example 7. From a box, remove a face and then cut off 4 convex points, see Fig. 12 top-left. The resulting surface is s-reversible. To reverse this surface, first subdivide the surface as in Fig. 12 top-right. Then by repeating foldout-operations, make the surface very short. Then, we can push down the \u2018ceiling\u2019. Finally, by pullout-operations, we get the surface reversed. Theorem 4. The surface obtained from an s-reversible polyhedral surface M by applying a tubeattachment operation is also s-reversible. Proof. By repeating fold-out-operations, make the attached tube very short, and flatten it on the face, see Fig. 13. Then the resulting surface is regarded as a part of M , and we can s-reverse it" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003308_s1064230710040143-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003308_s1064230710040143-Figure1-1.png", + "caption": "Fig. 1. Schematic diagram of man\u2013machine interface for realization of haptic and visual interaction with virtual object.", + "texts": [ + " In all these cases special man machine interface should be used which provides the possibility of realistic haptic per ception of the so called computer synthesized virtual object. The basic functional unit of the haptic interface is a special multidimensional controlled manipulation interaction mechanism. It is equipped by sensors that determine the value of joint coordinates of the mech anism, their velocity, and then the spatial position of the free end of the manipulator and its velocity. More over, the mechanism is equipped with the sensor (Fig. 1) measuring the force and time instant of inter action of the human hand with the handle fastened at the free end. Removable models of virtual objects can be used as the handle. The shape, size, and texture of these mod els are similar to those of the simulated virtual objects. The mass inertia characteristics and weights of these models should be as small as possible, it is desirable that they should be smaller than the corresponding characteristics of simulated objects. In international practice, this interaction mecha nism is called Haptic display", + " If the system is stable, the acceptable qual ity of transition processes at perturbations, i.e., admis sible process establishment time, small \u201covercontrol\u201d, absence of oscillations, should be provided. For more realistic character of perception, it is desirable to add these feelings with visual feelings received using the \u201cvisual\u201d part of the man\u2013machine interface [1]. The latter generates realistic image of the virtual object in the position and orientation coincid ing with that of the model. The man perceives this image via two displays in the helmet for the left and right eyes (Fig. 1). The aspect angle and scale of the image correspond to the current position and orienta tion of the human head. The task of analysis of the control system for inter action mechanism drives is quite complicated due to high order of the differential equation representing the model of the interaction process. This model covers the description of the behavior dynamics of the inter action mechanism with the chosen control law and the virtual object moving under the action of the applied force and torque", + "10c) JTF and F are the vectors of generalized forces applied to the model and reduced to the generalized coordi nates included in the blocks g and w, respectively; Kg, K\u03c4, Kw are the positive definite symmetric constant matrices of friction coefficients; U is the six dimen sional control vector, i.e., the vector of forces formed by the drives of the mechanism reduced to the coordi nates of the block g; Ed(q, , ), E\u03c4(q, , ) are the (6 \u00d7 18) blocks of the Euler operator of the Lagrange function L = T \u2013 \u03a0. Assuming, as shown in Fig. 1, that the rotors of the drive motors do not possess linear velocities, but just rotate about the motionless axes, the total kinetic energy T of mobile elements of the mechanism can be represented as (1.11) where , , and \u03b8m are the inertia matrices of \u0440th rotor, motor of the \u0440th link, and model respectively; and Mm are the masses of the \u0440th link and the model, respectively; and = , and are the vectors of angular velocities of the \u0440th rotor of the \u0440th link and the vector of linear velocity of the \u0440th link in coordinate systems of the \u0440th rotor and \u0440th link, respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003799_kem.417-418.313-Figure5-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003799_kem.417-418.313-Figure5-1.png", + "caption": "Fig. 5 Fatigue crack growth curve of the rail with a squat defect (initial crack size=5.0mm)", + "texts": [ + " The fatigue crack growth during every train passage can be calculated from the relation between the effective stress intensity factor range, \u2206Keff and crack growth curve of rail material as following; q p eff R K Cd da )1( / \u2212 \u2206 = (3) where C, p and q are empirical constants that depend on material. In this research, the material constants are obtained from the reference [5] as C=7.63\u00d710-10 cyclemmmMPa /\u2212 , p=3.45 and q=1.63. The \u2206Keff can be calculated from the crack size, a and stress range, \u2206\u03c3 as following; )( 12.1 1 kE a MKeff \u03c0\u03c3\u2206=\u2206 (4) 2 2/ 0 22 )/(1 ,sin1)( cakdkkE \u2212=\u2212= \u222b \u03b8\u03b8 \u03c0 where the M1 is an empirical factor to account for the finite dimensions of the rail cross section and the E(k) is the elliptical integral of second kind. Fig. 5(a) shows the squat defect model for crack growth analysis. Fig. 5(b) is the crack growth curves calculated by supposing the initial crack sizes as 5mm. Estimation of critical crack size and crack growth rate is an essential to prevent rail from failure and develop cost effective railway maintenance strategy. In this study, we predict crack growth rate of a rail with a squat defect by considering the stress components in rail head. From the analysis results, we conclude that a rail with squat which has 5mm initial crack size may not be rapidly propagated until it reaches much longer crack in normal operating condition" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003122_978-90-481-3141-9_4-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003122_978-90-481-3141-9_4-Figure1-1.png", + "caption": "Fig. 1 Illustration of procedure to obtain delamination beam specimens from filament wound composite cylinders", + "texts": [ + " Delamination specimens were machined from composite cylinders consisting of E-glass fibers in an epoxy resin. The internal diameter of the cylinders was 160 mm and the nominal wall thicknesses were 6 and 12 mm (12 and 24 layers). The lay-ups of the cylinders were \u0152\u02d9\u2122 6 and \u0152\u02d9\u2122 12, where \u2122 D 30\u0131; 55\u0131 and 85\u0131. A 58 mm long and 13 m thick, release agent coated aluminum film, was inserted at the mid-plane of the cylinders during the filament winding process to define a starter delamination crack, see Fig. 1. The film insert was wrapped around the circumference of the cylinder to enable machining of multiple test specimens from each cylinder. After filament winding, the cylinders were cured at 160\u0131 C for 3 h. The average fiber volume fraction was 0.61 for all cylinder lay-ups. Beam fracture specimens of a nominal length of 200 mm and a nominal width of 18 mm were cut from the cylinder wall for the subsequent delamination tests as schematically illustrated in Fig. 1. The beam axis was parallel to the cylinder axis producing straight beams with a curved cross-section. Figures 2\u20134 show the DCB, ENF and MMB cylinder specimens and loading principles. In order to accommodate the curved cross-section of the beam fracture specimens, contoured aluminum loading tabs were fitted to the DCB and MMB specimen and contoured loading pins and supports were attached to the ENF and MMB test fixtures as shown in Figs. 2\u20134. Further experimental details are provided in Refs. [10\u201312]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001205_pime_proc_1972_186_090_02-Figure3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001205_pime_proc_1972_186_090_02-Figure3-1.png", + "caption": "Fig. 3. Sliding apex mechanism", + "texts": [ + "comDownloaded from ROTARY PISTON MACHINE SUITABLE FOR COMPRESSORS, PUMPS AND I.C. ENGINES 745 Fig. 4. Precessing rotor with four working chambers necessarily limited for space. However, it avoids the use of gears. The motion of the rotor resembles that of precession with $ = -24 and the possible applications overlap with that case. ment devices by providing means for altering the angles. In the former the angle between the shafts is variable and in the latter the angle between the shaft and the normal to the face F is variable. 2.3 Sliding apex The third mechanism, Fig. 3, causes the rotor to have exactly the same motion as in the Hooke\u2019s coupling but the constraint of the second shaft is replaced by sliding contact between the flat circular face, F, and the cylindrical apex, C, of the rotor. This contact is practical in some cases because the inertial loads transmitted through the contact can be made small and fluid pressures in the two chambers exert no extra loads at that point. The avoidance of a second shaft and hub sphere clearly simplifies the construction enormously when compared with that of Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000204_naecon.2000.894996-Figure5-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000204_naecon.2000.894996-Figure5-1.png", + "caption": "Figure 5 - Manipulation of cosine finetion: (a) originnl~finctiorz; (b) inverted; (e) amplitude added", + "texts": [ + " The critical points which are required to be free of discontinuities, occur at the end of segment A and the beginning of segment C, both high speed transitions. The solution to this problem is found with the cosine profile. This ultra-smooth solution provides for an extended performance envelope in regards to the craft\u2019s ability to withstand higher winds and greater noise relative to other profiles. The manner in which the base profile is obtained involves first inverting the cosine function and then adding a value equal to the function\u2019s amplitude (figure 5 ) . After manipulation of equations, the acceleration and deceleration profiles can be described by 3 fundamental variables, the peak velocity ( Vp), the duration of acceleration (tA), and the duration of deceleration (tc). How these variables are determined is dependant upon which of three specific cases exists. The cases in question are entirely dependant on geometry between the two points in question which the craft is navigating between. The first case is when the path is considered \u201clong\u201d. This is the initial assumption until proven otherwise" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002115_s00170-007-1350-z-Figure6-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002115_s00170-007-1350-z-Figure6-1.png", + "caption": "Fig. 6 a Meshed body and boundary conditions. b Temperature field inside the body for t=te", + "texts": [ + " So the imposed boundary conditions were identical to the ones adopted in the development of the analytical model for the upper flat end surface: krT\u00f0 \u00de \u00bc Q\u00fe H Tgas T where Q \u00bc F0 for t te 0 for t > te \u00f021\u00de while the other surfaces were considered as adiabatic: krT\u00f0 \u00de \u00bc 0 \u00f022\u00de x = 0,1,2 mm x = 4,6,8 mm 1400 The rod was meshed in linear solid tetrahedral elements with eight nodes. The thermo-physical properties (k, \u00f1, cp and T0) were the same used in the analytical model and are listed in Table 1. In Fig. 6.a the meshed body and the boundary conditions are reported, while in Fig. 6.b the temperature field of the body when t = te is shown. In Fig. 7 the comparison between the analytical and the numerical model is reported for small (x=0, 1, 2 mm) and for large distances (x=4, 6, 8 mm) from the upper surface. As can be seen, there is a fairly good relationship between the analytical and the numerical results for small (x=0, 1, 2 mm) and for large (x=8 mm) distances from the upper end surface, whereas for intermediate distances (about 4 and 6 mm) the analytical model underestimates the temperature" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000845_00022660410526033-Figure7-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000845_00022660410526033-Figure7-1.png", + "caption": "Figure 7 Ideal roll", + "texts": [ + " The boundary conditions in entry section are t \u00bc 0; vx \u00bc 0; _vx \u00bc 0 t \u00bc t1; vx \u00bc vxc; _vx \u00bc 0 Thus, it is assumed roll that the angular velocity in this section is vx1 \u00bc a1t 3 \u00fe b1t 2 \u00fe c1t \u00fe d1 0 , t , t1 \u00f026\u00de Roll angular velocity in this section is vx2 \u00bc vxc t1 , t , t2 \u00f027\u00de The boundary conditions in exit section are t \u00bc t2; vx \u00bc vxc; _vx \u00bc 0 t \u00bc t3; vx \u00bc 0; _vx \u00bc 0 therefore roll angular velocity is vx3 \u00bc a2t 3 \u00fe b2t 2 \u00fe c2t \u00fe d2 \u00f028\u00de Using the above relations, kinematic parameters in roll maneuver can be calculated. The entry Mathematical modeling of helicopter aerobatic maneuvers Yihua Cao, Guangli Zhang and Yuan Su Aircraft Engineering and Aerospace Technology Volume 76 \u00b7 Number 2 \u00b7 2004 \u00b7 170\u2013178 and exit transient proportional factor may be taken as 0.1 or so. In order to calculate the time histories of the load factor, the functions of simulating the flight path and velocity are required. Here the ideal roll is considered (Figure 7). The velocity can be expressed by a fifth-order polynomial, similar to that in a loop maneuver. In an ideal roll, if the roll angle gs around the helicopter longitudinal axis is assumed to be harmonized with the phase angle ur of ideal roll in air, the following relation can be gained ur\u00f0t\u00de \u00bc gs\u00f0t\u00de \u00bc Z t 0 vx\u00f0t\u00dedt \u00bc Z t 0 _ur\u00f0t\u00dedt or _ur\u00f0t\u00de \u00bc vx\u00f0t\u00de; ur\u00f0t\u00de \u00bc gs\u00f0t\u00de \u00f029\u00de At any time, helicopter\u2019s circumferential velocity is Ut \u00bc _ur\u00f0t\u00deRe \u00bc vx\u00f0t\u00deRe \u00f030\u00de Thus the load factor is nT\u00f0t\u00de \u00bc ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 1\u00fe _u 2 rRe=g 2 \u00fe2 _u 2 rRe=g cos ur r \u00f031\u00de An articulated helicopter performs a roll maneuver at initial velocity V 0 \u00bc 260 km=h; with the minimum velocity Vmin \u00bc 204 km=h; steady roll angular velocity vxc \u00bc 1:22 rad=s; the transient proportional factor Km \u00bc 0:1 and the effective radius Re \u00bc 11:8m: The following results are obtained: t1 \u00bc 1 s; t2 \u00bc 5 s; t3 \u00bc 6 s, i" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003845_s10409-009-0254-6-Figure7-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003845_s10409-009-0254-6-Figure7-1.png", + "caption": "Fig. 7 Contour of deformation results in x-direcion under different compressive forces. a 1.25 kN; b 1.75 kN; c 2 kN; d 2.5 kN", + "texts": [ + " The critical collapse force is determined by the following method. The curve of compressive force versus x-direction deflection is shown in Fig. 6 where the compressive force is increased from zero to 2.5 N and the other parameters are fixed. The compressive force corresponding to the crossing point of two tangent lines is defined as the critical collapse force [14] in this analysis. In Fig. 6, the critical collapse force is 2.022 N. For the case of 0.015 m initial radius and 1.5 \u00d7 105 Pa pressure, the deformed contour results are shown in Fig. 7. The large deflection in x direction and the bending of the longitudinal inflated booms can be obviously observed from our results. Especially at the loading action point, the significant inverse deformation occurs under high compressive force. On the other hand, the hoop inflated toruses have an inconspicuous configuration change. The stress fields are shown in Fig. 8. The shear stress occurs mainly on the hoop inflated toruses, and the normal stress occurs mainly on the longitudinal inflated booms" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001692_acc.1995.529782-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001692_acc.1995.529782-Figure2-1.png", + "caption": "Figure 2: Layout of the various components in the OH-58A gearbox. The figure also shows division of the gearbox into subsystems for diagnosis.", + "texts": [ + " Diagnosis is performed by propagating the n flagged values of the vibration features f i ( t ) through the SBCN, and obtaining as outputs the fault possibility values associated with individual gearbox components as: n p k ( t ) = x f i ( t ) w i k (1) i= 1 where the Wik represent the weighting factors determined based on the lower and upper bounds of the fuzzy influences ( l i k and U ( k ) between the itk sensor and kth component as: w i k = l i k + ( U i k - k k ) . f i ( t ) - (2) In SBCN, in order to make uniform interpretation of the fault possibility values p k ( t ) , they are normalized to have values between 0 an 1 as: (3) 3. EXPERIMENTAL The effectiveness of the SBCN was demonstrated using vibration data from an OH-58A helicopter main rotor gearbox (see Fig. 2). Vibration data was collected at the NASA Lewis Research Center as part of a joint NASA/Navy/Army Advanced Lubricants Program. Various component failures in an OH-58A main rotor transmission were produced during accelerated fatigue tests [3]. The vibration signals were recorded from eight piezoelectric accelerometers (frequency range of up to 10 KHz) using an FM tape P. 2 1 624 recorder. The signals were recorded once every hour, for about one to two minutes per recording (using a bandwidth of 20 KHz)", + " These paths consisted of: (1) Duplex Bearing to Triplex Bearing through Spiral Bevel mesh, (2) Duplex Bearing to Ring Gear through the Sun-Planet mesh, (3) Mast Roller Bearing to Mast Ball Bearing through the Main Shaft, (4) Ring Gear to Mast Ball Bearing through Planet Bearing, and ( 5 ) Duplex Bearing to Mast Ball Bearing through the Sun-Planet mesh. Based on the lumped mass model of these paths, the RMS values were calculated for excitation sources at each gearbox component. The fuzzy influences between each of the components and the accelerometers were then obtained using these RMS values. Diagnosis of the OH-58A gearbox was performed in three different hierarchies. In the top hierarchy, faults in the three subsystems of the gearbox (see Fig. 2) were isolated. The fuzzy influences between the three subsystems and the eight accelerometers were obtained by averaging the influences of the components in each subsystem, as shown in Table 1, and were incorporated as the weights of the top SBCN subsection. The inputs to this sub-section of SBCN were the averaged values of all abnormal features from each aecelorometer, and its outputs were the fault possibility values for the three Subsystems. In the second hierarchy, faulty component families (gear and bearing) in each subsystem were isolated" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001164_1-84628-559-3_20-Figure6-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001164_1-84628-559-3_20-Figure6-1.png", + "caption": "Fig. 6. Phtograph of the modularized actuating unitFig. 4. Sectional view of the internally mouted actuating unit.", + "texts": [ + " For examples, the actuating units set on the top side of xy plane and marked with (a) and (b) are used for actuating the stage to rotate with respect to z-axis; simlarily, the actuating units set on the xz plane and marked with (f) and (e) are used for actuating the stage to rotate with respect to y-axis as shown in Fig. 4 which is the sectional view A-A in Fig. 3. Figure 5 shows the perspecitve schematic drawing of the semi-shperical 3-DOF positionig stage. The friction adjusting mechanism set on the top surface of the stage is used for providing a suitable holding fricitonal force for the rotaional stage. Figure 6 shows a photograph of the modualrized springmounted PZT actuator, in which the PZT actuator has the dimension of 5 5 10 mm (Tokin). The stiffness of the spring is 0.023 N/mm. Six actuating units are symmetrically mounted to the positioning stage having a radius on the bottom side. The positioning stage made of stainless steel having a mass of 2.6 kg is then set on the base having the same radius on the contace surfaces. Figure 7 shows the positioning stage and the base. Figure 8 shows the 3-DOF positioning stage" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003695_17543371jset49-Figure3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003695_17543371jset49-Figure3-1.png", + "caption": "Fig. 3 Forces acting on the oar and gate", + "texts": [ + "comDownloaded from forces and system mass\u2019. FX Gate is the component of the gate force acting in the fore\u2013aft direction, which is determined from the handle force by applying the no-slip assumption. The blade is assumed to rotate about a point which remains stationary in the axis of the boat\u2019s forward movement. This is the no-slip condition and is used to calculate the relationship between the gate force FGate and the handle force FHandle by considering moments about the centre of the action of FBlade [10] (see Fig. 3). Recorded handle force data are used as an input to the model according to FGate \u00bc FHandle LOar lOut \u00f06\u00de LOar \u00bc lIn \u00fe lOut \u00f07\u00de FHandle \u00bc FPeak Handle sin pt t1 ; 0< t6 t1 \u00f08\u00de 0; t1 < t6T \u00f09\u00de 8< : The five-segment model proposed by Caplan and Gardner [15] is extended to a seven-segment model by including the head, with X position assumed to coincide with the shoulders, and the feet assumed to be fixed relative to the boat. The masses of each of these body parts as a portion of total athlete mass m are given by mFoot \u00bc 0:0145m mShank \u00bc 0:093m mThigh \u00bc 0:2m\u00femSeat mTrunk \u00bc 0:578m mArm \u00bc 0:056m mForearm \u00bc 0:044m mHead \u00bc 0:0145m \u00f010\u00de The mass of the seat is included with the mass of the thigh since the motion of the seat is nearly coincident with the motion of the hip" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001871_iecon.2006.347447-Figure3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001871_iecon.2006.347447-Figure3-1.png", + "caption": "Fig. 3. Equivalent circuit with u1 being applied", + "texts": [ + " The signals from di/dt sensors reflect the anisotropy of rotor, which introduces unbalance in the phase values of the stator leakage inductances. The leakage inductances may be assumed to vary according to: )ncos(lll anana \u03b8\u03940\u03c3 += (1) ))(ncos(lll ananb 3\u03c02\u03b8\u03940\u03c3 \u2212+= (2) ))(ncos(lll ananc 3\u03c04\u03b8\u03940\u03c3 \u2212+= (3) where l0 is the average inductance, and \u0394l is its variation caused by the rotor anisotropy (nan = 2 for saturation induced anisotropy or nan= nrs =Nr/p for rotor slotting, where Nr is rotor slot number and p the pole pairs). The equivalent circuit when vector u1 is applied to the machine is shown in Figure 3. For this case the following equations can be derived: )u( a )u( ab as )u( abd e td id lriU 1 1 \u03c3 1 ++= (4) )u( b )u( bc bs )u( bc e td id lri 1 1 \u03c3 10 ++= (5) )u( c )u( ca cs )u( cad e td id lriU 1 1 \u03c3 1 ++=\u2212 (6) where Ud is the DC link voltage, rs is the stator per phase resistance, ea is the back emf in phase a, iab etc are the phase currents in the delta connected machine, and the subscript u1, u0 etc relates to the switching vector imposed at that instant in time. Similar, if inactive voltage vectors u0 or u7 are applied, the following equations can be derived: )u( a )u( ab as )u( ab e td id lri 0 0 \u03c3 00 ++= (7) )u( b )u( bc bs )u( bc e td id lri 0 0 \u03c3 00 ++= (8) )u( c )u( ca cs )u( ca e td id lri 0 0 \u03c3 00 ++= (9) If the sampling instants, at which u0 and u1 are applied, are close, it can be assumed that: )u( a )u( a ee 10 \u2248 , )u( b )u( b ee 10 \u2248 , )u( c )u( c ee 10 \u2248 Additionally, the voltage drops across the stator resistance can be ignored due to their small values compared with Ud" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001563_095440605x8397-Figure3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001563_095440605x8397-Figure3-1.png", + "caption": "Fig. 3 Coordinate systems for surface generation", + "texts": [ + " IMechE Vol. 219 Part C: J. Mechanical Engineering Science at RMIT UNIVERSITY on July 12, 2015pic.sagepub.comDownloaded from As reflected in Fig. 1b, the surfaces of the special rack cutter can be acquired by revolving the profile in Fig. 1c about axis B\u2013B with a radius of rc. The surfaces of the special rack cutter with respect to coordinate system Sc(XcYcZc) can then be represented by R(0) c \u00bc + d \u00fe u tana\u00f0 \u00de u\u00fe rc\u00f0 \u00de cosb rc u\u00fe rc\u00f0 \u00de sinb 2 4 3 5 (2) where b represents the revolving angle, shown in Fig. 3a, and is also a surface parameter of the special rack cutter. The normal vector pointing to the inward side of generating surface S0 can be determined with N(0) c \u00bc @R(0) c @u @R(0) c @b \u00bc + rc \u00fe u\u00f0 \u00de (rc \u00fe u) cosb tana (rc \u00fe u) sinb tana 2 4 3 5 (3) Thus, the unit normal vector is n(0)c \u00bc + cosa cosb sina sinb sina 2 4 3 5 (4) To derive mathematical models for the surface generation and curvature analysis, coordinate systems are defined in Fig. 3. Coordinate systems Sc and Ss(XsYsZs) are attached to the special rack cutter and the gear blank respectively, and Sh(XhYhZh) becomes the fixed coordinate system. When Ss is rotated with an angle u about axis Zs, Sc will be translated a distance rgu along the direction of the Xc axis. By using homogeneous transformationmatrices Proc. IMechE Vol. 219 Part C: J. Mechanical Engineering Science C11604 # IMechE 2005 at RMIT UNIVERSITY on July 12, 2015pic.sagepub.comDownloaded from [20], if matrix Mij denotes the coordinate transformation from system Sj(XjYjZj) to system Si(XiYiZi), the matrices for coordinate transformation from coordinate system Sc to Sh, from coordinate system Ss to Sh, and from coordinate system Sc to Ss can then be represented respectively by Mhc \u00bc 1 0 0 rgu 0 1 0 rg 0 0 1 0 0 0 0 1 2 6664 3 7775 (5) Mhs \u00bc cos u sin u 0 0 sin u cos u 0 0 0 0 1 0 0 0 0 1 2 6664 3 7775 (6) Msc \u00bc cos u sin u 0 rg( sin u u cos u) sin u cos u 0 rg( cos u\u00fe u sin u) 0 0 1 0 0 0 0 1 2 6664 3 7775 (7) According to the theory of gearing [21], the meshing equation can be expressed as n(0)h V(0,1) h \u00bc 0 (8) where n(0)h is the unit normal vector of the cutting tool surface S0 with respect to the fixed coordinate system Sh and is identical to that of the generated surface at the contact point; V(0,1) h then becomes the relative velocity between the generating and the generated surfaces at their points of contact" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003363_j.ejc.2009.11.023-Figure3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003363_j.ejc.2009.11.023-Figure3-1.png", + "caption": "Fig. 3. A non-rigid toroidal surface with pentagonal meridian.", + "texts": [ + " For the purpose of insight view, the rotational parameter v is taken to be less than 2\u03c0 . Fig. 1 shows the initial surface from example 1without any deformations \u03b5 = 0.0 and v \u2208 [0, 43\u03c0 ]. Fig. 2 shows the deformed surface from example 1 with deformations \u03b5 = 0.15 and v \u2208 [0, 43\u03c0 ]. 2. The second example of a non-rigid toroidal surface is a surface with a convex pentagon with apices A(\u22121, 2), B(\u22122, 3), C(0, 4), D(2, 3) and E(1, 119\u2212 \u221a 2641 45 ). The polygon rotates around the u-axis of the coordinate system uO\u03c1. Fig. 3 shows the initial surface from example 2without any deformations \u03b5 = 0.0 and v \u2208 [0, 32\u03c0 ]. Fig. 4 shows the deformed surface from example 2 with deformations \u03b5 = 0.25 and v \u2208 [0, 32\u03c0 ]. 3. The third example of a non-rigid toroidal surfaces is a surface with nonconvex polygon with 9 apices A(\u22122, 2), B(\u22124, 3), C(\u22123, 4), D(\u22122, 6), E(0, 8), F(3, 12), G( 72 , 3839 304 + (\u2212173450877+836 \u221a 43583479149)1/3 3042/3 \u2212 50005 304(3(\u2212173450877+836 \u221a 43583479149))1/3 ),H(0, 3) and I(\u22121, 1). The polygon rotates around the u-axis of the coordinate system uO\u03c1" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003497_cnsr.2010.26-Figure4-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003497_cnsr.2010.26-Figure4-1.png", + "caption": "Figure 4. Relationship among mo", + "texts": [ + " A complete node in the network is shown in figure 3. It includes Agent Execution Environment (AEE) and Mobile Agents (MA), over a sensor node. The agent execution environment provides the conditions necessary for agent execution and migration. Database structure is present in the AEE at sensor nodes. The updations from low-level to high-level MA are in the form of DB queries. DB is shown in figure 3 and it is updated on the basis of the relationship between mobile agents of different hierarchal levels, shown in figure 4. EDB is created at deployment time by the administrator and maintained by querying between the MAs. Mobile agents appear in following two roles. 3.2.1. Sniffers: Sniffers observes the behavior of the network at different levels. A sniffer is essentially a software agent that migrates from one node to the other (to make observations) when required, and decide whether the nodes visited are faulty or not. 3.2.2. Correctors: Correctors are the repair agents present at different levels in the network", + " Metropolitan Agents (MA): Metropolitan Sniffers (MSs) and Metropolitan Correctors (MCs) are initially deployed at the gateways. Here the activities of the entire super-cluster are monitored. There the EDB is present which assists the metropolitan agents in fault diagnosis at the larger level. 3.3.3. Global Agents (GAs): Global agents are deployed at the sink. Global sniffer (GS) and global corrector (GC) have the view of the whole network from the sink. Here the entire EDB is present which carries about faults and resulting failures possible. The hierarchical relationship between for communication is shown in figure 4 detecting fault informs correctors about t fault f. Corrector after correcting the fault to sniffer. There is a possibility that a failu by the sniffer but it is unable to detect fault. In this case, it intimates the sniffer n that it has encountered an unknown fault. hierarchy then replies with the solution of t Faults may occur in variety of ways in proposed approach we have classified categories. This is the logical categoriz irrespective of their nature. A fault may permanent based on duration or it ma omission, timing or arbitrary fault depen nature thereof as classically defined" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000911_tia.2004.834136-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000911_tia.2004.834136-Figure1-1.png", + "caption": "Fig. 1. Typical circuit breaker unit in GIS switchgear.", + "texts": [ + "16-kV GIS equipment was higher than for conventional MC switchgear. However, the purchase price for the 26.4-kV GIS equipment was significantly lower than the price of conventional MC switchgear. After consideration of the other benefits associated with GIS equipment, the overall evaluation favored the use of the GIS switchgear. V. GIS GIS differs greatly from traditional MV MC switchgear widely used in the U.S. A functional schematic view of one pole of a typical unit of GIS switchgear is shown in Fig. 1. As in the familiar MC switchgear, vacuum circuit breakers are used for interruption. MV GIS switchgear differs from high-voltage GIS switchgear in that the SF gas is used for its insulating properties, not for interruption. Conventional MC switchgear relies on a combination of air and solid insulating materials, but GIS switchgear uses bare bus conductors on insulating supports, immersed in insulating gas. Hence, GIS switchgear is much smaller than equivalent MC switchgear. Since SF gas provides the major insulation, it is not feasible to incorporate drawout provisions for the circuit breaker" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002858_s0001925900005047-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002858_s0001925900005047-Figure1-1.png", + "caption": "Figure 1. Satellite attitude description in terms of Euler's characteristics. XYZ are body-fixed axes (principal axes of inertia), xyz is the fixed frame of reference.", + "texts": [ + "1017/S0001925900005047 Downloaded from https://www.cambridge.org/core. Tufts Univ, on 24 Jun 2018 at 09:42:10, subject to the Cambridge Core terms of use, available at DYNAMICS OF A SPINNING SATELLITE \u2212\u23a9 . . . (6) 1 1 2 2 1 1 2 2 A A A A ( ) ( ) ( ) ( ) yc yc yc zc zc zc a t a t a a t a t a = \u00b5 + \u00b5 = \u00b5 + \u00b5 . . . (7) } } where UM denotes the fuzzy variable of the control signal u. Each linguistic term is associated with a fuzzy set as shown in Fig. 5. The parameters of these sets are determined by the optimisation algorithm. A scaling factor is also used to normalise the fuzzy sets. The fuzzy sliding surface of a second-order system is shown in Fig. 6, where \u03b71 represents the scaling factor. For normalised fuzzy sets, S1,NB = \u20131 and S1,PB = 1. The maximum control energy umax is also bounded by physical limitations, that is U1,NB = \u20131 and U1,PB = 1. If the symmetry of fuzzy terms corresponding to s and UM is assumed, the remaining design factors are S1,PM, U1,PM, and the scaling factor \u03b71. Therefore, the optimisation parameters corresponding to the main controller of the first phase include S1,PM, U1,PM, \u03b71, and the positive constant \u03bbM. The lead angle method, used in the midcourse phase, may occasionally guide the missile near the boundaries of the tracking beam" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000327_20.950991-Figure5-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000327_20.950991-Figure5-1.png", + "caption": "Fig. 5. Schematic of the slider-disk contact in a quasistatic deformation process under impact (O is initial contact point,O is the point with maximum compression).", + "texts": [ + " However, the disk wear does not occur until another number of L/UL operation which could be the critical cycles. Fig. 3 shows the degraded 0018\u20139464/01$10.00 \u00a9 2001 IEEE slider and worn disk surface image measured by the scanning reflectance analyzer. after 72 000 ramp load/unload cycles. The detailed picture is shown as Fig. 4. Consider the corner of a slider impact with disk, the impact can be regarded as an impact of elastic spheres with finite radii of curvature. The schematic of the contact and the compressing process of spheres with radii of curvatures and are plotted in Fig. 5. A quasistatic approximation based on static contact elastic theory is used to deal with the elastic impact of spheres. Consider a normal contact without tangential velocity components at the contact point, the kinetic energy of the slider at the starting instant of contact is where and are the approaching velocity and the effective mass of slider, . Assume is the instant contact force between the spheres and is the instant compression, their relationship can be represented as [5], where is the composite Young\u2019s modulus of elasticity for the slider and disk" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000801_j.jmatprotec.2004.07.013-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000801_j.jmatprotec.2004.07.013-Figure1-1.png", + "caption": "Fig. 1. Illustration of search coil locations in the section of the test lamination.", + "texts": [ + " Laminations of non-oriented, 0.65 mm thick electrical steel stamped with typical induction motor stator geometry E-mail address: mosesaj@cardiff.ac.uk (A.J. Moses). were given a decarburising anneal prior to assembling in a 924-0136/$ \u2013 see front matter \u00a9 2004 Elsevier B.V. All rights reserved. oi:10.1016/j.jmatprotec.2004.07.013 pack around 50 mm in height. Epstein strips of the same material (800-65-D5) were found to have loss of 4.5 W/kg at 1.3 T, 50 Hz (sinusoidal flux) after a decarburising anneal. Fig. 1 shows a part of the stator laminations geometry. The outside and inside diameters were 126 and 92 mm, respectively and the tooth length was 20 mm, there are 36 teeth on each lamination. Orthogonal search coils are positioned as shown in the figure before placing the test lamination in the middle of a pack of seven identical laminations in a magnetising rig to form a thin section of the stator core. A three phase excitation winding was used to magnetise the laminations in a similar fashion to that occurring in an actual three phase, four-pole stator core assembled from the same size laminations" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002747_3.3801-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002747_3.3801-Figure1-1.png", + "caption": "Fig. 1 Simply supported beam resting on a hard spring.", + "texts": [ + ", displacement at i due to a force of unit amplitude at j , oscillating with circular frequency 12 -\u0302) amplitude of the external force at j time The force induced at 0 and the displacement w at 0 may be Received February 28, 1966; revision received May 6, 1966. * Research Specialist. *7 i TTtf run i _ . . . \u2014\u2014\u2014\u2014 , \u2014\u2014\u2014 \u0302 Equations (8) and (9) provide the desired response equation! at 0. The aforementioned technique first will be appliec to the case of a simple beam resting on a nonlinear spring a some position along the span, as shown in Fig. 1 . Using standard methods available in Structural Dynamics it mav be shown that sin(mra/L) \u2014 \u2014\u2014\u2014pL(l - rn 2 (10 where rn = Q/pn, pn being the natural frequency of the nil mode of a simple beam. Using Eqs. (10) and (8), shown to be, for the case a = b = L/2, (11 Defining the dimensionless parameters ai2 = p^/k/m (spring characteristic) y = kwbi/po (displacement) M = |e (p0//c)2 (characteristic force) D ow nl oa de d by U N IV E R SI T Y O F C A L IF O R N IA - D A V IS o n Fe br ua ry 1 0, 2 01 5 | h ttp :// ar c" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003709_icinfa.2010.5512052-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003709_icinfa.2010.5512052-Figure2-1.png", + "caption": "Fig. 2 gives the process of constructing and analyzing VP. Firstly, parameterized finite element model is created in finite element software and solved to generate model neutral files. Then the files are introduced into dynamic analysis software and flexible model is obtained. Following this approach, the VP is achieved after applying boundary conditions. At last, simulation and debugging are carried out and the results are analyzed. So the whole process is finished in three stages: construction of flexible model (Fig. 2a), application of boundary conditions (Fig. 2b) and simulation and analysis (Fig. 2c).", + "texts": [], + "surrounding_texts": [ + "The mesh of rotor, backplates and pads was generated by eight-node hexahedral element. Rotor was divided into 5942 1478978-1-4244-5704-5/10/$26.00 \u00a92010 IEEE elements and unilateral backplate and pad was divided into 5886 elements. Then the model was analyzed to obtain the information of modal coordinates, modal transformation matrix, modal stiffness matrix, modal mass matrix and modal frequency. Finally, this information was introduced into dynamics analysis software MSC.ADAMS and flexible model was gained, as shown in Fig. 3." + ] + }, + { + "image_filename": "designv11_69_0003062_9780470447734.ch1-Figure1.2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003062_9780470447734.ch1-Figure1.2-1.png", + "caption": "FIGURE 1.2 Crystal faces of a highly ordered crystal of graphite and the formation of an edge-plane step defect.", + "texts": [ + " It has been shown that when |n m|\u00bc 3q, where q is an integer, the CNT is metallic or semimetallic and the remaining CNTs are semiconducting [7,8]. Therefore, statistically, one-third of SWCNTs are metallic depending on the method and conditions used during their production [7] and can possess high conductivity, greater than that ofmetallic copper, due to the ballistic (unscattered) nature of electron transport along a SWCNT [9]. If one considers the structure of a perfect crystal of graphite, two crystallographic faces can be identified, as shown in Figure 1.2. One crystal face consists of a plane containing all the carbon atoms of one graphite sheet, which we call the basal plane; the other crystal face is a plane perpendicular to the basal plane, which we call the edge plane. By analogy to the structure of graphite, two regions on a CNT can be identified (and are labeled in Figure 1.1) as (1) basal-plane-like regions comprising smooth, continuous tube walls and (2) edge-plane-like regions where the rolled-up graphite sheets terminate, typically located at the tube ends and around holes and defect sites along tube walls" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001950_1-84628-179-2_5-Figure5.5-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001950_1-84628-179-2_5-Figure5.5-1.png", + "caption": "Fig. 5.5. Control direction for a tandem helicopter.", + "texts": [ + " To minimize the aerodynamic interference created by the operation of the rear rotor in the wake of the front, the rear rotor is elevated on a pylon (0.3r to 0.5r above the front rotor). In a tandem rotor helicopter, pitch moment is achieved by differential change of the main rotors thrust magnitude (by collective pitch), roll moment is controlled by lateral thrust tilt using cyclic pitch (Figure 5.4), yaw moment is obtained by differential lateral tilt of the thrust on the two main rotors with cyclic pitch (Figure 5.5), and the vertical force is achieved by the change of the main rotor collective pitch. For simplicity we will present here the dynamic model of a tandem main rotor helicopter in hovering. We propose a dynamic tandem helicopter model based on Newton\u2019s equations of motion [59] with the assumptions of the standard helicopter with the following changes: 1T The nose rotor blades are assumed to rotate in an anti-clockwise direction when viewed from above and the tail rotor blades rotate in a clockwise direction, see Figure 5" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002896_1448837x.2008.11464166-Figure4-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002896_1448837x.2008.11464166-Figure4-1.png", + "caption": "Figure 4: Load current and voltage waveforms for the RBHCC with bipolar", + "texts": [ + " As shown in fi gure 3(a), two hysteresis controllers are used to generate random band heights. In this section, a hysteresis band height of unipolar and bipolar modulation schemes are calculated in terms of system parameters using mathematical analysis and the results are compared. control. waveforms. block diagram of random band controller, and (b) load current and voltage waveforms. E08-PE01 Nami.indd 2 30/1/08 4:57:29 PM Australian Journal of Electrical & Electronics Engineering Vol 4 No 1 Figure 4 shows the load current and voltage waveforms for the bipolar hysteresis band modulation. According to this fi gure, the rate of load and reference current changes can be expressed as follows: * 1 1 1 2 2a a adi di di diHB HB T T dt T dt dt dt a (1) 2 * 2 2 22 T dt diT dt diHB dt di T HB dt di aaaa (2) dc a V dt diL (3) dc a V dt diL (4) where L is the load inductance, and i+ a and ia are the rising and the falling segments of the load current, respectively. With regard to fi gure 4: c c f TTT 1 21 (5) where T 1 and T 2 are switching intervals, and fc is the switching frequency" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000123_s00604-003-0064-7-Figure5-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000123_s00604-003-0064-7-Figure5-1.png", + "caption": "Fig. 5. Cyclic voltammograms of the resulting modified electrode (15 mL of a mixture viologen-dehygrogenase spread and dried under vacuum on a 5 mm carbon disk: see experimental section) obtained by controlled potential electrolysis for 20 minutes at 0.9 V (Q\u00bc 1.96 mC) and immersed in Tris buffer solution (0.2 M, pH\u00bc 7.7) containing 0.2 M NaHCO3 and oxoglutarate (12 mM). (A) under argon; (B) under CO2", + "texts": [ + " The absence of absorbance increase with time at 340 nm clearly indicates the efficient immobilization of isocitrate dehydrogenase and the absence of a phenomenon of enzyme release. The poly 1-isocitrate hydrogenase electrode was immersed in a Tris buffer solution (0.2 M, pH 7.7). The cyclic voltammogram of this enzyme electrode clearly shows the reversible one-electron reduction of the polymerized viologen groups. Oxoglutarate (12 mM) was added to the buffer. Under argon, the electroactivity of the enzyme electrode remains unchanged (Fig. 5A). In contrast, in the presence of CO2, an increase in cathodic and a decrease in anodic currents of the V2\u00fe=V \u00fe system is observed (Fig. 5B). This electrocatalytic phenomenon indicates that isocitrate dehydrogenase can catalyze the fixation of CO2 in oxoglutarate using the cation radical form of the polymerized viologen groups. It should be noted that these experiments were repeated with three enzyme electrodes providing a similar electrocatalytic phenomenon in the presence of CO2. The replacement of CO2 by argon leads to the complete disappearance of this electrocatalytic behavior, the cyclic voltammogram of the enzyme electrode being identical to the initial one before CO2 bubbling" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002819_6.2007-3813-Figure26-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002819_6.2007-3813-Figure26-1.png", + "caption": "Figure 26. The Effect of Slot Pressure.", + "texts": [ + " Tangential Jets As described above, the model projectile was modified such that a rearward facing slot was created that could provide a tangential aft jet. Figure 25 shows that the force and moment generated were relatively small except at Mt = 0.72 for which a significant force was developed as well as a nose up moment. Both the force and moment generally increased with an increase in slot pressure, but there a reduction in the effectiveness of tangential blowing occurs at the highest jet pressures. The pressure contours shown in Figure 26 help explain the data in Figure 25 as it appears that the primary effect of the tangential blowing is to increase the pressure immediately behind the slot. This pressure increase is likely due to the jet reducing the separation behind the slot, which was the intended effect. At some point the jet becomes too strong to remain attached to the body and thus the benefits of tangential blowing begins to diminish as is evidenced by the pressure contours in Figure 26 and the data in Figure 25. The fact that there were only limited changes in the force at lower tunnel Mach numbers may also be due to the fact that the separation was not as strong at these lower Mach numbers. As described above, both a round and a rectangular pin were manufactured that could be inserted into any of the pressure tap locations. The pressures generated on the model were measured when the rectangular pin was inserted into Tap 14 for various tunnel Mach numbers as shown in Figure 27" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003815_s147355041000025x-Figure6-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003815_s147355041000025x-Figure6-1.png", + "caption": "Fig. 6. Hands", + "texts": [], + "surrounding_texts": [ + "The presented model of life sees the human body as a rationally generated structure, independent of the environment, rather than the result of random errors in the genome. Apart from the philosophical implication of the self-organizational origin of life, this understanding can lead to progress in various areas of science (Kauffman, 1993; Pivar, 2009). In medicine, the mechanically causative theory of disease is a promising field. Self-organization presents the possibilities of embryonic plastic surgery to prevent birth defects, as well as for aesthetic advantages. The technology of tissue engineering may benefit from being informed as to the natural directional vectors of tissue growth, leading to the synthesis of entire organisms." + ] + }, + { + "image_filename": "designv11_69_0000966_978-1-4419-8887-4_6-Figure6.11-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000966_978-1-4419-8887-4_6-Figure6.11-1.png", + "caption": "Figure 6.11. Artificial Athlete Stuttgart.", + "texts": [ + " It does, however, provide a standard test procedure for comparison of different surfaces. Chap. 6. Shoe-Surface Interaction in Tennis 139 The rotational friction is influenced by the properties of the court surface and the shoe sole and also by the area of contact between the shoe and surface. Rotational friction can be quantified using a test device that measures the resistance to an applied constant torque (Stuttgart device), or the measurement of the torque required to rotate a weighted foot from a stationary position (BS7044). A typical test setup is illustrated in Fig. 6.11. The mechanical test procedures described have been adopted by sports governing bodies, such as the I.T.F., because they provide repeatable tests that are relatively straightforward to perform. However, the relevance of the tests depends on whether the results correlate with human interaction with the tennis court surface. One limitation of the majority of standard mechanical test procedures used for tennis surfaces is that they typically do not allow testing of the influence of different tennis shoes on the surface", + " In contrast, Krabbe, Farkas and Baumann (1992) reported that the estimation of internal loads highlighted differences in ankle joint loading on different surfaces, with frictional properties of the surfaces and running style of the subject being identified as influential factors. It therefore seems that our understanding of the force reducing ability of tennis court surfaces during human interaction needs to be improved if specific shoe-surface characteristics that are desirable are to be identified. Figure 6.11. CoIJection ofbiomechanical data for a tennis shoe-surface combination. An interesting finding in the Dixon and Stiles (2003) study was that, although changes in the surface had no influence on impact force, a change in footwear was influential. Two different tennis shoes were worn on all the Chap. 6. Shoe-Surface Interaction in Tennis 143 surfaces - one a basic model and the other a top of the range shoe from the same manufacturer. It was found that the impact force consistently exhibited a lower rate of loading for the basic shoe model", + " This adaptation to a change in surface by adjustment of movement may increase the energy cost of performing the task, it may reduce the effectiveness of the performance and it may also influence the chances of injury. The ability of humans to adapt to different levels of friction provided between the shoe and surface has also been demonstrated in a recent study by Dura, Hoyes, Martinez and Lozano (1999). However, these authors suggest that performance is not influenced adversely. The performance of a 180-degree turn was compared on five different playing surfaces. It was found that, despite differences in coefficient of friction indicated by the mechanical test procedures illustrated in Fig. 6.11; the time taken to change direction on the different surfaces was comparable. For surfaces with relatively high coefficients of friction, the time taken in the braking phase of Chap. 6. Shoe-Surface Interaction in Tennis 145 the turn was highest. This allowed time for more knee flexion than that observed for the surfaces with lower coefficients of friction. The authors suggested that this increase in knee flexion is a protective mechanism against potential high loads that may result from the limited sliding on these surfaces" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002691_pime_conf_1967_182_422_02-Figure23.1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002691_pime_conf_1967_182_422_02-Figure23.1-1.png", + "caption": "Fig. 23.1. Notation", + "texts": [], + "surrounding_texts": [ + "method of solution of the elastohydrodynamic problem may be summarized as follows : (1) The bearing pressures were derived from the Reynolds equation assuming a constant film shape. The finite difference equation is given in Appendix 23.1. (2) The pressure distribution was expressed as a double Fourier series. (3) Using distortion coefficients derived for a unit non-dimensional pressure amplitude, the total distortion at each point of the bearing was calculated and the results were added to the initial clearance to obtain the film shape. Step (1) was then repeated, and the reiterative procedure continued until a compatible film shape and pressure distribution were determined. The use of this method reduces the elasticity problem to one of determining the distortion coefficients for a given bearing geometry. This is carried out once at the beginning of the survey of results. Full details of the derivation of these coefficients are given in Appendix 23.2. Proc Instn Mech Engrs 1967-68 In order to obtain a stable convergence it was necessary to introduce a relaxation factor into the change of film shape for each reiteration. EXPERIMENTAL APPARATUS Fig. 23.2 shows the general arrangement of the journal bearing test machine. The shaft is driven by an electric motor through an infinitely variable-speed pulley system. A pressurized oil supply was used, the pump being a simple diaphragm pump operated from a cam on a secondary shaft. Fig. 23.3 shows the apparatus in more detail, The load was applied to the bearing by means of a link chain and a brass shoe. A shear pin between the drive coupling and the shaft allowed for possible bearing seizure. The shaft could be moved axially through the bearing by a small lead screw and cradle, the two roller bearing assemblies being fitted with linear bearings for this purpose. The shaft was fitted with an inductive film thickness transducer, a quartz crystal pressure transducer, and a thermocouple. The signals were carried out through a silver contact slip-ring assembly at the end of the shaft. Fig. 23.4 is a sectional view of the shaft showing the positions of the instruments in the shaft and a typical connection. The steel cover for the instruments was ground integral with the shaft to ensure no discontinuity in the shaft surface. The bearings were of 1 in diameter and 1 in long. The steel bearings with the radial clearances of 0.0015 and Vol182 Pt 3N at Gazi University on December 12, 2015pcp.sagepub.comDownloaded from 194 D. K. BRIGHTON, C. J. HOOKE, AND J. P. O'DONOGHUE 0.0056 in were tested to determine to what extent the published theoretical results agreed with experiment. The recommended radial clearance for bearings of this diameter is 0.0014 in (6). The Delrin bearing (bearing 3) had a radial clearance of 0.0012 in at room temperature. The change in clearance with temperature owing to thermal expansion was significant and was accounted for. Bearing running conditions Oil used SAE 30 Oil inlet pressure 4 Ibff:in2 Total maximum load bearing 1 170 lbf bearing 2 70 lbf bearing 3 220 Ibf Maximum bearing temperature 35\u00b0C (95\u00b0F) Ambient temperature (approx.) 25\u00b0C (77\u00b0F) Shaft speed 250 rev/min bearings 1 and 3 bearing 2 1000 rev/min" + ] + }, + { + "image_filename": "designv11_69_0002979_978-3-540-70534-5_7-Figure1.4-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002979_978-3-540-70534-5_7-Figure1.4-1.png", + "caption": "Figure 1.4: Braitenberg vehicles avoiding light (phototroph)", + "texts": [ + " Many biped robot designs have five or more motors per leg, which results in a rather large total number of degrees of freedom and also in considerable weight and cost. Braitenberg vehicles A very interesting conceptual abstraction of actuators, sensors, and robot control is the vehicles described by Braitenberg [Braitenberg 1984]. In one example, we have a simple interaction between motors and light sensors. If a light sensor is activated by a light source, it will proportionally increase the speed of the motor it is linked to. In Figure 1.4 our robot has two light sensors, one on the front left, one on the front right. The left light sensor is linked to the left motor, the right sensor to the right motor. If a light source appears in front of the robot, it will start driving toward it, because both sensors will activate both motors. However, what happens if the robot gets closer to the light source and goes slightly off course? In this case, one of the sensors will be closer to the light source (the left sensor in the figure), and therefore one of the motors (the left motor in the figure) will become faster than the other" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000260_tmag.2003.810424-Figure9-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000260_tmag.2003.810424-Figure9-1.png", + "caption": "Fig. 9. Showing contours ofA at 0.02 s after start.", + "texts": [ + " This is required as input data for the calculations. The total inertia of the rig during this test was 0.25 kg m . The measurements are compared with FEA and coupled circuit calculations in Figs. 5 and 6. The curve on Fig. 5 furthest away from the measurements is the coupled circuit result. Only the envelope of the currents is shown on Fig. 6. For clarity, a close up of the start of these transients are shown in Figs. 7 and 8. Note there is a mechanical resonance in the rig at about 16 rads/s. Fig. 9 shows contours of at 0.02 s after starting. The FEA results start with known circuit impedances, inertia, friction, and windage versus time and supply voltage, and the machine is accelerated according to the computed torque at each timestep, to get the computed machine speed. Measured and computed results agree reasonably well. Both FEA and circuit theory methods compare fairly well with measurements. It is hoped that the measured data in this paper can be used by others. [1] D. Rodger, H. C. Lai, and P" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001293_rob.20099-Figure5-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001293_rob.20099-Figure5-1.png", + "caption": "Figure 5. Step 2 of guidance example: a step 2.1 and b step 2.2, respectively.", + "texts": [ + " After the long-range motion of the platform, the guidance algorithm first relocates the platform so that at least one laser beam hits the center of a corresponding detector as close as random errors would allow . The subsequent corrective actions move the platform in such a way as keeping the first LOS hitting the center of its respective detector, while repositioning the platform such that the remaining LOS also hit their targets. The overall proofof-concept algorithm, thus, consists of three steps: Step 1 Figure 4 : 1.1 Determine the largest PSD offset, d1. 1.2 Translate the platform along the x or y direction by d1. Step 2 Figure 5 : 2.1 Measure the new offset, d2, along the same detector considered in step 1. Calculate the actual orientation angle of the platform from the x axis, a, and determine the difference = d\u2212 a, where d is the desired platform orientation. 2.2 Translate, by b, and, then, rotate, by , the platform about the center of the fixed side of the platform to compensate for d2 and . This movement should result in the LOS considered in step 1 to keep hitting the center of its detector and, furthermore, the platform achieving its final desired orientation" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001244_rd.166.0556-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001244_rd.166.0556-Figure1-1.png", + "caption": "Figure 1 System diagram for the cascaded nonrecursive equalizer,", + "texts": [ + " Several examples with numerical results obtained via an APL simulation program are also included '\"Some of the results contained in this paper were presented at the Sixth Annual Princeton Conference on Information Science and SYS1ems, March 23 - 24, 1972. to illustrate the effectiveness of the approach. It should be noted that base-band pulse transmission is assumed throughout. A number of related topics of practical im portance, such as time jitter, digital round-off errors, and other implementation details are beyond the scope of this paper. 2. Equalizer structures \u2022 Nonrecursive equalizer with cascaded stages The first equalizer structure considered in this paper con sists of n cascaded stages of transversal filters as shown in Fig. 1. The channel output feeds the first stage of the equalizer. The output of the first stage in tum feeds the second stage, and so on. Let us denote the channel im{ (O)} { (0)pulse response by the sample sequence J.Y.k = a_No , ,0) \\0 '0) 10)\"' \u2022 1 \u2022. '., a_ 1 ' (~o !, 0'1' , \u2022.\u2022, aN }. Similarly, the output of \u2022 \u2022 i { !il} {til tilthe lth stage IS denoted by Uk = a_No , .\u2022\u2022, et_ 1 ' I eto m, a/ii, . . \" aN(ii}. On the other hand, tap-gain settings \u2022 '. (il ( T be a sufficiently large, finite and fixed time. The problem of interest is to find a sequence of control inputs )}1(,),(),0({ \u2212Tutuu iii KK , and ii Utu \u2208)( , ]1,0[ \u2212\u2208\u2200 Tt , such that \u2126\u2208)(Tqi , Vi\u2208\u2200 and\u2126 describes a given convex and compact set for a desired consensus protocol" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003818_vss.2010.5544663-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003818_vss.2010.5544663-Figure1-1.png", + "caption": "Fig. 1. Helicopter mounted on a platform.", + "texts": [ + " The norm \u2016x\u20162, with x \u2208 R n, denotes the Euclidean norm and \u2016x\u20161 = |x1|+ |x2|+ \u00b7 \u00b7 \u00b7+ |xn| stands for the sum norm. The minimum and maximum eigenvalue of a matrix A \u2208 R n\u00d7n is denoted by \u03bbmin{A} and \u03bbmax{A}, respectively; and \u2016A\u2016 = \u2016A\u20161 = maxj \u2211n i=1 |aij | stands for the 1-norm. The vector sign(x) is given by sign(x) = [ sign(x1), . . . , sign(xn)]T where the signum function is defined as sign(xi) = 1 if xi > 0 [\u22121, 1] if xi = 0 \u22121 if xi < 0 \u2200xi \u2208 R. II. DYNAMIC MODEL AND PROBLEM STATEMENT The full mathematical model of the 3-DOF VARIO scale model helicopter, depicted in Fig. 1, can be described as follows [1]: c0z\u0308 = c8\u03b3\u0307 2u1 + c9\u03b3\u0307 + c10 \u2212 c7 + wz (1) m(\u03b3)\u03d5\u0308 = \u2212c4 [c12\u03b3\u0307 + c13] u1 + c5c11\u03b3\u0307 2u2 \u2212 c6[2c5\u03b3\u0307 + c4\u03d5\u0307] sin(2c3\u03b3)\u03d5\u0307 \u2212 c4 [ c14\u03b3\u0307 2 + c15 ] + w\u03d5 (2) m(\u03b3)\u03b3\u0308 = [ c1 + c2 cos2(c3\u03b3) ] [c12\u03b3\u0307 + c13]u1 \u2212 c4c11\u03b3\u0307 2u2 + c6 [( c1 + c2 cos2(c3\u03b3) ) \u03d5\u0307 + 2c4\u03b3\u0307 ] sin(2c3\u03b3)\u03d5\u0307 + [ c1 + c2 cos2(c3\u03b3) ] [c14\u03b3\u0307 2 + c15] (3) 978-1-4244-5831-8/10/$26.00 \u00a92010 IEEE where m(\u03b3) = c1c5 \u2212 c2 4 + c2c5 cos2(c3\u03b3). In the above equations, z \u2208 R is the height, \u03d5 \u2208 R is the yaw angle, \u03b3 \u2208 R is the main rotor azimuth angle and ci (i = 0, ", + " Substituting the above equation into (35) and if the desired trajectories and initial conditions are chosen in such a way that the terms including z\u0308d, \u03d5\u0308d, \u03d52 2, and \u03d52 can be neglected we have the following simplification: \u03b3\u03072 = a1\u03b3 2 2 + a2 + a3 \u03b32 + a4 \u03b32 2 = a1\u03b3 4 2 + a2\u03b3 2 2 + a3\u03b32 + a4 (37) where a1 = c\u22121 5 c14 = 2.415 \u00d7 10\u22124, a2 = c\u22121 5 c15 = 5.2914, a3 = \u2212(c5c8) \u22121(c0c12g) = \u2212518.8359, and a4 = \u2212(c5c8) \u22121(c0c13g) = \u2212432\u00d7 104 (see Table I). The solutions of (37) when \u03b3\u03072 = 0 are: \u03b3\u0307\u2217 1 = 355.03 [rad/s], \u03b3\u0307\u2217 2 = \u22124 \u00b1 380.99i [rad/s], \u03b3\u0307\u2217 3 = \u2212347.02 [rad/s]. Only the last of these values have a physical meaning for the system (see Fig. 1 for the rotation sense of main rotor). To analyze the stability of the zero dynamics, define \u03b3\u03032 = \u03b32 \u2212 \u03b3\u0307\u2217 3 (38) whose time derivative is \u02d9\u0303\u03b32 = a1\u03b3 2 2 + a2 + a3\u03b3 \u22121 2 + a4\u03b3 \u22122 2 = a1\u03b3 2 2 + a2 \u2212 |a3|\u03b3 \u22121 2 \u2212 |a4|\u03b3 \u22122 2 . (39) Rewriting the above expression in terms of velocity error \u03b32 = \u03b3\u03032 + \u03b3\u0307\u2217 3 we have \u02d9\u0303\u03b32 = a1(\u03b3\u03032 + \u03b3\u0307\u2217 3)2 + a2 \u2212 |a3|(\u03b3\u03032 + \u03b3\u0307\u2217 3 )\u22121 \u2212 |a4|(\u03b3\u03032 + \u03b3\u0307\u2217 3)\u22122. (40) The equilibrium points of the above equation are \u02d9\u0303\u03b3\u2217 1 = 702.05 [rad/s], \u02d9\u0303\u03b3\u2217 2 = 343.02 \u00b1 468.2i [rad/s], \u02d9\u0303\u03b3\u2217 3 = 0 [rad/s]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001318_elan.200603689-Figure3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001318_elan.200603689-Figure3-1.png", + "caption": "Fig. 3. Structure of the FI-ECL flow-through thin-layer cell.", + "texts": [ + " The gas pressure pump (in Figure 2) consisted of the pressure nitrogen bottle (A), the gas buffer chamber (B), and the regulator (C) with the water level, h, used for flow rate Electroanalysis 19, 2007, No. 2-3, 181 \u2013 184 J 2007 WILEY-VCH Verlag GmbH&Co. KGaA, Weinheim injection control. This pump had 1 % noise of the ECL intensity. The ECL detection unit consisted of an ECL flow cell, a photomultiplier tube ((PMT) H6780 \u2013 04, Hamamatsu, Japan), and a potentiostat (1112 Huso, Japan). Current at the PMT was converted into voltage, with a homemade current follower. The voltage was recorded with a computer through an AD-converter. The main body of the ECL cell (Figure 3) was made from a black \u201cT\u201d shaped three-way plastics tube (3.0 cm i.d. and 4.0 cm o.d.). A Pt counter electrode, an Ag jAgCl reference electrode and a stainless steel pipe serving as the solution outlet were mounted from the top of the cell. At the right side of the cell, a quartz window was mounted in front of the PMT. A Pt wire of 0.1 mm diameter and 10 cm long was reeled around the outer wall of the capillary (0.5 mm i.d. and 1.0 mmo.d.), and was coated with epoxy resin except for the surface near the tip of the capillary" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003746_s00170-008-1853-2-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003746_s00170-008-1853-2-Figure1-1.png", + "caption": "Fig. 1 Geometry of ball end milling", + "texts": [ + " The purpose of this paper is to study the dynamic behaviors of tool and the critical speeds of rotor system with the linear and nonlinear dynamic cutting forces for ball end milling during milling operation. The stability of rotor system is also investigated. Moreover, the relationship between critical speeds and natural frequencies is discussed with respect to various feed-per-teeth in linear and nonlinear rotor systems with the linear and nonlinear cutting forces. 2.1 Linear dynamic cutting force The linear cutting forces proposed by Lee and Altintas [1] are employed in this work. The detailed geometry of a helical ball end milling tool is shown in Fig. 1. The elemental tangential, radial, and axial dynamic cutting forces dFt, dFr, dFa acting on the cutter are given by dFt \u00bc Kte ds\u00fe Ktc t y i; bk db dFr \u00bc Kre ds\u00fe Krc t y i; bk db dFa \u00bc Kae ds\u00fe Kac t y i; bk db \u00f01\u00de where t y i; bk \u00bc h y i\u00f0 \u00de sinbk is the uncut chip thickness normal to the cutting edge, and the angle bk is defined in Nomenclature. h y i\u00f0 \u00de is instantaneous chip thickness, db \u00bc dz . sinbk is differential cutting edge length in which dz is differential height in axial direction of tool" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001486_j.enzmictec.2004.04.026-Figure10-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001486_j.enzmictec.2004.04.026-Figure10-1.png", + "caption": "Fig. 10. Hydrolysis of l-PheAFC (0.021 mM) with 1000 U -chymotrypsin in toluene: (a) l-PheAFC in toluene; (b) 3 h after enzyme addition; (c) 9 h after enzyme addition; (d) difference spectrum (a) \u2212 (c); (e) course of the reaction (detection of AMC, \u03bbEX = 365 nm/\u03bbEM = 430 nm).", + "texts": [ + " Their fluorescence maxima were located close to each other; a differentiation between the coumarine substrates and products during an enzymatic reaction in toluene was therefore complicated. 3.9. Enzymatic parameters of the hydrolysis of the coumarine substrates with \u03b1-chymotrypsin and PLE in toluene To investigate the kinetic parameters of the enzymatic hydrolyses in toluene, each coumarine substrate was cleaved with -chymotrypsin and PLE in toluene and KM and Vmax were determined respectively. Fig. 10 shows the 2Dfluorescence spectra of the hydrolysis of l-PheAFC with -chymotrypsin in toluene. Twelve milliliters of l-PheAFC (0.021 mM, 0.5% KPP) were placed within the reactor and the reaction was started with the addition of 1000 U - chymotrypsin. The temperature was 30 \u25e6C. Fig. 10a displays the fluorescence maxima of the AFCcoumarine substrate at \u03bbEX = 345 nm/\u03bbEM = 400 nm with an RFU of 238. Three hours after the enzyme addition, an increasing AFC product peak at \u03bbEX = 365 nm/\u03bbEM = 430 nm was visible. Nine hours after the start of the enzymatic reaction, the AFC-peak displayed an RFU of 1110, the time d e l o p a i t a t o d-PheAMC were not visible in the 2D-spectrums of toluene, heir peaks maxima were located to close to the scatter ight and could therefore not be detected" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002567_6.2007-2277-Figure3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002567_6.2007-2277-Figure3-1.png", + "caption": "Fig. 3 Boundary conditions for displacements and forces.", + "texts": [ + " (34)-(36), to give )]()()()([)( 443322113 \u03be\u03c6\u03be\u03c6\u03be\u03c6\u03be\u03c6\u03be AAAARS +++= (37) )]()()()([)( 43322112 \u03be\u03be\u03c8\u03be\u03c8\u03be\u03c8\u03be AAAARM +++= (38) where 3 3 032 2 02 , L c EIR L c EIR == (39) and )()( 41 \u03be\u03c6\u03be\u03c6 \u2212 , )()( 41 \u03be\u03c8\u03be\u03c8 \u2212 are functions of \u03be related to f1(\u03be)\u2212f4(\u03be), h1(\u03be)\u2212h4(\u03be) of Eqs. (34) and (36). Having developed the expressions for the bending displacement W, bending rotation \u0398, shear force S, and bending moment M, the dynamic stiffness matrix can now be formulated by applying the boundary conditions. Following the sign convention shown in Fig. 2 for positive shear force and bending moment, and referring to Fig. 3, the boundary conditions for the displacements and forces (for each of the two cases, i.e., 1=n and 2) are given as follows. Displacements: ==\u2212== ==== 22 11 ;:)1..( ;:)1..(0 \u0398\u0398\u03be \u0398\u0398\u03be WWceiLyAt WWeiyAt (40) Forces: \u2212=\u2212=\u2212== ==== 22 11 ;:)1..( ;:)1..(0 MMSSceiLyAt MMSSeiyAt \u03be \u03be (41) Substituting Eqs. (40) into Eqs. (34) and (35) gives the following matrix relationship. \u2212\u2212\u2212\u2212 \u2212\u2212\u2212\u2212 = 4 3 2 1 4321 4321 4321 4321 2 2 1 1 )1()1()1()1( )1()1()1()1( /)1(/)1(/)1(/)1( )1()1()1()1( A A A A chchchch cfcfcfcf LhLhLhLh ffff W W \u0398 \u0398 (42) or American Institute of Aeronautics and Astronautics 10 AH\u03b4 = (43) Similarly substituting Eqs" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002403_1.3601696-Figure21-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002403_1.3601696-Figure21-1.png", + "caption": "Fig. 21 Turbine-driven 6 0 K V A alternator \" t u r b o n a t o r \" developed by G.E. C o m p a n y", + "texts": [ + " Further simplifications of the system are possible by elimination of the reduction gear (Fig. 20(a)) and integrating the turbocompressor with the alternator (Figs. 20(6,c), [20]). At one time the power supply in the B-52 used bleed-off air from the main jet engines to power turbine-driven 60 K V A alternators. These APU's employed turbine-driven gear boxes and low-speed 6000 rpm alternators. In new applications it was desirable to reduce appreciably the weight of the system without sacrificing reliability. This was accomplished by the \"Turbonator\" design illustrated in Fig. 21. This design employs a directly driven 24,000 rpm, 60 K V A alternator. However, from rotor dynamics and temperature standpoint, a rolling-element bearing could not be used next to the turbine; thus a hybrid bear- 674 /' O C T O B E R 19 68 Transactions of the A S M E Downloaded From: http://tribology.asmedigitalcollection.asme.org/ on 01/27/2016 Terms of Use: http://www.asme.org/about-asme/terms-of-use ing was employed. This journal bearing used the same gas supply as the turbine. The bearing on the cold end of the machine was a rolling-element, bearing that took the radial and thrust loads" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002411_tt.39-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002411_tt.39-Figure1-1.png", + "caption": "Figure 1. Schematic of integrating sphere with sample", + "texts": [ + " A ball was held against a horizontally rotating shaft at one point (designated here as its north pole) and a measured volume of the lubricant solution was applied to the rotating surface from a gas-tight syringe. After the solvent evaporated, the weight gained by the ball was determined by a microbalance with an accuracy of \u00b12 \u00b5g. The smallest weight of lubricant that could be reliably quantified in this manner was in the 10 \u00b5g range. Published in 2007 by John Wiley & Sons, Ltd. Tribotest 2007; 13: 129\u2013137 DOI: 10.1002/tt Figure 1 shows a simplified schematic of the integration sphere with ball bearing sample. The intensity of the incoming beam is not uniform, but is a complex function of the spectrometer aperture and beam steering optics within the FT-IR bench. The beam-directing mirror is not located directly over the centre of the sample causing the beam to fall onto the sample at a small angle (on the order of 10\u00b0) from normal. For a flat sample at the reflectance port, the beam converges on the centre of the sample, but for three-dimensional samples, the higher the sample, the further the centre of the sample will be from the centre of the beam" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000428_6.1979-2033-Figure4-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000428_6.1979-2033-Figure4-1.png", + "caption": "FIGURE 4. THE KAWASAKI K A G - ~ ( ~ )", + "texts": [], + "surrounding_texts": [ + "References\nBelavin, N. I . , \"Ekranoplany\" (Shie ld P l anes ) , Sudostroyeniye P r e s s , Leningrad, 1968, 176 pp. (AD 694-582).\nd e B a r t i n i , R. L. , \"Tomorrow's Transpor t\" , Sov ie t Union, 1974, No. 12, pp. 50-51.\nCzerniawski, S., \"Ekranoloty ZSRR\" (Screen Planes of t h e USSR), Skrzydla ta Polska , August 14, 1977, pp. 8-9.\nMantle, P. J., \"Background t o A i r Cushion Vehicles\", Hovering C r a f t & Hydrofoil , Vol . 15 , No. 2, November, 1975, pp. 5-16.\n\"Ram-Wing Veh ic l e s as Water-Borne Transpor t F a c i l i t i e s \" , Hovering C r a f t and Hydrofoi l , Vol. 14 , No. 2, November, 1974, pp. 11-17. O r i g i n a l source: Budownic tiwo Okretowe, No. 3, 1974.\nMcLeavy, R., e d i t o r , J a n e ' s Su r f ace Skimmers. Hovercraf t and Hydro fo i l s , 11 th ed., Sampson Low, Marston & Co., Ltd . , London, 1978; passim. Refer t o Hennebutte, X-114, ESKA, Seag l ide , and Lo ckheed . Ol'shamovskiy, S.V., \"Navigation and Rules of Navigation on In land Wa terw.zysl', Transpor t Publ i sh ing House, Moscow, 1975, p. 293.\nR i t c h i e , M. L., he Research and Development Methods of Wilber and O r v i l l e Wright\", Astron a u t i c s and Aeronaut ics , Vol. 16, No. 7/8, July/August , 1978, pp. 56-67.\nFil ipchenko, G. G . , \"Ekranoplan KAG-3 i yego i spytaniya\" (The KAG-3 and its t r a i l s ) , Katera i Yakhty, No. 15 , 1968, pp. 21-22. NISC. T rans l a t i on No. 3737.\nMcLeavy, J ane ' s S u r f a c e Skimmers, 6 t h ed . , 1972-73; p.29. Re fe r s t o X-113.\nGrunin, E. , \"Nad vodoy paryashchiy\" (Soaring above t h e water ) , Tekhnika molodezhi, No. 12 , 1974, pp. 30-34.\nPipko, D., \"Ekranoplans a r e Winged Ships of t h e Futurenn, Nauka i zh izn , 1966, No. 1, pp. 33-41, t r a n s l a t e d a s FTD-ID (RS) T-1432-76.\nB. Blinov, \"Ekranoplan\" , I z o b r e t a t e l i Ra t s iona l i za to r , No. 3, 1965, pp. 18-19.\nKocivar, B., a am-wing X-214 F l o a t s , Skims, and F l i e s \" , Popular Sc ience , Vol. 211, No. 6 , . 'December, 1977, pp. 70-73.\nMoore, J. W. , \"Conceptual Design Study of Powered Augmented Ram Wing-in-Ground E f f e c t A i r c ra f t \" , AIAA A i r c r a f t Systems and Technology Conference. August 21-23, 1978. AIAA Paper No. 78-1466.\nMcLeavy, J ane ' s Surface Skimmers, 9 t h ed., 1975-76; passim. Re fe r s t o Be r t i n v e h i c l e s and WSEV.\nBelavin, N. I. , \"~kranop lany\" (Screen p l a n e s ) , 2nd Ed., Sudostroyeniy P re s s , Leningrad, 1977,\nMcLeavy, J ane ' s Su r f ace Skimmers, 1st ed . , 1967-68; passim. Refers t o K4G-3, RAM 1 and RAM 2.\nBert i n , J . , \"Une nouvel le a v i a t i o n de t r a n s p o r t lourd par\" (A unique l a r g e t r anspor t a i r c r a f t ) . No. 46, 1973-4, pp. 2-8.\nCa lk ins , D . E., \"A F z a s i b i l i t y Study of a Hybrid Airsh ip Operat ing i n Ground ~ f f e c t \" , J o u r n a l of A i r c r a f t , Vol. 14, August, 1977, pp. 809-815.\n\"Aerofoil : A Marine-' Schif f ' d e r Zukunf t ? \" , Wehrtechnik, No. 11, November, 1975, p. 642.\nSchmit t , D. , \"Airf oi l-Flugboot ~ 8 r g II\", Flug Revue, No. 5, May, 1978, pp. 70-77.\nMcLeavy, J ane ' s Su r f ace Skimmers, 7th ed . , 1973-74. Refers t o Hennebut t e .\nP r i v a t e Communication, David Rousseau, Naval Ship Research and Development Center, Carderock, '\nMaryland, August 30, 1978..\n\"Energy Cost i n Transpor ta t ion\" , Science Dimension NRC, Ottawa, Canada, Vol. 6, No. 1, 1974, pp. 4-9.\nD ow\nnl oa\nde d\nby F\nre ie\nU ni\nve rs\nita et\nB er\nlin o\nn D\nec em\nbe r\n12 , 2\n01 6\n| h ttp\n:// ar\nc. ai\naa .o\nrg |\nD O\nI: 1\n0. 25\n14 /6\n.1 97\n9- 20\n33", + "FIGURE 1. T. KAARIO'S SURFACE-EFFECT VEHICLE \"AEROSANI\" NO. 8(1)\nKey: 1 -forward wings; 2-articulated controlling wings; 3-hull with driver cockpit; 4-side stabilizers; 5-rudder; 6-tail stabilizing beams with planes; 7-flap; 8-main lifting wing; 9-skis\nFIGURE 2. A SWEDISH WATER-BORNE WIG VEHICLE (\"AEROBOAT\") DEVELOPED BY TROENG\nIN THE LATE 193dl)\nKey: 1-hull with crew cabin; 2-floats; 3-stabilizer with controlling surfaces; 4-outboard engines; !&lifting wing; 6-propeller a ~ e ~ i b ! y\nD ow\nnl oa\nde d\nby F\nre ie\nU ni\nve rs\nita et\nB er\nlin o\nn D\nec em\nbe r\n12 , 2\n01 6\n| h ttp\n:// ar\nc. ai\naa .o\nrg |\nD O\nI: 1\n0. 25\n14 /6\n.1 97\n9- 20\n33", + "FIGURE 5. GENERAL ARRANGEMENT OF THE RFB X-113 AM AEROFOIL BOAT(^^)\nFIGURE 6. THE OllMF-2 SINGLE-SEAT WING-IN-GROUND EFFECT RESEARCH CRAFT(^)\nD ow\nnl oa\nde d\nby F\nre ie\nU ni\nve rs\nita et\nB er\nlin o\nn D\nec em\nbe r\n12 , 2\n01 6\n| h ttp\n:// ar\nc. ai\naa .o\nrg |\nD O\nI: 1\n0. 25\n14 /6\n.1 97\n9- 20\n33" + ] + }, + { + "image_filename": "designv11_69_0003700_tac.1964.1105701-Figure4-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003700_tac.1964.1105701-Figure4-1.png", + "caption": "Fig. 4-PoSsible minirnvm-energy trajectories.", + "texts": [ + " The state space can be divided into various regions, the largest of which is controllable only by a bang-coast-bang policy; there is no \\yay to ax-oid the extra 2E, energy expenditures. Xnother region has a minimum integrated power using a coast-bang policy; such srates must now suffer the additional cost of E,. For the remaining initial states, the bang-coast-bang policy uses less integrated po\\\\-er than the coast-bang; however, the extra turn-on packet may be larger than the saving in integrated poner. Since the remaining states are controllable by either type of policy, the additional energy for warmup may change the optimal policy for some of these states. Consider Fig. 4. Obx-iously, there is only one coast-bang policy that can control a given state. The unique coast-bang policy must then be compared with the minimum-energy bang-coast-bang policy, which is just that one which minimizes the integrated pon.ert since the warmup energy is the same for all. Thus the bang-coast-bang trajectory used is the one that was optimum in the absence of warmup energy. The energies for the t\\vo trajectories are (c.f. [A] j + 2E,. (12) Equating the right-hand members of both equations yields an espression for the locus of states that may be controlled either \\\\-a>- using the same energy" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000599_047168659x.ch4-Figure4.4-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000599_047168659x.ch4-Figure4.4-1.png", + "caption": "Figure 4.4 802.11b channel model.", + "texts": [ + " From the obtained results, it can be expected that the first step in the two-session correlation occurs at the boundary of the physical carrier sensing range (i.e., 180 m , PCS_range 200 m), while the second step (i.e., d(2, 3) approximately 300 m) occurs when the radiated energy from one session has almost no effect on the other. A more detailed discussion on the physical carrier sensing range can be found in [Anastasi et al. 2004]. To summarize the above results, in [Anastasi et al. 2004] the channel model shown in Figure 4.4 to characterize an IEEE 802.11b WLAN is proposed. The channel model describes how a transmitting station S affects the stations around it, depending on the distance, d, from S and the rate j ( j \u00bc 1, 2, 5.5, 11 Mbps) used by S for its transmissions. In the figure, Tx( j ) denotes the station S transmission range when it is transmitting at rate j. As shown in Figure 4.4, stations at a distance lower than Tx( j ) are able to receive the S data. Nodes at a distance Tx( j ) , d , 200 m are in the S physical carrier sensing range, and hence observe that the channel is busy when S is transmitting. Finally, nodes at a distance 200 d , 300 m are affected by the energy radiated by S. From this channel model it follows that the hidden-station phenomenon, as it is usually defined in the literature (see Section 4.2.1), is almost impossible with the above ranges. Indeed, the PCS_range is about twice TX_range(1), i" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001227_j.jsv.2005.04.023-Figure10-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001227_j.jsv.2005.04.023-Figure10-1.png", + "caption": "Fig. 10. Elastomer plates subjected to a shear test.", + "texts": [ + " 9c and d, which show the extracted real and imaginary parts of the complex stiffness by using relations (18)\u2013(21). The restoring force model proposed satisfactorily describes the experimental harmonic behaviour of the cylindrical mount under traction\u2013compression. The hardening\u2013softening behaviour of the loops plotted in Figs. 8 and 9a is mainly due to the geometric configuration imposed by the compression\u2013traction test. The shear tests were carried out on a pair of elastomer (also carbon black filled rubber) plates placed opposite each other, see Fig. 10. They were subjected to axial harmonic deflection having different amplitudes and forcing frequencies. Only a positive triangular deflection was applied for the quasi-static test (frequency of 0.01Hz and amplitude of 3.5mm). The dynamic tests were performed using a sine excitation with a (10\u2013100Hz) frequency range and an (0\u20131.0mm) amplitude range. The shear specimen contained more carbon black filler than the specimen used for the compression test. The measurement of the dynamic modulus shows that amplitude dependency is more pronounced for this specimen and that the evolution of the dynamic modulus function of the amplitude is substantially nonlinear, see Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001309_6.2006-6599-Figure4-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001309_6.2006-6599-Figure4-1.png", + "caption": "Figure 4. Body Axis Systems and Aerodynamic Angles", + "texts": [ + " x\u0307 = Ax + Bu state equation y = Cx + Du output equation (2) where x \u2208 R n\u00d71 is a state vector, A \u2208 R n\u00d7n is a plant matrix, u \u2208 R m\u00d71 is an input vector, B \u2208 R n\u00d7m is a control distribution matrix, y \u2208 R p\u00d71 is an output vector, and C \u2208 R p\u00d7n and D \u2208 R p\u00d7m are matrices that determine the elements of the output vector. It is assumed that the longitudinal motions and lateral/directional motions are uncoupled, which is validated by the small angle assumption. The dynamic models are expressed in the stability axes, which is shown in Figure 4. The flight condition used for identification was the power approach configuration with flaps down, gear down, and the parameters shown in Table 1.24 The states of the lateral/directional model are sideslip angle (\u03b2), body axis roll rate (p), body axis yaw rate (r), and body axis roll attitude angle (\u03c6). Aileron deflection (\u03b4a) and rudder deflection (\u03b4r) are the lateral/directional controls. A 3-2-1-1 aileron maneuver followed immediately by a 3-2-1-1 rudder maneuver was used to identify the lateral/directional model" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0000299_iros.1990.262416-Figure6-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0000299_iros.1990.262416-Figure6-1.png", + "caption": "Figure 6: The example of the manipulation by thervehi+ changing contacting points. The initial and final velocity of the target object is zero.", + "texts": [ + " Obtain the candidates of contacting points which give the nearly minimum value of lhl for q j , which is the average acceleration i n the period between t j N t j t l . Consider the sum of the fuel consumption AA4k in changing contacting points from the present points to the kth candidate cpr, =.(pkl, a . . , c p k ~ ) along the surface of a target object. Select the contacting points which give the minimum value of the following evaluation function: where - 396 - References The example of manipulation is shown in Fig.6. We can see that changes of contacting points occur in order to obtain small IkI and large DMM in Fig.& 5 CONCLUSION In this paper, we have described how to determine contacting points in manipulating a huge object by vehicles in space. We have proposed two indices for the determination. One is the norm of the contacting force which indicates how a c i e n t l y a target object is accelerated. Another is the DMM (direct manipulability margin) which indicates how widely a target object is accelerated around the desired acceleration" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003800_isda.2009.183-Figure2-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003800_isda.2009.183-Figure2-1.png", + "caption": "Figure 2. Stator and rotor flux vectors", + "texts": [ + "183 988 During one period of sampling Ts, the voltage vector applied to the machine remains constant, and thus we can write: sssssss TVTVkk ..)()1( \u2248\u0394\u21d2+\u2248+ \u03c8\u03c8\u03c8 (3) Because the rotor time constant is larger than the stator one, the rotor flux changes slowly compared to the stator flux. Thus torque can be controlled by quickly varying the stator flux position by means of the stator voltage applied to the motor. At any instant, the torque is proportional to the stator flux magnitude, the rotor flux magnitude, and the sinus of the angle\u03b4 (see Fig.2). It is expressed by: )sin()( \u03b4rsrse kkT \u03a8\u03a8=\u03a8\u00d7\u03a8= (4) Where: rs m LL LPk \u03c32 3= and rs m LL L 2 1\u2212=\u03c3 The estimated electromagnetic torque can be expressed as: )( 2 3 \u03b1\u03b2\u03b2\u03b1 \u03c8\u03c8 sssse iiPT \u2212= (5) III. NEURAL NETWORK STATOR RESISTANCE ESTIMATOR The stator resistance of an induction motor can be estimated with the adaptive estimator using neural networks as illustrated in Fig. 3. Two independent observers are used to estimate the rotor flux vectors of the induction motor. Equation (6) is referred to as \u201cvoltage model\u201d which is based on measured stator voltages and stator currents from the induction motor" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003010_20100906-5-jp-2022.00015-Figure4-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003010_20100906-5-jp-2022.00015-Figure4-1.png", + "caption": "Fig. 4. Horizontal guidance.", + "texts": [ + " Hence, the guidance is switched off while the path-planner is employed. Since the path-planner produces a path in two dimensions, we focus in this work on the guidance on the horizontal axis while the trim altitude should be hold. The altitude command for the autopilot is then simply the required trim altitude. Since the path-planner has to be switched off when approaching the waypoint, the last planned path is used to reach the waypoint. The following proportional controller for lateral guidance is used (see Fig. 4 for illustration): ( )ref 0 corrcommand \u03be \u03be \u03be \u03be= \u2212 \u2212 (5) where \u03beref is the pre-calculated angle at the reference point. The reference point is determined by finding the point closest to the current position on the last planned path and then taking the point on the path 50m ahead of this point. \u03be0 is the current angle, \u03becorr is the angle from the current position to the reference point. As we can see from (5) and Fig. 4, the guidance law consists of two terms which are explained as follows: The first term, \u03beref, describes the command/reference that the autopilot should follow if the aircraft was on the actual path with the correct heading; The second term, (\u03be0-\u03becorr), is a corrective term that guides the aircraft back on the path if it deviated from it. Remark 6: The maximum rate of change of flight path azimuth is limited to \u00b17\u00b0/s (see section 6), which results in a minimum turn radius of roughly 210m at a speed of 25 m/s", + " (6) is then solved at each sampling instant to obtain the sequence of optimal control actions and the first control of the sequence is applied on the control surfaces of the aircraft (see Fig. 2). To demonstrate the inter-effects of the layers we demonstrate the implementation of the three-layer control structure on the Aerosonde small fixed-wing aircraft by performing simulations of the implementation on the Aerosim block set in Matlab/Simulink. An example of a flight path with three waypoints is shown in Fig. 4. The aircraft starts at the trim conditions (altitude 1500m, wingtips level, 25m/s groundspeed, zero wind) at position (0m, 0m) facing straight north. The trim altitude and groundspeed shall be hold during the complete flight. Wind from the west with the von Karman turbulence is introduced as a ramp starting at t=0s and reaching 6m/s at t=6s which then persists for the whole of the simulation (Fig. 7). The path is re-planned at every time instance (every 0.4s) while far enough away from the waypoint" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0001293_rob.20099-Figure8-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0001293_rob.20099-Figure8-1.png", + "caption": "Figure 8. A spatial LOS sensing module.", + "texts": [ + " For example, a vehicle docking at a point along a straight line dock e.g., a train on tracks would require only one planar LOS, e.g., via a 1 DOF galvanometer mirror providing a rotation about a single axis, while a vehicle docking in 3D space e.g., a space satellite would require three spatial LOS, or more if there are potential obstructions or redundancy is required for increased certainty e.g., sensor fusion . In Table I, a spatial LOS refers to a laser beam aligned in 3D space by a 2 DOF galvanometer mirror Figure 8 . Correspondingly, a 2D array detector is necessary to receive this spatial LOS. One must note that optimal placement of the LOS laser sources and detectors on the vehicle would be necessary to facilitate effective guidance of the vehicle, namely, maximize visibility and, thus, reduce uncertainty in sensing data.16\u201318 As noted above, the acquired LOS-based external proximity-sensing data would not be in terms of the vehicle\u2019s relative pose with respect to the dock, but rather defined as offsets in the individual LOSdetectors\u2019 frames of reference", + " A computer simulation environment was developed in order to illustrate the working objective of the generic guidance algorithm in minimizing systematic errors and converging within the random noise of an autonomous system. Numerous simulations were run for a vehicle with 6 DOF mobility. The vehicle\u2019s shape is a cube with edges of 0.1 m. The three spatial LOS light sources are placed symmetrically around the perimeter of the docking workspace of the vehicle. Each source uses two 1 DOF galvanometer mirrors to provide a spatial LOS Figure 8 . The three array detectors, 50 mm by 50 mm, are placed centrally on the three faces of the vehicle, respectively. The inaccuracy of the motion of the vehicle is represented twofold: systematic errors and random errors. Instead of modeling each source of systematic error, their combined effect is represented herein as a function of the overall motion of the vehicle: Systematic error = inaccuracy/full range displacement of vehicle, 34 where inaccuracy/full range was chosen as 6 mm/m for translation and as 56 milli-deg/deg for rotation" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003210_j.mechmachtheory.2009.09.007-FigureC.1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003210_j.mechmachtheory.2009.09.007-FigureC.1-1.png", + "caption": "Fig. C.1. The three aligned points Ai;Aj and Ak .", + "texts": [ + " ; hj 1\u00de; \u00f0C:1a\u00de Aj \u00bc Uj;i\u00f0hi\u00fe2; . . . ; hj\u00de; \u00f0C:1b\u00de we consider the angle n defined by n \u00bc d Aj 1Aj ! ;AiAj ! ! : The value of n depends only on Ha \u00bc \u00f0hi\u00fe2; . . . ; hj\u00de. Setting in the reference frame \u00f0Aj; ~ij; ~jj\u00de: Ak \u00bc Uk;j\u00f0hj\u00fe2; . . . ; hk\u00de \u00f0C:2a\u00de we consider the angle w defined by w \u00bc d ~ij;AjAk ! ! : \u00f0C:2b\u00de The value of w depends only on Hb \u00bc \u00f0hj\u00fe2; . . . ; hk\u00de. Considering G\u00f0Ha; Hb\u00de \u00bc n w; \u00f0C:3\u00de the points Ai;Aj and Ak are aligned if and only if hj\u00fe1 \u00bc G\u00f0Ha; Hb\u00de \u00bdp : \u00f0C:4\u00de Proof. Referring to Fig. C.1, let hj\u00fe1 \u00bc d Aj 1Aj ! ;AjAj\u00fe1 ! ! \u00bc n\u00fe d AiAj ! ;AjAk ! ! w: ! The points Ai;Aj and Ak are aligned if and only if d AiAj ! ;AjAk ! \u00bc 0\u00bdp . h If hj\u00fe1 is nonsaturated, it belongs necessarily to the interval h j\u00fe1; h \u00fe j\u00fe1 . By using notations of Lemma C.1, the following definition can be given: Definition C.2. The multivalued operator H h j\u00fe1; h \u00fe j\u00fe1;Ha;Hb is defined as the set (with cardinality 0, 1 or 2) of the elements belonging to h j\u00fe1; h \u00fe j\u00fe1 which are congruent to G\u00f0Ha;Hb\u00de modulo p. Proposition C" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003903_(asce)as.1943-5525.0000057-Figure1-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003903_(asce)as.1943-5525.0000057-Figure1-1.png", + "caption": "Fig. 1. Interception geometry", + "texts": [ + " JOURNAL OF AEROSPACE ENGINEERING \u00a9 ASCE / JANUARY 2011 / 89 as a multiobjective optimization problem in Section 2 where the mathematical of the interception geometry and the architecture of the proposed guidance law are described. The different modules of the proposed multiobjective optimization algorithm are presented and discussed in Section 3 . The simulation results of the proposed approach are discussed in Section 4 . Finally, Section 5 concludes the paper. For the design purpose, only the two-dimensional engagement geometry as shown in Fig. 1 is considered where the missile and the target are assumed as constant-velocity point masses. The engagement model can be represented by the following equations Mishra et al. 1994 : \u0307 = VM sin \u2212 M \u2212 VT sin \u2212 T /r r\u0307 = \u2212 VM cos \u2212 M + VT cos \u2212 T \u0307M = aM VM x\u0307M = VM cos M y\u0307M = VM sin M \u0307T = aT VT x\u0307T = VT cos T y\u0307T = VT sin T 1 where aM and aT=missile and target accelerations, respectively. The proposed guidance law is shown in Fig. 2. It consists of three fuzzy-based guidance laws for launch, midcourse, and terminal phases" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0003424_s12206-009-1173-y-Figure6-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0003424_s12206-009-1173-y-Figure6-1.png", + "caption": "Fig. 6. Signal wires installed on the steel sheet spring.", + "texts": [ + " In order to prevent the strain gauge from damping, the strain gauge was covered with silicone sealing, as shown in Fig. 4. We ground and drilled the ring grooves and the piston crown, as shown in Fig. 5, then installed the rings and strain gauge, and finally drew signal wires from the strain gauge through the inside of the piston, and out of the piston crown. To prevent the signal wires from breaking under engine operation, we attached the signal wires to a steel sheet spring [1] with epoxy adhesive, as shown in Fig. 6. Signals from the strain gauge were sent via a bridge box to a dynamic strain amplifier, and finally displayed and recorded on an oscilloscope. In each experiment, we applied sufficient two-stroke oil to the cylinder wall at the engine bottom dead center (BDC), and then calibrated a ring strain of zero at the engine top dead center (TDC). We then ran the engine while measuring ring strain throughout the engine cycle. Experiments were carried out at room temperature with engine speeds of 100, 200, 300, and 400 rpm" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002839_6.2007-5094-Figure3-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002839_6.2007-5094-Figure3-1.png", + "caption": "Figure 3. Rotordynamic gas foil bearing test rig.", + "texts": [ + " One could hastily attribute the sub harmonic whirl motions to a typical rotordynamic instability induced by hydrodynamic effects of the gas film, i.e. generation of too large cross-coupled stiffness coefficients that destabilize the rotor-bearing system. However, as learned from the measurements in Ref. 24, rotor imbalance triggers and exacerbates the severity of subsynchronous motions. The subsynchronous behavior is a forced nonlinearity due to the foil bearing strong nonlinear (hardening) stiffness characteristics, as is demonstrated below. IV. Description of test rig and measurement of rotor-GFB dynamic responses Figure 3 shows the GFB test rig for rotordynamic measurements24,25,26. The hollow rotor, 209.5 mm in length, is of 38.10 mm in diameter at the foil bearing locations. A steel housing holds the GFBs and contains an internal duct to supply air pressure up to 7 bar (100 psig) for cooling the bearings, if needed. A 0.75 kW DC motor with maximum speed of 50 krpm drives the rotor through a flexible coupling and quill shaft. The motor torque is not large enough to overcome the bearings stall torque at rotor startup" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_69_0002403_1.3601696-Figure15-1.png", + "original_path": "designv11-69/openalex_figure/designv11_69_0002403_1.3601696-Figure15-1.png", + "caption": "Fig. 15 Clean-up compressor deve loped by M T I for (he AEC-LASL", + "texts": [], + "surrounding_texts": [ + "6 7 0 /' O C T O B E R 19 68 Transactions of the A S M E Downloaded From: http://tribology.asmedigitalcollection.asme.org/ on 01/27/2016 Terms of Use: http://www.asme.org/about-asme/terms-of-use Low Temperature. With the growing interest in superconductivity, lasers, masers, parametric amplifiers, and other low temperature electronic equipment, there have been requirements for efficient, miniature, cryogenic refrigeration systems. To date, many of the systems employ reciprocating components which, unfortunately, are not very reliable and contribute to contamination of the system. For turbomachinery components to be efficient in this application, they must operate at very high speeds (100,000-500,000 rpm) and the bearing and windage losses must be minimized. In fact, at liquid helium temperatures every watt of loss in the expander corresponds to approximately 2000 watts of power input to the compressor. Oil contamination in the system cannot be tolerated because it will adversely affect the performance of the heat-exchanger components. These reasons have stimulated several developments in high-speed gasbearing compressors and expanders. Fig. 8 presents a schematic of a closed cryogenic refrigeration cj-cle with a gas-bearing compressor and two expanders which utilize alternators for load absorption. The gas-bearings in the expanders are subjected to cryogenic temperatures. High Temperatures. Liquid-film (excluding liquid metals) and rolling-element bearings have been limited to relatively lowtemperature operation. This is primarily due to lubricant breakdown and decreased viscosity at elevated temperature. In the case of rolling elements, the bearing materials also limit the operating temperature. These barriers can be overcome with the use of gas-bearings. In addition, the design can be considerably simplified and the reliability improved when gas-bearings are used in high-temperature turbomachinery. Seals can be removed, thermal gradients can be minimized, machine length and weight can be reduced, and rotor dynamics can be improved, to cite just a few advantages. It is also well recognized that the efficiency of turbomachinery can be appreciably improved by going to higher turbine inlet temperatures (Fig. 9, [6]). Super alloys and other blade cooling techniques are being investigated which should enable turbines to operate for extended periods of time with gas temperatures in the 2000-;i000 deg F range. These trends, of course, should be paralleled by high-temperature bearing development. (One such study sponsored by the U. S. Army (AVLAB) is presently underway for the application of gas-bearings to a small two-spool aircraft gas turbine of very high speed and temperature.) In general, the high-temperature machines are also high-speed machines. This factor also gives gas-bearings an advantage over rolling-element and liquid-film bearings. The former loses lifecapacity with increases in speed while the latter consumes too much power. Several problems, however, must be solved in order to successfully apply high-temperature gas-bearings. These include: materials, coatings, methods of bearing support, manufacturing cost, etc. The three machines shown in Figs. 10-12 illustrate the applicability of gas-bearings to meet the high-temperature requirements. Fig. 10 shows a compressor test loop with a 1000 deg F, 18,000 rpm helium compressor which was developed for the United States Department of the Interior, Bureau of Mines [18). Here, the high-temperature requirements and \"zero\" contamination governed the choice of gas-bearings. This compressor utilizes self-acting gas-bearings. Fig. 11 shows parts of a 1900 deg F, 60,000 rpm gas turbine also developed for the United States Department of the Interior, Bureau of Mines. The bearings here are hybrid,3 using the same gas as is supplied to the gas turbine. This machine operated successfully for many hours at design conditions. Fig. 12 shows a gas turbine design with 2200 deg F turbine inlet temperature which utilizes either selfacting or hybrid bearings. The simplicity of this design should 3 F o r d e f i n i t i o n o f h y b r i d , re fer t o T a b l e 7 . Journal of Lubrication Technology O C T O B E R 1 9 6 8 / 6 7 1 Downloaded From: http://tribology.asmedigitalcollection.asme.org/ on 01/27/2016 Terms of Use: http://www.asme.org/about-asme/terms-of-use indicate clearly the advantages stemming from the use of gasbearings in high-temperature turbomachinery [19]. Radiat ion D a m a g e . For a nuclear heat source to be effective in power generation, the heat must be transferred from the reactor or isotope to the power conversion equipment. In the case of gas-cooled reactors this transfer is accomplished by circulators which must often operate in a high radiation field. Most conventional lubricants break down in high radiation fields; on the other hand, most gases are very stable. Space turbomachinery in the Van Allen belt must also be protected from radiation. Here, too, the use of gas-bearings eliminates the problem of lubricant breakdown. Contamination of the nuclear heat source or heat transfer equipment by a lubricant is also highfy detrimental. Very often, this machinery is \"hot\" and must be handled with automatic equipment which, in general, is veiy difficult. For these reasons the equipment must be highly reliable, have long operating life, and be safe to operate [20]. Figs. 13 and 14 show motor-driven circulators that have been developed for the Los Alamos Scientific Laboratory's project UHTREX [21]. Fig. 13 shows the first-generation machine which has now accumulated more than 6030 hrs of trouble-free operation under conditions more severe than those for which it was designed. Fig. 14 shows the second-generation machine which has incorporated in its design a number of improvements. Three such machines have been delivered to LASL and are part of the reactor system which went critical this year. These compressors have accumulated in excess of oOO hr of trouble-free operation (Table 1). Reliabi l i ty and Long Life. Without question, the user is always interested in equipment which is reliable and has long life. Inherently, turbomachinery is more reliable than reciprocating machinery. This is one of the reasons why gas turbines, rotary compressors, and expanders are replacing reciprocating machinery. In general, improvements in reliability result from evolutionary development. The jet engine is an excellent example of such evolutionary development. Within a span of 10 years, the jet engine has had a tenfold increase in operating hours between overhauls. Since a chain is as strong as its weakest link, the reliability of each individual component must be very much higher than the required system reliability: R,. = h'JUltR, PH Thus, from the standpoint of high reliability and low maintenance, it is good design practice to integrate and simplify the components or subsystems and to use as few parts as possible. As indicated earlier, the use of gas-bearings eliminates the use of several of the low reliability components (e.g., seals, pumps, controls, etc.) and greatly simplifies the mechanical design, thus inherently offering potential for higher reliability [22]. There are a number of examples where the reliability consideration is paramount; these cases are often associated with safety. Nuclear, space or deep-sea manned equipment are typical examples. Here, the user is willing to pay for the improvement in reliability. Figs. 15 and 16 illustrate two compressors which have been developed for nuclear applications where reliability and long-life were the deciding factors in the choice of gas-bearing turbomachinery. The first of these is a cleanup compressor for project 6 7 2 /' O C T O B E R 19 68 Transactions of the A S M E Downloaded From: http://tribology.asmedigitalcollection.asme.org/ on 01/27/2016 Terms of Use: http://www.asme.org/about-asme/terms-of-use LOW DP \"DETECTOR UUTIiEX [21]. This compressor uses self-acting bearings and it lias, thus far, accumulated more than 6459 hr of trouble-free operation (Table 1). The second figure shows several compressors which were designed for parallel operation in order to increase the systems reliability. These compressors are hybrid bearings and operate in helium or nitrogen. The compressor reliability has been so high that all of tlie standby compressors have been removed from the loop [23]. The compressors used have accumulated, to-date, in excess of 25,000 hr of trouble-free operation. High Speed a n d Long Life. Ill applications requiring high speed and long-life, major attention has to be paid to proper bearing choice and design. As pointed out in the foregoing, rolling-element and liquid-film bearings are not well-suited for high-speed, long-life applications. On the other hand, gas-bearings are; in fact, the presence of high-speed beneficial effects of increasing the load-carrying capacity. The power losses are two to three orders of magnitude smaller than with liquid-lubricated bearings. Low power level turbomachinery such as cited in the space applications (Figs. 6 and 7), cryogenic expanders (Fig. 8) or compressors, generally require both high speed (in order to achieve optimum efficiency) and long life. Also, in most of these applications, lubricant contamination must be avoided because it too will adversely affect the heat exchanger equipment and, in turn, system efficiency. Fig. 17 illustrates a compressor development which made the use of gas-bearings mandatory in order for it to meet the high-speed, long-life, high-efficiency, \"zero\" contamination specification requirements. Simplif ication. It was pointed out in the foregoing that the use of process fluid lubrication simplifies the design and increases the reliability of turbomachinery. Some applications benefit greatly Fig. 18(a) Fig. 17 1 5 0 , 0 0 0 rpm cyrogenic compressor developed by M T I EXTERNAL NITROGEN S U P P L Y Fig. 18(b) Fig. 18 Schematic d i a g r a m of: (a) CSN-2 lube-oi l a n d seal-gas system a n d (b) CSG-1 gas-bear ing turbocompressor lubrication system Journal of Lubrication Technology OCTOBER 1 9 6 8 / 673 Downloaded From: http://tribology.asmedigitalcollection.asme.org/ on 01/27/2016 Terms of Use: http://www.asme.org/about-asme/terms-of-use Fig. 20(c) Gas-bear ing system consisting of t w o - s h a f t turbocompressor a n d turboal temator Fig. 2 0 5 0 0 k w ( e ) turbomachinery configurations from the use of one or more bearing types and fluids. The following three examples will illustrate this point. The Advanced Power Conversion Branch of the U. S. Army Nuclear Power Field Office is conducting a technology program on compact Closed Brayton-Cycle electrical power equipment. Any oil getting into the Closed Brayton-Cycle Loop contaminates the heat exchanger surfaces and adversely affects the power generation. At the present time, an oil lubrication system and complex seals are used in the turbomachinery. The electric [lower output of the Closed Brayton-Cycle is controlled by the variation of the loop's gas pressure. As the loop's pressure is varied, the oil pressure to the journal and thrust bearings must also be varied or the oil will leak into the loop. This, in fact, has occurred inadvertently on several occasions. The application of gas-bearings, of course, eliminates this problem (Fig. 18, [24]). In addition, it simplifies the system making it more compact, more efficient, and of higher reliability [24]. The bearings in such a system can be hybrid and can use the compressor discharge pressure. A dynamic simulator was designed and tested to verify the applicability of gas-bearings to a 500 kw(e) Closed BraytonCycle power system. Fig. 19 illustrates the simulator which consists of a two-stage axial turbine and a single-stage radial compressor. The rotor weight was ninety (90) lb and the speed was 28,000 rpm with an over-speed of 35,000 rpm. The thrust bearing was designed for an estimated aerodynamic load of five hundred (500) lb. Tests were extended up to seven hundredfifty (750) lb thrust load. The program was successfully completed and the results clearly indicated the applicability of gas- ^ Discharge Turbine-Driven Pump supported on one 6team bearing and one water bearing Fig. 2 2 Turbine-driven feedwater pump designed by M T I bearings to 500 kw(e) Closed Brayton-Cycle turbomachinery [25]. Further simplifications of the system are possible by elimination of the reduction gear (Fig. 20(a)) and integrating the turbocompressor with the alternator (Figs. 20(6,c), [20]). At one time the power supply in the B-52 used bleed-off air from the main jet engines to power turbine-driven 60 K V A alternators. These APU's employed turbine-driven gear boxes and low-speed 6000 rpm alternators. In new applications it was desirable to reduce appreciably the weight of the system without sacrificing reliability. This was accomplished by the \"Turbonator\" design illustrated in Fig. 21. This design employs a directly driven 24,000 rpm, 60 K V A alternator. However, from rotor dynamics and temperature standpoint, a rolling-element bearing could not be used next to the turbine; thus a hybrid bear- 674 /' O C T O B E R 19 68 Transactions of the A S M E Downloaded From: http://tribology.asmedigitalcollection.asme.org/ on 01/27/2016 Terms of Use: http://www.asme.org/about-asme/terms-of-use ing was employed. This journal bearing used the same gas supply as the turbine. The bearing on the cold end of the machine was a rolling-element, bearing that took the radial and thrust loads. This design satisfied all the objectives of the program and made the machine very simple and highly reliable. Conventional turbine-driven hydraulic pumps use three fluids, the turbine fluid (generally oil), and the hydraulic fluid. It should be recognized that, in most cases, the turbine fluid (steam) and the hydraulic fluid can be effectively used in bearings. In fact, the steam supply and the pump discharge pressure will generally be sufficiently high so that hybrid bearings can be designed into the equipment. Fig. 22 illustrates a turbine-driven, feed-water pump which employs a hybrid water bearing on the pump end and a steam bearing on the turbine end [22]. This design eliminates seals, fire hazards, and accessor}- lubrication equipment, and reduces the overall size and weight of the machine. Recent studies of naval machinery [22, 27, 28] have focus ed attention on the use of process fluid as a lubricant. The advantages gained from the application of process fluid lubrication are listed in the discussion section of this paper and in the foregoing references. Some of the process fluids available in the engine room are air, steam (at or near saturation), and water. In the development of new or improvement of existing machinery lubrication with any of these fluids should be considered. Ideally, in order to minimize external piping and the problem of mixing of the fluids, steam lubrication should be used to lubricate machinery where steam is the process fluid, while water lubrication should be used where water is the process fluid. In regard to steam lubrication, two choices are available. The first of these is to superheat, the steam, ahead of bearing inlet, sufficiently to insure that there will be no consideration anywhere in the bearing system. Alternately, the steam can be furnished to the bearings in its saturated condition. The second alternative avoids the use of auxiliary heating of high-pressure steam. However, it introduces the additional problems of two-phase flow, condensation, and evaporation in the bearing and restrictor passages [29], Recently a design study for a U. S. Navy forced draft blower utilizing steam lubrication was completed. The machine is now under construction. To the best of the author's knowledge, this is the first machine to utilize steam lubrication (Fig. 23). General Remarks The preceding examples have shown several cases where gas- bearings were mandatory and other cases where the advantages warranted the development effort. As more experience with gas-bearing turbomachinery is gained, the development effort will decrease and it is expected that gas-bearing turbomachinery will receive very wide acceptance. There is ever}' reason to believe that in quantity production gas-bearing machinery should be lower in cost to other comparable turbomachinery and thus offer the additional economic advantage. The development effort in gas-bearing turbomachinery centers around compressible analysis of hydrodynamic and hybrid fluid films with special emphasis on spring and damping coefficients for rotor response and instability analysis [30]. The various pa- M T I for U. S. N a v y Fig. 2 4 Several molor-dr iven rotors a n d compressors Journal of Lubrication Technology O C T O B E R 1 9 6 8 / 6 7 5 Downloaded From: http://tribology.asmedigitalcollection.asme.org/ on 01/27/2016 Terms of Use: http://www.asme.org/about-asme/terms-of-use T a b l e 5 G o v e r n i n g , d i m e n s i o n l e s s p a r a m e t e r s Gas Lubricant (Compressible) ~ Length-To-Diameter Ratio 12sli: Bearing Number: A = \u2014 - \u2014 _ , W Load: ^ Feeding Parameter: Ac Supply Pressure Ratio : 6uNa